REFRIGERATION SCIENCE AND TECHNOLOGY SCIENCE ET TECHNIQUE DU FROID

Low temperatures and electric power

Transmission Motors, transformers and other equipment

Cryogenics and properties of materials

Cryoelectrotechnique Transport d'electricite

Moteurs, transformateurs et autres equipements Cryogenie et proprietes des materiaux

PERGAMON PRESS Oxford • New York • Toronto

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Copyright © 1970 Pergamon Press Ltd., and The International Institute of Refrigeration

All Rights Reserved. No part of this publication may be reproduced, stored in a retrieval system, or transmitted, in any form or by any means, electronic, mechanical, photocopying, recording or otherwise,

without the prior permission of Pergamon Press Ltd. First edition 1970

Library of Congress Catalog Card No. 74-143142

Printed in Belgium by Ceuterick 08 016370 X

FOREWORD

MEETING OF COMMISSION I OF THE I.I.R. LONDON (U.K.)

MARCH 24-28, 1969

Commission I of the 1.1. R. held two symposia from 24-28 March, 1969, in London, under the general title of "Cryogenics in Fuel and Power Technology". These meetings reflect the growing importance of cryogenics and refrigeration in the distribution of natural gas for power supplies, electrical engineering, and future methods for the distribution of electricity.

One symposium was devoted to "Liquefied Natural Gas" and was attended by over 400 participants, including 150 visitors to Britain. This conference was organised by Commission I in conjunction with the British Cryogenics Council, with the assist­ance of the Institution of Mechanical Engineers. The conference on electricity, "Low Temperatures and Electric Power", was attended by 300 participants, of whom 145 came from 14 different countries outside Britain. This meeting was organised by Commission I in conjunction with the British Cryogenics Council, with the assistance of the Institution of Electrical Engineers.

The present volume is an account of the meeting on Low Temperatures and Electric Power, the proceedings of the other symposium are being published separately. The Commission is much indebted to the organising committee of the meeting*, and particularly to its Chairman, Professor N. Kurti, F.R.S., for their care and attention in the planning of the programme.

ORGANIZING COMMITTEE*

Mr. A.D. Appleton (International Research & Development) Mr. D.R. Edwards (British Insulated Calender's Cables) Dr. B.B. Goodman (British Oxygen Company) Dr. J.W.L. Kohler (Philips N.V.) Prof. N. Kurti, Chairman (University of Oxford) Mrs. M.K. McQuillan (Imperial Metals Industries) Dr. W.T. Norris (Central Electricity Research Laboratories) Dr. T. Raine (Associated Electrical Industries-General Electric Company) Mr. R.G. Cox, Secretary (Institution of Electrical Engineers)

8

AVANT-PROPOS

REUNION DE LA COMMISSION I DE L I . I. F. LONDRES (R.U.)

24-28 MARS, 1969

La Commission I de l'l. I. F. a tenu deux reunions du 24 au 28 mars 1969 a Londres, sous le titre general de « La cryogenie dans la technique des combustibleset de 1 'energie». Ces reunions refletent l'importance croissante de la cryogenie et du froid dans la distribution du gaz naturel pour la fourniture d'energie, pour l'electrotechnique et pour les methodes futures de distribution d'electricite.

Une reunion a ete consacree au gaz naturel liquefie, a laquelle assistaient plus de 400 participants, dbnt 150 non Britanniques. Cette conference a ete organisee par la Commission I en liaison avec le «British Cryogenics Council)), avec l'appui de F «Institution of Mechanical Engineers)).

A la reunion sur la «Cryoelectrotechnique» assistaient 300 participants, dont 145 non Britanniques venaient de 14 pays differents. Cette reunion a ete organisee par la Commission I en liaison avec le «British Cryogenics Council», avec l'appui de F«Institution of Electrical Engineers)).

Ce volume est le compte rendu de la reunion sur la Cryoelectrotechnique; les comptes rendus de l'autre reunion seront publies separement. La Commission exprime ses remerciements au comite d'organisation*, et en particulier a son President, le Professeur N. Kurti, F.R.S., pour les soins qu'ils ont apportes a l'etablissement du programme.

COMITE D ' O R G A N I S A T I O N *

Mr. A.D. Appleton (International Research & Development) Mr. D.R. Edwards (British Insulated Calender's Cables) Dr. B.B. Goodman (British Oxygen Company) Dr. J.W.L. Kohler (Philips N.V.) Prof. N. Kurti, President (University of Oxford) Mrs. M.K. McQuillan (Imperial Metals Industries) Dr. W.T. Norris (Central Electricity Research Laboratories) Dr. T. Raine (Associated Electrical Industries-General Electric Company) Mr. R.G. Cox, Secretaire (Institution of Electrical Engineers)

9

LISTE DES PARTICIPANTS LIST OF PARTICIPANTS

ALLEMAGNE (Ouest-Rep. Fed.) — GERMANY (West-Fed. Rep.)

BLECHSCHMIDT, H., Diisseldorf BOGNER, G., Erlangen BUCHHOLD, T., Wiesbaden DOOSE, C , Fuelich GANN, A., Stuttgart-Vaihingen GANSKE, R., Hamburg HARTWIG, G., Karlsruhe HELLER, L, Aachen HILDEBRANDT, U., Hollriegelskreuth HILLMANN, H., Hanau KUHLMANN-SCHAEFER

LEMMERICH, J., Berlin LUCKING, H.W., Koln-Mulheim SASSIN, W., Jiilich SAUR, E.J., Giessen SCHEFFLER, E., Hannover SCHMIDT, F., Erlangen SELLMAIER, A., Hollriegelskreuth VOIGT, H., Frankfurt-Niederrad WANSER, G., Hannover WEINHOLD, J., Hamburg

ALLEMAGNE (Est) — GERMANY (East)

BEWILOGUA, L., Dresden HAUNSTEIN, W., Jena

MULLER, G., Dresden NEUBERT, J., Dresden

BELGIQUE — BELGIUM

CODLING, N., Bruxelles FENEAU, C , Hoboken-Antwerpen

LECOINTE, G., Hoboken-Antwerpen

CANADA

CASS-BEGGS, D., Ottawa RAMSHAW, R.S., Waterloo

RINGER, T.R., Ottawa

BAK, C , Lyngby BALSLEV, N., Lyngby

DANEMARK — DENMARK

KOFOED, B., Lyngby

ETATS-UNIS —U.S.A.

ARP, V.D., Boulder BOOM, R., Madison CHELTON, D.B., Boulder COFFEY, H.T., Menlo Park

DAUNT, J.G., Hoboken JEANMONOD, J., Hyde Park LAVA, V.S., Portland MATISOO, J., York town Heights

10

MINNICH, S.H., Schenectady NICOL, J., Cambridge PASTUHOV, A., Cambridge SNOWDEN, D.P., San Diego

STEETER, M.H., Cambridge WALKER, C , Murray Hill WATSON, J.H.P., Corning WHITMAN, C.H., New York

FRANCE

ADAM, E., Chatou AUPOIX, M., Marcoussis BARRETT, P., Clamart BERTHET, M., Marcoussis BLANC, J.M., Chatou BORDENAVE, J.P., Le Bourget BRUNEAU, P., Paris BURNIER, P.H., Massy CARBONELL, E., Sassenage CARVOUNAS, E.G., Jeumont CHABRERIE, J.P., Fontenay-aux-Roses CLAUDET, G., Grenoble CROITORU, Z., Clamart DAMMANN, C , Marcoussis DELILE, G., Clamart DUBOIS, P., Marcoussis FABRE, J., Paris FALLOU, (Mme) B., Fontenay-aux-Roses FERRIER, M., Clamart GALAND, J., Fontenay-aux-Roses GILCHRIST, J., Grenoble

GIRARD, B., Or say HARPE, A.P. de la, Sassenage HELLEGOUARC'H, J., Le Bourget LACAZE, A., Grenoble LAIR, P., Saint-Ouen LECOMPTE, J., Grenoble LEFEVRE, F., Sassenage LEHONGRE, S., Sassenage LESAS, P., Le Bourget MAILFERT, A., Fontenay-aux-Roses MARQUET, A., Clamart MOISSON, F., Marcoussis NEVEU LEMAIRE, D., Marcoussis PECH, T., Fontenay-aux-Roses SAMMAN, J., Grenoble SCHWAB, (Melle) A.M., Clamart SHIMAMOTO, S., Gif-sur-Yvette STAHL, Bruyeres le Chatel SYRE, R., Argenteuil THOMAS, P., Bruyeres le Chatel

GADDA, E., Milano OCCHINI, E., Milano

ITALIE — ITALY

SACERDOTI, G., Frascati

JAPON — JAPAN

YAMAMOTO, M., Kawasaki

PAYS-BAS — THE NETHERLANDS

CLAASSENS, A.M.J.M., Delft CLASON, R.J., Delft DELWEL, J.W., Eindhoven GOEMANS, P.A.F.M., Eindhoven GOLDSCHVARTZ, J.M., Delft KOHLER, J.W.L., Eindhoven LELIE, M.C., Delft MIJNHEER, A., Eindhoven

PRAST, G., Eindhoven RIEDIJK, W., Eindhoven ROMYN, J.G., \s Hertogenbosch STER, J. van der, Eindhoven VERBEEK, H.J., Eindhoven VOLGER, J., Eindhoven WIJKER, W.J., Arnhem

11

POLOGNE — POLAND

MAZUR, J., Wroclaw SKOWRONSKI, J.I., Wroclaw

ROYAUME-UNI — UNITED KINGDOM

ALEXANDER, V.J., Gt. Malvern ALHOUSENI, F., St. Andrews ALLAN, R., Manchester APPLETON, A.D., Newcastle-upon-Tyne ARTON, K.A.M., Glasgow ASHMOLE, P.H., London BARBER, A., Birmingham BARKER, B., London BATES, J.J., Swindon BATTAMS, (Miss) P.A., Guildford BECKLEY, P., Newport BURTON, R.A., London CAIRNS, D.N.H., Leatherhead CATTERALL, J.A., Teddington CHESTER, P.F., Capenhurst CHORLTON, A., London CHU, S.C., Brims town CLARKE, M.E., London COLYER, B., Didcot CORNISH, D.N., Abingdon COUPLAND, J.H., Didcot Cox, J.R., London CROSSLEY, I., London DAVIDSON, D.F., Manchester DAWSON, C.N., Brighton DUNN, W.I., Glasgow EADIE, G.C., Newport EATWELL, A.J., London EDEN, P.M., London EDWARDS, D.R., London ELLIOTT, J.M., Newcastle-upon-Tyne EVANS, R.M., London FEDORKO, G., New Maiden FIRTH, I., St. Andrews FLEWITT, P.E.J., Cockfosters FORREST, A., Paisley FRENCH, R.A., Cuffley FURTADO, C , Oxford GARDNER, J.B., London GARNELL, A., Walsall GAYDON, B.G., Leatherhead GOLUB, R., Brighton GOODMAN, B.B., London GORE, D.C., London GRAFF, C , Southampton

GREEN, I.M., Glasgow GREIG, D., Leeds GRIGSBY, R., London HANCOX, R., Abingdon HARRIS, A.G., Malvern HASELDEN, G.G., Leeds HASELFOOT, A.J., London HAYDEN, J., Cranfield HENSEL, P., Oxford HLAWICZKA, P., Glasgow HOULIHAN, J., London HUSSEY, R.T., Bath INSTONE, C.S., London JAMES, T., Abingdon JEWELL, P., London JOHNSON, R.H., London JONES, J.E., London KENDALL, P.G., Leatherhead KNAPTON, A.G., Manchester KUGLER, S., London KURTI, N., Oxford LANE, F.J., Brighton LAWRENCE, G.S., Southampton LAWRENSON, P.J., Leeds LEWIS, D.S.G., London LINDLEY, B.C., Leatherhead LONDON, H., Harwell LORCH, H.O., Stafford MACNAB, R.B., Newcastle-Upon-Tyne MADDOCK, B.J., Leatherhead MALE, J.C., Leatherhead MARRIOTT, A.P., Southampton MARTIN, F.P., London MASCHIO, G., Eashleigh MCQUILLAN, (Mrs) M.K., Birmingham MEATS, R.J., Leatherhead MELVILLE, P.H., Leatherhead MIDDLETON, A.J., Didcot MIDGLEY, F.B., London MONROE, A.G., London MORTON, I.P., Southampton MORTON, N., Salford MORTON, P.H., Birmingham MULHALL, B.E., Newcastle-upon-Tyne NORRIS, R., Manchester

12

NORRIS, W.T., Leatherhead OFTEDAL, E., Sittingbourne PARK, J.G., London PEARCE, D.G., Didcot PLESSNER, K.W., London PREECE, C , Durham QUAYLE, J.P., London RAINE, T., Manchester RAY, J.P., Guildford REINECK, K.M., Durham RHODES, R.G., Coventry RICHARDS, T.L., London RICKETSON, B.W.A., Pitchcott ROBERTS, D.C., London ROGERS, E.C., London ROSE, H., Newcastle-upon-Tyne Ross, J.S.H., Newcastle-upon-Tyne SALDANHA, S.A., Southampton SALIM, A.J., Manchester SALMON, D.R., Teddington SCURLOCK, R.G., Southampton SHEPPARD, H.J., London SMITH, G.A., Leicester

SMITH, P.F., Chilton SOTT, M., Oxford SPEAR, C.H., Barnet SPENCE, S.T., Oxford STEARN, J.W., Malvern STEEL, A.J., London STOVOLD, R.V., Didcot STRANGEWAY, P.J., London SUTTON, J., Leatherhead SWIFT, D.A., Leatherhead TAYLOR, M.T., Leatherhead THOMPSON, D.S., Dundee TINLIN, F., Newcastle-upon-Tyne WALKER, A.J., Sittingbourne WALTERS, C.R., Didcot WEEDY, B.M., Southampton WIGLEY, D.A., Southampton WILKINSON, K.J.R., Rugby WILKS, J., Oxford WILLIAMS, J.E.C., Oxford WILLIAMS, M., Wembley WILSON, M.N., Didcot ZVEGINTZOV, M., London

SUfiDE — SWEDEN

CARLSSON, R.G.I., Gothenburg FORSBERG, A.A.E., Malmo

HALLENIUS, K.E., Gothenburg MADSEN, K.D., Vasteras

SUISSE — SWITZERLAND

BEBI, H.H., Zurich BENZ, H., Zurich BRATOLJIC, T., Birr FASEL, R., Zurich GUREWITSCH, A.M., Zurich KULL, U.N., Baden MARTIN, G., Cossonay

NEIDHOEFER, G.J., Baden OLSEN, J.L., Zurich SICKENBERG, H., Geneve SMITH, J., Geneve SZASZ, G., Zurich TREPP, C , Winterthur VECSEY, G., Zurich

TCHECOSLOVAQUIE — CZECHOSLOVAKIA

CESNAK, L., Bratislava CHYTRACEK, V., Prague KAISER, Z., Prague

KURKA, J., Pilsen TROCHTA, Z., Bratislava

13

U.R.S.S. —U.S.S.R.

BELIAKOV, Y.P., Moscou EPIFANOVA, (Mrs) V., Moscou GRIGORIEV, V., MOSCOU IVANZOV, O.I., Moscou KLIMENKO, E., MOSCOU POZVONKOV, F.M., Moscou

PRONKO, V.G., Moscou SAMOILOV, B. SHURGALSKY, E.F., Moscou SICHOV, V.V., Moscou SYTCHEV, V.

I.I.F. — I.I.R.

THEVENOT, R., Directeur

14

SURVEY PAPER

PROBLEMS OF THE ELECTRICAL POWER INDUSTRY

A. CHORLTON

Central Electricity Generating Board, London (United Kingdom)

Problemes de Pindustrie de la production d'electricite

RESUME : Le rapport se limite aux problemes des systemes electriques. II passe d'abord en revue les facteurs economiques qui ont motive la mise au point du systeme jusqu a son etat actuel a"integration complete, expliquant la politique de conception du reseau adoptee pour faire face aux problemes techniques impliques dans Vintegration. On illustre la disposition fondamentale des reseaux a" integration pour montrer les fonctions impliquees dans la transmission principale, la sous-transmission et la distribution et pour expliquer le concept du classement de la tension pour Vexploitation technique et economique des niveaux de courant. On donne les principales statistiques portant sur Vefftcacite de la transmission et les tendances des inves-tissements en transmission.

Le rapport considere ensuite en detail les concepts techniques expliquant Vinterdependance des caracteristiques du materiel et la coordination necessaire pour concilier les besoins contra-dictoires de stabilite du systeme, de qualite de la tension et des questions de court-circuit et pour repondre aux besoins de reglage de la frequence et de la tension. On indique les tendances des caracteristiques du materiel et du systeme avec iaugmentation des dimensions et Vevaluation et le role joue par la resistance et la reactance dans le comportement dynamique du systeme.

Le rapport etudie finalement les besoins et les problemes futurs. II examine la possibilite de developpement permise par la tension actuelle de transmission de 400 kV et combien les pra­tiques de conception classiques repondraient aux besoins du reseau de distribution, les problemes particuliers poses par V utilisation croissante de cables sou terrains pour la transmission princi-pale et secondaire et les perspectives de la transmission de courant continu dans ces conditions. On examine Vintegration des chaines de transmission de courant continu a usage special a la lumiere de leurs caracteristiques de fonctionnement particulieres et Vinfluence qui en resulte sur le comportement dynamique du systeme hy bride. On fait ressortir le probleme fondamental d'harmoniser des appareils de caracteristiques radicalement dijferentes dans le systeme de courant alternatif integre classique comme base de Vexamen des perspectives d'application des techniques supraconductrices.

Le rapport conclut par des remarques sur les problemes de mise au point des systemes et sur les pratiques en cours au Royaume- Uni en comparaison des autres pays.

INTRODUCTION

When I was invited to present this survey I took comfort in the thought that the broadness of the title gave me licence to confine my examination to some problems of the electrical power industry—I certainly could not deal with all of them.

I shall deal with engineering problems in the planning and design of the integrated grid and supergrid systems for England and Wales, recognising that many of the technical constraints imposed apply also to other systems, although with varying degrees of influence.

I shall first briefly review the background to the emerging system, outlining the planning and design philosophy, and then dwell in some detail on the technical constraints—because these constraints condition development of systems and proper understanding of them is essential to any consideration of system economics.

I shall conclude with a look ahead at the more distant future, which is always risky, examining system development potential and problems, based on current technology.

BACKGROUND TO THE EMERGING SYSTEM

The high-voltage 'transmission grid' is a feature common to public electric supply systems of most industrialised countries. In many of these countries the

17

prime purpose has been to exploit natural energy resources geographically remote from consumption centres, such as hydro power. In this country, the prime motivation has been the substantial economies to be gained from pooling of generation and consumer loads to minimise spare generating capacity and to maximise usage of the best plants.

The original 132 kV transmission grids, conceived in the 1930's, were built expressly for this purpose. They yielded seven regional power pools and enabled the spare plant over the country as a whole to be reduced from 70 % to 26% and unit production cost to be cut by over half. There were fringe interconnectors between these regional networks intended solely for emergency use. Although of limited capability, their enforced regular usage just before and during the war, demonstrated the further economic advantages of national pooling. This experience led in the 1950's to the building of the 275 kV superimposed grid, expressly to achieve this national pooling on an adequate and continuing basis.

The economies from integration stem in the first place from generating savings outweighing consequential transmission costs. These are made possible in most countries by the comparative cheapness of the overhead transmission line and are rendered substantial in this country by the relatively short distances involved. However, notwithstanding the short distances, pooling would not have been economically feasible on anything like its present scale using the costly underground cable trans­mission.

There are other important benefits from integration, the most oustanding being the economies of scale. The greater stored energy of the power pool makes it technically feasible to use larger generating units and the grid network provides the basic frame­work on which to build the extra transmission facilities to deal with them. The 132 kV transmission grids developed this way; they permitted larger generating stations to be accommodated and, with appropriate strengthening for the extra function, were used to distribute the output from these new stations to both existing and new con­sumption centres within the individual regions. Thus, they emerged into dual purpose grids fulfilling the functions of both interconnection and local high-power distribution.

The history of the supergrid has been basically similar. It was conceived at the outset as a dual purpose system, but the second function was that of longer distance bulk transmission, to save fuel transport. A third function has now been added—that of shorter distance high-power distribution, to accommodate the output from still larger sets and stations that have become economically feasible in an expanded system. This has involved using built-in potential of the 275 kV lines, by converting them to 400 kV, and physically extending the network to reachTmore consumption centres and link with the new central power stations.

The main framework of transmission lines which will provide this expansion is now approaching completion, as is evident from the map of the 400 kV main network in figure 1. The network will cover the whole country, providing capacity for its three functions. It will be noticed that the principal conurbations are ringed rather than traversed by the 400 kV network. The planning problem with these highly populated areas is to transmit thousands of MW's over short distances from peripheral sources to the very local 132 kV or 66 kV sub-networks, with an absolute minimum call on space. This is best done by continued development at 275 kV, often by cable, within the conurbation, supplying these 275 kV networks by large capacity transformer feeders from a 400 kV power source ring around the conurbation. By this means are achieved the benefits of a very high-power infeed circuit (at 400 kV) and, at the same time, smaller site and cable requirements within the conurbation.

The following are salient statistics of the planned 400/275 kV supergrid network and the parts so far in service.

18

Table 1

Equipment Existing 1969 Planned for 1975

Overhead Line-Route Mileages Working at 275 kV 1,650 1,050 Working at 400 kV 1,580 3,300

Total 3,230 4,350

Underground Cable-Circuit Mileages Working at 275 kV Working at 400 kV

Total

Switching Station-Number 275 kV 400 kV

Transformer Capacity-MVA 400/275 kV 400/132 kV 275/132 kV or

Site Total

lower voltage

230 10

240

97 33

122

28,000 15,600 53,900

294 29

323

101 71

158

60,000 39,000 66,000

The creation and strengthening of the 400/275 kV supergrid network has accounted for the bulk of the physical expansion of main transmission facilities in England and Wales over the past 15 years. Concurrent with this development, the original 132 kV grid networks have been progressively sub-divided and contracted, and re-organised into discrete reticular sub-networks supplied from the supergrid through transforming stations. Their primary function now is medium power distribution. While some new sections of 132 kV overhead line have been necessary, others have been dispensed with, and most of the extra 132 kV circuit mileage has been in the form of underground cabling.

P L A N N I N G AND DESIGN PHILOSOPHY

The way in which the forementioned developments will change the usage of the supergrid is shown by figure 2 which traces the growth in system power movement. By 1975 rather more than half the power production at time of peak demand will be flowing through the supergrid network, compared with a little over a quarter at the present time. Most of the increase arises from the high-power distribution function. Since it is the newer base load stations that will be feeding direct into the Supergrid, the proportion flowing through that network will be greater at off-peak times, reaching close on 100% at the lowest system loads. This trend is expected to continue—thus, when the system has doubled in size, more than three-quarters of the total peak production will require to be conveyed over the supergrid.

This almost completes reliance on the supergrid, making it the main artery for conveyance of electricity supplies, puts great emphasis on reliability and on the provision of adequate standby capacity for security. The reconciliation of these require­ments with the conflicting and pressing amenity problem has greatly influenced system design concepts and, in particular, the structural arrangement of the supergrid network. The schematic diagram in figure 3 demonstrates this operative structure. It has the familiar grid-iron pattern. Power will be fed into the network from the

19

control station (shown as open rectangles) in blocks between 1000 and 2000 MW and taken off in smaller quantities from the far more numerous transforming stations (shown by the blacked in rectangles). The 400/275 kV transformer feeders into the various conurbations areas are clearly shown.

Fig. 1 — Main 400 kV network.

The diagram illustrates the marshalling of circuits in substations at various line route intersections. This marshalling is necessary to achieve sufficient electrical shortening of circuits, and a short circuit level commensurate with the power carrying capability required of the circuits. The short circuit level reflects the reactive power available to sustain the voltage and hence the transmitted power. The 275 kV lines converted to 400 kV have two conductors per phase and a thermal rating of 1100 MVA per circuit; this is about twice the surge impedance loading and requires roughly 2.5 MVARs per mile to sustain it. The newer 400 kV lines have twice as many conduc­tors and twice the thermal rating, requiring roughly four times the reactive power. These reactive power needs call for the most intensive marshalling where the heaviest loads are, notably the feeding points to the conurbations.

Security philosophy recognises that reliance on the supergrid will be continuous, extending over periods when circuits must be taken out of service for maintenance or repair. It also recognises that use of double circuit construction for the overhead lines exposes them to the risk, albeit small, of both circuits being faulted simultaneously. The general policy is to provide cover for a double contingency, the system being designed to withstand the consequences of, simultaneous disconnection on fault of both circuits of a double circuit line, or simultaneous outage of any two circuits.

20

The same criteria is applied to the transformer connections to the subsidiary networks, to the extent that any demands in excess of 300 MW are not put totally at risk in the event of two coincident circuit outages. The practical effect of these security standards

Fig. 2 — Trend in power transfers.

is that all demands in excess of 300 MW have at least three supply circuits with ratings further dependent on the demand level.

TECHNICAL CONSTRAINTS

Performance of power systems, i.e., their supply capabilities are conditioned by the stability of individual generators and of interconnected groups of generators. As systems grow in power density, the consequences of instability become more unpredictable and potentially more severe, hence design criteria require to be more stringent. Performance of consumer apparatus both influences and is influenced by the quality of voltage, and conditions the design of supply connections. In the British grid, these technical constraints have a differing influence at the approach to the extremes of system loading.

When consumer load is high and most plants are generating, the transmission network provides a stiff coupling between plants, and they react closely together;

21

-CM DSIZEWELL

Fig. 3 — Schematic outline of main 400 kV network.

Fig. 4 — Inter-network transformation arrangements.

22

•♦OO KV OR 275 KV

2 i

DISTRIBUTION NETWORKS

Fig. 5 — Network subdivisions.

RELATIVE U N IT S I ZE

Fig. 6 — Transmission plant specific costs relative to unit size.

23

also the generators generally must have strong field excitation to furnish reactive power for the high consumer load and the heavily loaded network; this gives them greater stability. Generally, the problem at this time is not stability of individual generators or stations but stability between large groups of stations with limited interconnection. The requirements for this system stability are, network configura­tions which do not yield large changes in transfer impedance following circuit outages, and quick clearance of faults through fast protective relaying. The latter is now becoming the most influential parameter. With system growth and stronger inter­connections, the voltage depression from faults can be widely reflected, resulting in transitory loss of load on many plants. By keeping the depression time as short as possible, stability is enhanced.

When consumer load is low, and consequently only a few plants are generating in widely separated parts of the system, the interconnection capacity is lessened by distance; also the consumer needs little reactive power and the network none at all—in fact it has a surplus. Consequently the field excitation of the generators must be reduced to allow them to absorb reactive power, and they are then less stable. Generally, the problem at this time is stability of individual generators and stations. The network reactances and design parameters of the generators influence this behaviour, but with increasing size of unit, there are economic pressures to accept natural values yielding less inherent stability, and increasingly now the main reliance is on fast automatic generator excitation controls, and on fast protective relaying.

Another facet to stability imposing different constraints are the frequency excur­sions that immediately follow the sudden disconnections of plant or load. The excursions are conditioned by the kinetic energy of the system and tend to assume most significance at low consumer loads, when the running generation and hence the system kinetic energy is also low. System performances in these situations condition the size of individual generating unit in a given system, also transmission arrangements where these put generation or load at risk to a single fault incident. They can also condition the employment of asynchronous interconnecting links in systems and their control arrangements for contribution to system frequency regulation. The direct current cross-channel link which interconnects the British and French systems is one such as asynchronous link. However, this particular interconnection was conceived as a pilot scheme and it is too small to make a worthwhile contribution to the frequency regulation needs of either system. Hence it has no form of frequency bias.

Complementary to the control of system frequency is, of course, the control of system voltages. The basic mode of control is the same for both, that is, the balan­cing of consumption with production—in the one case the 'active power' and in the other, the 'reactive power*. The requirements of the transmission and distribution networks enter into this balancing. While these requirements are very small in the case of active power—for example the losses over the supergrid and grid networks vary only from 1.0 to 3.0% over the full range of system loading—they are widely variable and at times very large in the case of reactive power. At high system loads there is substantial consumption of reactive power by the networks as well as the consumer load calling for extra production from the generators. At low system loads the consumer wants very little and there is substantial surplus production from the lightly loaded network, calling for some absorption by the generators. The problem shows up most during the rapid fall-off in consumer load in the late evening. The reactive burden on the generator then changes rapidly frcm production to absorption, this absorption duty being imposed on fewer generators.

The surplus reactive production arises from the capacitative reactances in the networks when energised, growing with expansion of networks, with the use of higher voltages, and particularly with use of underground cable. In the CEGB system.

24

it is now much more than can be absorbed with stability by the available generators and various modes of compensation are applied—some involving special purpose plant. Timely introduction of the compensation is sssential for proper voltage control, and it entails advance instructions based on predictive studies of system needs.

One of the principal problems in the continuing development of high-power density systems, and particularly multi-purpose systems, is that of containing the short circuit stress imposed on switchgear with growth in power density. With that problem, there is the attendant one of ensuring reasonable confinement of the effects of system disturbances. The desirable features are means of containment that do not impair stability and yield a reasonably uniform stress profile over the network as a whole, and over the range of system loading.

SUB-TRANSMISSION

Because of their influence on voltage quality, the technical constraints imposed by short circuit containment and voltage control, can extend to the design and develop­ment of the subsidiary networks and their supply connections from the Supergrid. However, increasingly now, economics are becoming the primary influence. The various inter-network transformer supply arrangements are portrayed in figure 4.

The transformation from 400 kV to 275 kV is by auto-transformers in units of 500, 750 and 1000 MVA; subsequent transformation to 132kV is also by auto-transformers in units of 120, 180 and 240 MVA; the further transformation to Area Board distribution voltages, which are mostly 33 kV and 11 kV, is by double winding transformers, in a range of standard unit sizes. Increasingly, supplies are being given direct from 275 kV to the Area Boards at distribution voltages, either 66 kV or 33 kV, by double winding transformers-substations are somewhat larger but fewer are needed and, in general, siting problems inside the conurbation areas are eased. Outside the conurbation areas the transformation is direct to 132 kV by auto-transformers mostly in 240 MVA units but with 360 MVA units in prospect. There is so far little direct transformation for 400 kV to Area Board voltages.

Short circuit levels in these subsidiary networks can be held to desired values, by suitable subdivision and by use of an appropriate number and size of supply connections to each subsection. By this means, short circuit levels can be held indefi­nitely to one chosen value, load growth being accommodated by progressively creating more subsections, a development practice which offers distinct advantages. The transformer affords a convenient and generally economic means of effecting the subdivision as shown in figure 5. The left-hand arrangement illustrates the principle as applied to the 275 kV conurbation networks, which for security reasons are supplied from two or more geographically separate points in the main 400 kV network. The diagram shows four such supply feeds and it also illustrates how further subdivision of a network subsection can be affected to accommodate generation. The right-hand arrangement in figure 5 illustrates the principle as applied to the subsidiary 132 kV and Area Board distribution networks where, in general, supplies from the higher voltage network are given through a single transforming point.

Choice of short circuit value entails consideration of the value with all supply circuits in service, which decides the switchgear rating, and the value following circuit outage, which conditions the regulation and so the voltage quality. The closer the two values can be brought together the higher the load that can be supplied through switchgear of a given rating. This implies the use of multiple circuits of appropriate size, according to security criteria, an arrangement which tends to conflict with exploitation of scale. Thus, the problem resolves to balancing switchgear cost against

25

circuit costs. CEGB pratice has yielded broadly the following relationships between source unit size, power level and short circuit level.

Table 2

Voltage kV

400 275 132 132 66 33

Short Circuit Level MVA

35,000 15,000 5,000 3,500 2,500 1,000

Source Unit Size MVA

2,000- 1,000 1,000- 500

360 - 240 240- 180 180- 120 120 - 90

Power Level MVA

6,000 ■ 2,000 •

960-640-480-120-

8,000 4,000 1,280

960 640 180

PS. S/Stn

INTERCONNECTED 4 0 0 / 2 7 5 k V NETWORK

AVERAGE INVESTMENT

CATEGORY £/kW

POWER STATION ~ CONNECTIONS"

2.

T -SUPERGRID 6

T l32kV BULK

" SUPPLIES 5

!32kV

S-J -SUB-TRANSMISSION—10

AREA BOARD DISTRIBUTION

Fig. 7 — Transmission investment.

The general similarity of the relationships at the different voltages is readily apparent.

The above short circuit levels refer to the balanced three-phase values. Unbalanced earth faults can yield higher equivalent values in some parts of the higher voltage networks (400, 275 and 132 kV) and so constrain the permissible three-phase value.

26

This problem stems from the progressive increase in total number of transformers, which at these voltages have their neutral points solidly earthed to suit system insula­tion policy. It assumes more significance in underground sections than in overhead sections of networks, because impedance to the flow of earth fault currents is much less with cables than with overhead lines.

TRANSMISSION INVESTMENT

The graphs in figure 6 trace the general relationship between specific costs and unit size for various items of transmission equipment. The cost reduction with size is clearly shown but the diminution in the economic return with increasing size, in the absence of further technological advancement, is also clearly in evidence. This fact tends to react against employment of the larger units unless a high enough utilisation of the equipment capacity can be foreseen at the outset. On the other hand, practical considerations tend to encourage use of the larger units and in the case of overhead lines these considerations may prevail, particularly as the improve­ment in transmission efficiency offers partial recompense.

The net result of the foregoing technical and cost considerations on capital investment in main and sub-transmission facilities in England and Wales is shown by figure 7. Here the facilities are portrayed as falling into the following categories:

(1) Facilities for connection of new generating plant; (2) Construction and strengthening of the interconnected network; (3) Supply connections to the sub-transmission networks; (4) Sub-transmission facilities to supply the distribution networks.

The investment in each category is expressed in terms of expenditure per kW of new generating plant which the facilities are planned to accommodate. The figures are averages for England and Wales over the past decade.

The investment in the interconnected network covers its three functions of inter­connection, bulk transmission and high-power distribution. The expenditure incurred on each is not readily separable because they share so many common facilities; however, one-third or £2 per kW is judged to be a reasonable allocation to the inter­connection function. Although the figures are averages over 10 years, they have not varied widely when allowance is made for inflation. The reason is that a high propor­tion, around 60%, has been incurred on substation equipment, requirements for which are closely related to load growth. Of the remaining investment on transmission circuits, about half has been on underground cables, although they account for but a small fraction of the total transmission circuit mileage.

The figures show how the effects of scale are reflected in higher specific costs for sub-transmission, where the requirement is for reduction in power levels to manageable values for distribution.

The trend in these costs in the future will be very sensitive, of course, to the degree to which undergrounding becomes necessary for amenity or practical reasons, unless there is some technological breakdown which drastically brings down cable costs.

DEVELOPMENT POTENTIAL AND PROBLEMS

Looking ahead now to the character of these needs in the future system, trans­mission requirements and problems will be conditioned by the practicalities of power

27

station siting. They will be minimised by suitable dispersal of the newer generation —still in large stations but deployed in the emerging areas of unsatisfied demand. The supergrid has ample potential for accommodating the output of such stations and distributing it. The networks offer good coverage of the load areas and the lines can carry large amounts of power, over 4,000 MW in the case of the newer 400 kV lines; but to do so they must not be over-long and must have access to adequate dynamic sources of reactive power. Thus, distance is a key factor.

Fig. 8 — Areas of potential power deficit.

Preservation of the interconnection function is foreseen as being vital, not only to minimise capital outlay on generating plant, but also to exploit new types of plant with varied cost and operating characteristics. Interconnection requirements can be expected to increase in step with system growth, presuming a like increase in generating unit size, but even so, they will remain within the capabilities of 400 kV transmission for at least a quadrupling of the present demand of approximately 40,000 MW. The larger power exchanges for pooling involve transmission distances of up to 100 miles and preservation of the requisite capacity will require progressive contraction of the optional bulk transmission function; the supergrid will eventually revert to the dual purpose function of the original 132 kV grid, that is, interconnection plus high-power distribution.

Fuel transportation costs will have reducing significance in a future system, comprising an increasing proportion of nuclear plant, and so will have little influence

28

on deployment of the newer generation. The greatest restriction will be availability of adequate water for cooling purposes; and it is in the inland areas of this country, which are devoid of really large natural waterways and where there are competing claims to the water becoming available from conservation schemes, that this restriction could be the principal barrier to dispersal.

These inland areas are sketched and shown in the map in figure 8. In the north they embrace the inland conurbations of South-East Lancashire and West and South Yorkshire, and in the central area, the southern parts of the West Midlands conur­bation and the Western parts of the London conurbation.

The Northern area accounts for about half the consumer demand in Northern England, or roughly 7,000 MW at the present time. It will be traversed by three 400 kV lines which, as already described, will be supported by generation at each end. These facilities are foreseen as satisfying potential needs for a longish time but new generation central to this area would extend the life of the transmission still further.

The other potentially high deficit area embraces the central counties. It covers a considerable geographical area with a large but fairly widely dispersed consumer demand, at present over 10,000 MW and already supplied mostly by transmission from external sources. Planned development will permit the same pattern of supply to continue and with somewhat increased support from external sources. Most of this support will come from the Trent Valley area, with an increasing contribution from East Anglia and South Wales. The eventual requirement for more local gene­ration will arise sooner than normal load growth dictates because of the need to contract the optional longer distance power movement in order to preserve adequate transmission capacity within the established network for the now essential inter­connection function.

With reasonable dispersal of generation, and using traditional technical aids, the problems of short circuit containment and stability are fairly readily reconcilable and manageable within desired security criteria. These conflicting problems become more difficult to reconcile when there is need also for substantial movement of increa­sing concentrations of power. Even within the context of dispersed generation, problems of supplying the conurbations could remain because of the particular difficulties in finding suitable sites for large power stations in built-up areas—due sometimes to competing claims for the limited available land. Equally the conveyance of increasing quantities of power from outside sources could progressively present more and more technical and practical difficulty, even when using underground transmission.

It is this future possibility that has prompted the trial installation of direct current transmission in the London Area. One object of the trial is to demonstrate the dynamic behaviour of a d.c. transmission link as an integral part of the a.c. system, and to obtain better understanding of the technical constraints involved with the use this way of d.c, by reason of its functional dependence on the a.c. system. Another object of the trial is to advance the complex technology of a.c. to d.c. conversion and inversion, in an endeavour to reduce terminal substation dimensions and costs, these being at present the principal barriers to usage of d.c. transmission for the shorter distance underground power movement.

The difficulties in finding suitable sites for transmission substations within the conurbation areas are becoming equally formidable, and with the developing system, will place continuing emphasis on reduction of equipment dimensions and compact layouts to conserve costly space and facilitate strategic location of substations. High-voltage switchgear design has made appreciable advances in this direction of late, to the extent that physical dimensions of transformers and their ancillary plant now tend to determine spare requirements for substations.

29

CONCLUSIONS

It is perhaps worth noting by way of conclusion, that while there have been spectacular advances in the scale of electrical power systems, basic technology has remained unchanged since the advance of the three-phase system of supply; that is until recently. Now the newly developed high-voltage d.c. transmission is finding special purpose application in a. c. systems, to overcome technical constraints involved with very long distance overhead a.c. transmission and with the very much shorter distance underground a.c. transmission. Undoubtedly the potentialities of high-voltage d.c. transmission would be enhanced by an advance in technology yielding functional independence of the a.c. system.

Problems of power system development in this country will centre, in the longer term, on supplying the dense consumer demands in the inland conurbation areas from external power sources, on reducing substation siting requirements in general and in built-up areas in particular; and on providing strategic generation support to the 400 kV supergrid lines serving the central counties area and those serving the Lancashire-Yorkshire industrial complex.

30

SURVEY PAPER

PROBLEMS OF LARGE-SCALE REFRIGERATION

C. TREPP Sulzer Brothers Ltd., Winterthur {Switzerland)

Problemes du refroidissement a grande echelle

RESUME : On decrit les installations frigorifiques comprises entre 1,6 et 80 °K, avec des apports de courant electrique allant de plusieurs centaines de kW a quelques MW. On neglige ainsi les installations frigorifiques miniatures et les appareils frigorifiques de laboratoire. Ces installations doivent convenir a une exploitation industrielle continue, ce qui signifie qu'il faut preter attention a Ventretien, ainsi qu'a Vinstallation de secours et a ses elements.

Parmi les nombreuses sortes de cycles frigorifiques utilises actuellement, on examine les avantages et les inconvenients des cycles Stirling et Claude. Les progres dans Vapplication de la supraconductivite ont ete pousses tres loin dans le domaine de la conception des aimants et parmi les installations actuellement examinees et, dans certains cas en cours de construction, celles qui exigent une collaboration specialement etroite entre Vingenieur charge de la conception et Vutilisateur du froid retiendront specialement Vattention. On note en particulier le refroi­dissement des conducteurs creux avec de rhelium a Vetat supracritique, le fonctionnement des chambres a bulles a hydrogene liquide, munies d'un aimant supraconducteur, et les cavites supraconductrices frigorifiques au-dessous de 2°K.

On examine le cout des installations pour les appareils fonctionnant entre 2 et 20 °K. Des comparaisons semblables sont tirees d'un certain nombre de publications et Von fait ressortir ici quelques raisons des grandes differences des couts d'installation indiques. On s'etend aussi sur le rendement des differents systemes d''installation, c'est-a-dire sur la consommation speci-fique de courant.

1. INTRODUCTION

The methods usual today for attaining low temperatures utilize the thermodynamic properties of gases without exception. Over wide temperature and pressure ranges these properties are so well known that the thermodynamic conception and detailed calculation of refrigeration cycles making use of them is a routine exercise. Only for certain specialized problems, such as the heat transfer from high-pressure helium at low temperatures, are the calculation data still lacking.

Generally speaking, from the aspect of engineering and technology a large refri­geration facility is easier to translate into reality than a miniaturized plant, and the reason for talking about the " problems " of large-scale refrigeration is that these installations have not yet found widespread application. There are only very few examples of plants, or even projects, where the refrigeration capacity at 4.5 °K exceeds 1 kW.

2. REFRIGERATION CYCLE

At present, only refrigerating plants operating in the liquid helium range and the temperature range of 15-20°K can claim special importance for applications of potential interest. The use of cryo-resistors becomes economical only around 20°K, dictated by the behaviour of the electrical resistance in metals of high purity. And for the use of superconductors the temperature range of liquid helium is in any case advantageous. The majority of this paper concerns the 4°K range. It would be quite possible to extend the remarks to other temperature ranges as well.

31

Very often the question arises: why are the efficiencies of cryogenic plants in general so low? A few reasons are owed here (or one might call them excuses)—especially as these efficiencies play a meaningful part in the considerations of economy.

1 48,000 1 /- 300°

650 ' iJWvWWU

IWwflSi

4°

Stirling

hwfft^l

-T2)ln

Entropy

T,

Claude (Brayton)

Fig. 1 — Basic Temperature—Entropy—Diagrams of 3 most essential refrigeration cycles. At the left: Carnot; middle: Stirling (Kirk); right: Brayton, Claude or Ackeret—Keller (AK).

Both cycles are derived from the Carnot cycle (a), which is described by two isotherms and two isentropics. To achieve it in practice would be possible only at enormous expense. This will be immediately obvious from the values noted in figure 1. Isentropic compression of an ideal gas with specific heat ratio of x = 5/3 at 4°K and 1 atm. abs. would, with a final temperature of 300 °K, entail a pressure of some 48,000 atm.

On the other hand, the expansion of gas from 300 °K down to 4°K would mean increasing its volume 650 times over. These technically almost insuperable difficulties can be overcome by breaking the cycle down into a number of part cycles with smaller temperature ratios, but this would require complicated equipment with losses due to the number of intermediate heat transfer media.

Engineering has chosen a different approach, using the Stirling cycle (b) and the Claude or Brayton cycle (c). The ideal Stirling cycle works between two isotherms and two isochores, the ideal Brayton cycle between two isotherms and two isobars. Both are based on ideal refrigeration cycles, with which the Carnot efficiency can be attained in theory.

In its technical realization, the Stirling cycle works with reciprocating machinery for compression and expansion, which together with the regenerator makes possible in principle the heat exchange at constant volume. The heat exchange at constant pressure employed in the Claude cycle allows more choice with the mechanical equipment (compressor and expander), because the machinery does not directly participate in the heat exchange process. In contrast to the Stirling cycle, all particles of the refrigerant pass through the same process in the Claude cycle. Using it as an example, I shall point out some of the reasons for the relatively low efficiencies [1].

32

For a single-stage Claude cycle according to figure 2, the following relations are valid [2]: (ideal gas, ideal heat exchange with A T = 0)

'Icycle 'Icompr 'lexp 1 - T

n is - n a d

icompr \exp

*■ 'lexp 'Icompr

T = T/T ambient T = mean temperature at which refrigeration is supplied

At very low temperatures this means

'icycle

Hitherto attempts have been made to manage with as few plant components as possible with simple refrigeration cycles, in other words keeping the number of compression stages low for example. This leads automatically to relatively poor isothermal efficiencies, and figures of 50% or so are by no means rare. Turbine efficiencies reach values above 80% only in bigger installations, and 70% may be quoted as an average. But in fact there are further losses in addition, due to the heat transfer. This is set out in figure 2.

Losses and efficiency of cold gas circuit

1/'/K<p A"

Compressor

Expander

Heat exchanger

Fig. 2 — Approximative breakdown of losses in cold gas refrigerator depending on the temperature.

This diagram shows approximately the losses—and hence the efficiency—of an actual cryogenic plant. The heat exchange losses are of a similar order of magnitude

33

to those attributable to the turbine. The bottommost curve delimits the range in which at constant temperature difference in the heat exchanger refrigeration can be produced.

Only recently has the plant efficiency begun to play an increasingly important role, because the continuous operation of such equipment on a commercial basis is now being aimed at. Previously simplicity was the prime requirement in large cryogenic refrigerators, on account of its generally positive effect on operational reliability.

And so in the future it will be worthwhile investing more in the compressors of Claude cycles, in other words in the warm end of the plant, because they have a decisive influence on the plant efficiency.

Fundamentally similar considerations apply for the Stirling cycle. By rational standardization and judicious subdivision of the volumes of the working spaces in relation to each other, some companies have succeeded in raising the overall efficiency to a very high level: over 40% at 80 °K.

Efficiencies as high as these can be attained only in special developments, which are justified only where series production on a substantial scale is involved. In contrast, the Claude cycle has the great advantage of being able to manage with commercially available equipment as a general rule—apart from the expansion unit—and of allowing a large number of different cycles to be built up with small modifications, without special development costs. On the other hand, the Stirling cycle is suited for temperatures above 20 °K, primarily on account of the thermal properties of the metals though of course temperatures as low as desired can be reached by adding a JT stage. In this case the JT stage must be installed separately, whereas it can be integrated into the Claude cycle.

For transferring refrigeration—whether to a lower temperature level by means of a JT stage or over greater distances at any temperature—the Claude cycle is superior to the Stirling cycle. Such transfers of refrigeration can be accomplished without additional circulating equipment with the Claude cycle: the driving force is supplied by the main compressor.

The essential components of higher-capacity Claude plants are: reciprocating compressor, heat exchanger, cold gas turbine. The view is often advanced that turbo-compressors ought to be used in the interest of operational reliability. Now if we want an isothermal efficiency of say 75% from a turbocompressor giving compression from 1 to 15 atm, then on the most optimistic reckoning we shall need a machine with about 32 stages and sevenfold intercooling, and its blading will have to yield a poly-tropic efficiency of 83% for each stage. The mean throughput quantity would then be around 5 kg/sec, which means that a machine like this will only be developed when the drive input can amount to about 12 MW; this would make the refrigeration capacity at 4.5 °K about 23-45 kW.

Small, high-speed turbocompressors have been built or are under development, but their overall efficiency is not very encouraging—in some cases lying well below 30%.

For heat exchanges usually some type of extended surface exchange is used. As expanders for large-scale refrigerators usually turbines are chosen (the efficiency

rises with throughput). Gas bearings for large scale are inferior to oil bearings. Oil bearings are more

robust and allow application of loading devices such as oil brakes which can operate in a very wide range of capacities. This is especially important for fast cool-down of an installation.

3. PLANT AND OPERATING COSTS

In recent years, certain figures on plant costs have been published by various authors. These are reproduced in figure 3, taking the typical case of the 4.5 °K plants.

34

The average value is represented by the dotted line. It shows that at capacities smaller than 100 W the plant price varies with the .4th power, at 1 kW with the .5th and at 10 kW with about the .6th power of the cooling capacity of the refrigerator.

0 1 0 2 0 5 1 0 Capacity kW -

2 0 5 0

Fig. 3 — Relative capital cost of 4°K refrigerators. Numbers: see references.

The scattering is so wide that only very approximate economy calculations can be performed on the strength of these data. There are various reasons for this scatter:

— Liquid nitrogen precooling is not always included in the calculation. Besides the cost of the actual liquid nitrogen, there is the equipment for handling it.

— Calculating the performance of the liquefiers as refrigeration capacity. This is not always admissible, and especially with liquid-nitrogen-precooled installations, leads to sometimes incorrect data, because the influence of liquid nitrogen precooling in liquefaction and refrigeration is not the same.

— Some plant prices are quoted ex works without erection. — Some plant prices are quoted fully commissioned. — Only a few of the figures refer to actual completed installations, but most of them

to projects and tenders.

The costs of the components of a certain type of plant will not vary equally over the whole capacity range.

35

For the Claude plant, which is easiest to engineer for medium capacities, the total costs emerge roughly from the sum of the components as shown in figure 4.

Cost breakdown in percent of a Helium Refrigerator Plant

^mm^^

j/0/0

00*

milllM'L

m00^

IV

0////g/0M ' 1 III Hi'1

III

000^

II

00/M W/////M

■000WfM

00gMmmf^m W//////M

I Compressors and motors II Cold box and expanders III Controls IV Installation, start up

1 I I I . I • L 0 1 2 3 4

Capacity kW > 06692007 0

Fig. 4 — Approximative cost breakdown of medium size helium refrigerators.

The areas I-IV are the approximate percentages for the compressor station, cold box, control arrangements, and erection plus commissioning.

No figures at all can be given on other costs, which are governed by the particular installation, as these may vary within wide limits.

And so from what has been said, it becomes evident that it is not easy to quote really reliable cost figures for future cryogenic refrigerators. In principle the curves published by Kurti in 1967 still hold good today. But cost comparisons are rendered difficult above all by the fact that the requirements for refrigerating plants in electrical engineering are not known with accuracy. The figures given previously refer essentially to one-off jobs, where plants operating at 4.5 °K are involved. But the refrigeration industry now anticipates a substantial upsurge of its sales potential from applications in electrical engineering. This in turn will mean that it ought to be possible to produce future cryogenic plants in medium-sized runs of 10 to 50 units, which is bound to cut manufacturing costs to a certain extent. A further price reduction resulting from series production may be expected from narrowing the dimensional margins, which are often very generous at present.

Plant efficiencies, in so far as they are known today, lie in the region of 40% for 80°K plants, 15 to 35% according to plant size for 20°K, and up to about 10 to 20% for 4.5 °K. But efficiency and plant price cannot be viewed divorced from each other.

36

To enable the question of how important the efficiency is to be answered, a calcul­ation has been made. The result is plotted in figure 5.

3001 1 1 1 1 1 1

% I |

2 0 0 — x ^ ^ ^ ^ ^ "

100 {—/- '

fll 1 D = 20years

III 2 D = 15years

3 D = 10years SULZER 06692012 e

8 10 15 20 30 >/%

Fig. 5 — Tolerable capital cost variation of 4°K refrigerators depending on the efficiency.

The following data were taken as a basis:

Refrigeration capacity 1 kW at 4.4 °K Plant price £150,000 Efficiency 10% Amortization period D = 10-15-20 years Interest 6% Utilities 5l/2 centimes/kW.h

(this includes electricity, water, etc., but not maintenance, as this does not depend on the plant efficiency in any case). (1 Swiss centime = £ 0.001).

For such an installation, annual operating costs work out at about £ 60, 000, including capital charges. The plant efficiency can now be varied in order to find out how far the plant price may be raised together with the efficiency without putting up the annual costs. This can be read from figure 5, in which the plant price assumed previously is taken as 100%.

It is quite evident that with the low efficiency assumed, only a slight increase will justify a massive rise in the plant costs.

This is still clearer in figure 6, in which the first derivative from the curve in figure 5, is plotted against the efficiency. It can be seen that there is hardly any point in trying to raise the efficiency from 20 to 30%—anyhow difficult and expensive in this region—though it does appear necessary to lift it from 10 to 20%.

Figure 7 shows a selection of plant efficiencies from 4.4 °K refrigerators.

37

30

20

10

^ ^ 2

1 D = 20years

2 D = 15 years

3 D = 10 years

" " - —~ SULZER

8 10 15 20 30 ,,%

Fig. 6 — First derivative of curves figure 5.

0 1 0 2 0 5 1 0 2 0

Capacity kW ►

5 0 10

Fig. 7 — Efficiency of 4°K He-refrigerators. 1. see ref. [7]; 2. This work.

38

4. STANDBY EQUIPMENT, OPERATIONAL RELIABILITY

The standby equipment required in a particular case may also have considerable influence on the price. Especially with cable cooling is this so, for duplication of the auxiliary equipment is a standard requirement here. It is hardly likely that the refri­geration capacity will be provided in duplicate too. Instead one of the following two courses is probable:

— duplication of the moving plant components, such as compressors and turbines; — overdimensioning the refrigerators, so that if one plant fails, two or three neigh­

bouring plant units can take over its load.

With the values set out in figures 3 and 4 it should be fundamentally possible to estimate which variant is more favourable in general— the overdimensioned plant, running mostly at part load, or standby equipment for the moving plant components. But the values show so much scatter that no clear conclusion can be drawn, and the final decision will probably be taken on the strength of the operational requirements and not only because of the price of the refrigerator.

Exhaustive investigations and actually completed installations will be needed before it can be decided what kind of standby equipment will satisfy the particular operating conditions. The standby equipment is governed in many cases by one basic requirement: the refrigeration must be continuously available, without inter­ruption. Complete duplication of the systems means that the standby equipment too, including the transfer lines, must be kept constantly at the operating temperature —either by cooling from the main circuit or by running at part load. Part load operation, with no refrigeration output, will probably be uneconomical.

But in the author's opinion, duplication of the mechanical components—com­pressors and expanders—should suffice here. For complete autonomy can only be attained if other systems—like power and cooling water supplies—are also duplicated and installed independently of each other.

The optimum solution—at any rate of price—is considered to be the overdimen­sioning of the refrigerating plants where there is more than one anyhow, so that two or three neighbouring units can take over the duty and the load of the defective unit, or else duplication of the moving components only. This will guarantee uninterrupted cooling at any rate, except under disaster conditions.

As for the costs of these different ways of providing reserve capacity, it may be expected that the total costs of a single installation will be exceeded by about

a) 70-90% if the installation is duplicated; b) 30-40% if the mechanical components are duplicated; c) 25-35% with overdimensioning by 50% d) about 25% with overdimensioning by 35%.

For the operational reliability we have to consider the reliability of the individual plant components. As was said at the beginning, as long as there are so few plants in operation it is hardly possible as yet to obtain meaningful data on the MTBF of individual components.

Anyway there is some indication that these may be quite long, as there are examples of refrigerators in which turbines and compressors have so far run 15,000 hrs without any maintenance.

5. EXAMPLE

One particular plant type might be mentioned, namely the BEBC refrigerating plant

39

An interesting variant for solving the problem of operational reliability without incurring substantial additional expense is applicable to liquid hydrogen bubble chambers which are equipped with superconducting magnets. As an example, the refrigerating plants ordered by CERN for the large liquid hydrogen bubble chamber and its superconducting magnets may be described.

The principle data of the BEBC refrigerating plants are as follows: refrigeration capacity: 25 kW at 22 ° K

1.5kWat 4.4° K 2.5 g of liquid helium per second. (701/hr) It has been suggested by CERN that the two refrigerating plants should be

combined, so that the helium circuit is operated as the primary circuit. It will supply refrigeration at 4°K, precool the hydrogen required for cooling the chamber, and also deliver a small quantity of liquefied helium for cooling the electrical current leads.

Priority will be given to the helium refrigerating circuit, because it keeps the magnet cold, and the operating characteristics of this magnet are not yet known in detail. In the event of any malfunctioning of the refrigerating plants it is proposed to do without the chamber cooling first, and continue operating only the magnet itself.

Plant diagram (fig. 8)

Refrigerators for CERN BEBC

Bubble Chamber

Magnet

Fig. 8 — Basic flow diagram of the CERN-BEBC refrigerator (for big European bubble chamber).

40

In this diagram the principal components are shown schematically, i.e., one hydrogen compressor and two main helium compressors. The refrigeration is produced by Stirling machines at about 80 °K, and by cold gas turbines (duplicated for safety reasons) at temperatures of 40 and 12°K respectively. In addition, a small auxiliary compressor is installed; it is brought into service only in the event of certain plant components failing. The main helium circuit (in which the refrigeration is produced with turbines running on oil bearings) is a supercharged process, leading to relatively small heat exchangers and compressors. The turbines are duplicated, as are also their lubricating oil supply systems. If a helium compressor and any turbine should fail, after shutting down the cooling on the hydrogen side the full refrigeration output can still be applied to the magnet at 4.4 °K. In the event of a total breakdown of all machinery, the magnet can still be kept cold by feeding in liquid helium from a Dewar vessel for a time. One advantage of this process which should be emphasized is that a very large quantity of helium is available for cooling down the magnet, so that the requisite high refrigeration capacity can be attained even with relatively narrow temperature differences.

This installation bears much more resemblance to the kind which might find application in electrical engineering. Usually the requirements will not be so simple that some refrigeration capacity has to be produced at a certain temperature. Very often the process is characterized by the starting and cool-down conditions. In the case in question for example, up to lOOkW refrigeration capacity must be available for cooling the magnet down, and the plant configuration must be such that it can also be used to liquefy the cooling and experimental liquids needed. About 15, 000 Nm3

of helium and 40, 000 Nm3 of hydrogen must be liquefied before the experiment begins. This raises a special purification problem, because the contaminants separated out during liquefaction must not be present in the system during subsequent continuous operation. Similar problems will probably arise with cryo-cooled cables too, because here also exceptionally long spells of continuous operation will be demanded.

REFERENCES

[l] S. ERGENC and J. HANNY, Sulzer Techn. Review, 4 (1963). [2] P. GRASSMANN, Schweizerische Bauzeitungl9, (Nov. 1961), pp. 790-800. [3] A.R. WINTERS and W.A. SNOW, NBS Rep. 9259 by T.R. STROBRIDGE, D.B. MANN,

D.B. CHELTON (Oct. 1966), fig. 6. [4] C. TREPP, Bulletin SEV, 57 (1966), 18, pp. 817-823. [5] N. KURTI, Supercond. Conf. CERN, Geneva (1968). [6] G.P. COOMBS, NBS Rep. 9259 (Oct. 1966), see 3, fig. 7. [7] T.R. STROBRIDGE, D.B. MANN, D.B. CHELTON, NBS Rep. 9259 (Oct. 1966), fig. 7.

41

SURVEY PAPER

CRYOGENIC POWER TRANSMISSION

S.H. MINNICH and G.R. FOX Research and Development Center, General Electric Company, Schenectady (U.S.A.)

Transmission cryogenique

RESUME : On a etudie de facon assez approfondie les systemes de transmission d'electricite par cables souterrains cryogeniques en utilisant des conducteurs normaux dans Vazote et Vhydro-gene liquides et des supraconducteurs dans Vhelium liquide.

Vhydrogene et Vazote liquides presentent tous les deux de bonnes proprietes dielectriques, de sorte que Vutilisation de ces fluides dans des systemes dielectriques semblables au papier huile classique peut etre utile pour les systemes de cables cryogeniques a tension elevee. La conception du conducteur exige des techniques speciales pour reduire Veffet pelliculaire dans les cables a courant alternatif. De plus, il faut assurer une protection electromagnetique pour eliminer les pertes en ligne.

On a aussi etudie les systemes supraconducteurs a courant alternatif. Pour reduire les pertes en courant alternatif, on propose habituellement Vutilisation de niobium (supraconducteur mou). Comme les pertes dielectriques dans les systemes desolation a remplissage d'helium seraient probablement excessives, on propose habituellement Vutilisation d'une isolation par le vide ou d'helium liquide libre. Pour ces raisons, la plupart des conceptions avec supraconducteurs ont ete des systemes a basse tension.

Jusqu'ici aucun systeme rta present e de superior it e technique nette. Le choix du systeme et de la disposition semble reposer sur la comparaison des coiits ce qui exige des etudes detaillees de la conception pour etre evalues de facon satisfaisante.

During the past year, the Research and Development Center of the General Electric Company has undertaken a study of cryogenic underground power trans­mission systems under the sponsorship of the Edison Electric Institute and the Tennessee Valley Authority, as a part of the Electric Research Council's Underground Transmission Research and Development Program. This paper is based partly on the work carried out in this program.

INTRODUCTION

The electric utility industry faces a growing need for improvement in the technology of underground transmission systems. The blocks of power being generated and transmitted are steadily increasing, and there is increased pressure to eliminate the large overhead lines in suburban areas. Although it seems evident that the cost of transmission will continue to favor overhead systems, the amount and proportion of underground transmission will continue to increase.

Conventional high-voltage underground cable systems use electrical insulation consisting of a wrap of kraft paper (up to an inch or more in thickness) impregnated with high quality insulating oil. In the United States, the practice is to bury a steel pipe through which three cables are pulled in lengths up to approximately a half mile. After the cable has been installed, the pipe is filled with oil under pressure to preserve the electrical integrity of the insulation. This system is known as "pipe-type" cable.

In Europe, so-called "direct-burial" cable is used, in which three separate cables, protected by a lead or aluminum sheath, are laid side by side, separated by several feet. Direct-burial cables are pressurized with oil admitted through a central duct.

43

Both systems are limited in capacity by the operating voltage and by the dissipation of losses by conduction through the electrical insulation and the surrounding earth. Direct-burial cable has somewhat higher capacity because of its better heat dissipation. The highest capacity cable system in the United States is approximately 500 MVA at 345 kV, whereas European systems have capacities in the range of 1000 MVA.

A cryogenic cable system will be complex, requiring vacuum-insulated piping, continuous refrigeration, etc. Hence, cryogenic systems will find their utility by extend­ing the capacity of conventional systems beyond their economic capability rather than by the reduction of losses, per se. Because of the probability that overhead transmission will continue to be the least expensive system, it is envisioned that cryogenic systems will form relatively short links in an overall system, with line lengths of 10 to 50 miles. For this reason, this discussion is directed toward those cryogenic systems which will match directly the capability of overhead systems, and work has been concentrated on high-voltage a-c systems with capacities beyond present practice, namely 2000 MVA and higher.

CD

£ CD O

O O

I

0.1 t

tz 0.01 en

LU

0.001 20 40 60 80100 200 300 TEMPERATURE - Degrees K

Fig. 1 — Variation of Resistivity with Temperature at Cryogenic Temperatures.

The potential attractiveness of cryogenic electrical applications is based on the significant increase in electrical conductivity of metals at low temperatures. Tha resistance ratio, defined as the ratio of the resistance at room temperature to that at low temperature, can be 500 to 1000 for pure copper and aluminum. Figure 1 shows the trend in resistivity with temperature. The resistivity consists of two components, lattice vibrations which decrease rapidly at low temperature and residual impurities which are constant and form the "floor" of the curve in figure 1. For resistance ratios greater than 500, 99.999 percent purity is implied for copper and aluminum, (Such metal is available in pilot quantities.) Commercial grades of copper have resistance ratios up to 200.

44

Counteracting the advantage of high conductivity is the poor efficiency of refri­geration at cryogenic temperatures. This is due to the fundamental laws of thermo­dynamics coupled with practical achievable efficiencies in cryogenic refrigerators. A summary of the relative resistance of conductors and the efficiency of refrigerators is contained in table 1.

Liquid

Oil Nitrogen Hydrogen Helium

Table 1

CRYOGENIC SYSTEM PARAMETERS

Operating Temperature

{Degrees Kelvin)

293 77 20

4

Resistivity

1 1/8 1/500

Superconductors

Refrigeration Ratio ( Watts Input 1 Watts Load)

Theoretical

1 3

14 75

Practical

1 6- 10

40 - 100 300 - 1000

It can be seen from Table 1 that at liquid nitrogen temperature gains in conductivity are about balanced by the refrigeration penalty, while at liquid hydrogen temperature the increase in conductivity is about ten times the refrigeration penalty.

The above figures depict the theoretical possibility of trading losses and refriger­ation power under d.c. conditions only. Alternating-current conditions introduce additional losses. With high-conductivity normal metals, skin effect and eddy current losses must be considered. Superconductors also exhibit a.c. losses. In particular, the new high-field superconductors (such as niobium-titanium and niobium-tin), which are being used in superconductive coils for intense d.c. fields with considerable success, are lossy under a.c. conditions. These losses are of such magnitude that, coupled with the helium-temperature refrigeration penalty, the use of such materials under a.c. conditions is unattractive. On the other hand, niobium is a soft supercon­ductor whose critical field is in the range suitable for cable applications, and its a.c. losses can be made low. Hence, liquid helium cable systems using niobium can be considered.

Because the dielectric system of a high-voltage cable is of equal importance with the conductor characteristics, the dielectric properties (including losses) of materials at low temperatures are of concern. Finally, the economics of vacuum-insulated piping and refrigeration must be taken into account. The following discussion considers each of these factors in detail and shows how they can be integrated into various cable configurations.

DIELECTRICS

The existence of a satisfactory dielectric system is a key factor in a high-capacity cable system. Dielectric strengths of the cryogenic fluids are, fortunately, in the same range as those of good insulating oils. Also, dielectric losses are of prime significance in cryogenic systems because of the large refrigeration penalties involved.

Measurements of the dielectric strengths of the cryogenic fluids have been made using standard sphere-gap techniques and concentric cylindrical electrodes. The results of these two techniques are summarized in table 2. The loss tangents of the cryogenic fluids themselves are too low to be measurable.

45

Table 2

BREAKDOWN STRENGTH OF CRYOGENIC FLUIDS

Fluid Condition

1/2-inch spheres 0.025-inch gap-1 atmosphere Cylindrical 0.046-inch gap-1 atmosphere Cylindrical 0.046-inch gap-6 atmospheres

N2

1 120 370 650

AT 60 HERTZ

H2

1 000 260 650

He

560 300 500

Fig. 2 — Dielectric Test Specimens.

46

It is a well known fact that increased surface area of the electrodes leads to lower dielectric strengths, mostly by reason of statistics. The data in table 2 seem to confirm this trend. In the tests with concentric cylinder electrodes, the system was pressurized, leading to increased strength, as indicated. Presumably the increased strength is caused by the suppression of bubbles in the fluid. All data were taken at the normal boiling temperature.

Conventional insulating oil is seldom, if ever, used as a free fluid. Instead, barriers or fillers are used which have the effect of increasing the dielectric strength consi­derably. In conventional cable the barrier is paper, which is carefully taped onto the conductor and impregnated with insulating oil. It is essential that complete impregna­tion be obtained; the taping procedure assures that extended flaws in the paper structure cannot exist (i.e., local imperfections do not line up).

By analogy with oil-paper technology, it was reasoned that cryogenic insulation would be most effective with a similar fluid-filler system. Tests have been made on cylindrical electrodes taped with conventional insulating paper and with some of the newer synthetic papers. Loss tangents and breakdown strengths were measured. Typical results for several materials in liquid nitrogen and liquid hydrogen are shown in table 3. The samples used had a 0.040-inch thickness of dielectric material, and the maximum test voltage was 50 kV. Typical test samples are shown in figure 2.

Table 3 BREAKDOWN STRENGTH AND LOSS TANGENT FOR INSULATION MATERIALS

Liquid Nitrogen Liquid Hydrogen

Loss Tangent Breakdown Loss Tangent Breakdown Material at 500 Volts Stress at 500 Volts Stress

per mil (Volts per mil) per mil (Volts per mil)

Calendered Tyvek (duPont)

Spunbonded Polyethylene Fibre 0.0001 1 400 0.00014

Nomex (duPont) Aromatic Polyamide Fibre 0.002 1350 0.0009 Wood Insulating Paper

(Union Mills) 0.002 1 150 0.0004

1) Pressure: 75 psig. Relative Humidity: 50 percent

In many cases, the loss tangent increases with the voltage stress, and other complex effects are observed. Therefore, losses cannot be completely summarized with a single quotation of loss tangent, although the numbers in table 3 are felt to be achie­vable at a stress of 500 kV.

While the data obtained to date have been favorable, the values cannot be extra­polated directly to higher voltages. Until the time when test values are obtained at higher voltages, conclusions based on the use of these data are tentative. Where fluid filler systems are considered, use of the present data in design studies results in design stresses several times lower than those measured if dielectric losses are held to a consistent value. (The losses are a function of stress as well as of loss tangent.) Where free-fluid dielectric systems are considered, the measured values have been reduced arbitrarily by a factor of about three, in order to allow for reduced strength at high voltage. Cost and other studies based on these assumptions of voltage strength will indicate which configurations are most attractive. Favorable results from these studies will require confirmation by actual test at high voltages.

1 400

1 250

1 250

47

Since vacuum is sometimes proposed as a cryogenic cable insulation system, its characteristics merit some discussion. The use of vacuum is predicated on the suppo­sition that the vacuum is there anyway for the thermal insulation and that vacuum insulation is inherently lossless. At present, vacuum insulation is used mainly in particle accelerators or microwave tubes where it is the only insulating medium available. Achievement of high breakdown voltages across a vacuum requires the use of polished electrodes, surface conditioning, and extremely clean vacuum systems. In a recent survey, an authority on vacuum [1] states: "The voltage strength in vacuum soon reaches a nearly equilibrium value as the interelectrode gap is increased. This 'total' voltage dependence, recognized 30 years ago, is far more pronounced in vacuum than in solid or gaseous dielectrics. " In another place, this source says: "It is still difficult to insulate much more than 300 kV across a single gap of plane polished elec­trodes of moderate area (100 cm2) separated by 25 mm." The 300 kV referred to is a d.c. level, corresponding to 212 kV, RMS. These statements are interpreted as meaning that low-voltage systems may be possible, given sufficient development, but that systems of 345 kV and up may not be physically possible under any circum­stances.

In addition, vacuum insulation is not really inherently lossless. Field emission currents occur at high stresses; in normal practice they are a few tens of microamperes per square centimeter. At 200 kV, one microampere per square centimeter (RMS) represents a loss of 0.2 watt per square centimeter. As will be shown later, the desired surface loss for superconductive cable systems is of the order of 10 microwatts per square centimeter. Even in a liquid nitrogen system, typical conductor losses amount to something like 0.02 watt per square centimeter. Therefore, emission current levels must be reduced by several powers of ten. Recent measurements [2] suggest that this may be the case at low temperatures. Reduction in prebreakdown currents of a factor of 104 were quoted between room temperature and liquid helium tempera­ture. Various coatings on the electrodes are also helpful in this respect.

Hence, while there is some hope that vacuum insulation could be developed; at present, it seems safe to say that considerably more research will be required—more than that required for other fluid systems. If vacuum will be limited to the lower voltage systems, this may be disadvantageous in integrating a cryogenic cable into a total power system.

CONDUCTOR CONFIGURATIONS

Conductor configurations at liquid hydrogen temperatures are strongly influenced by skin effect. At high frequencies or with high conductivities, the current flows in a thin layer on the surface of the conductor, decaying exponentially with depth. The skin depth or exponential decay constant is given by:

5 = ^- V(P//)X10+9

where the skin depth is expressed in centimeters for resistivity in ohm centimeters. At 60 hertz and with a conductivity 500 times that of copper at room temperature, the skin depth is only 0.015 inch. This means that, if a cylindrical conductor is used, it may as well be in the form of a hollow tube—roughly two to three skin depths thick. The resistance of such a tube can be calculated by assuming that the current is uniformly distributed in a layer one skin depth thick.

More effective conductor utilization may be obtained by stranding, provided the strands are properly transposed. Transposition means the geometrical arrangement

48

of the strands so that the inductance of the strands will be equal, thus they carry equal currents. A simple hexagonal stranding sequence is the most elementary form of transposition, i.e., six elementary strands are wound around a central core (not used as conductor); six of these bundles are wound around a larger core; six of the resulting bundles are wound around a third core; and so on. In high-frequency work this configuration is called "litz " wire; its space utilization is poor if many strands are required. The individual strands are subject to losses caused by eddy currents flowing within them. Magnitudes of the eddy current effects are illustrated by the conductor model shown in figure 3. In this figure, a hollow former is wrapped with bundles of transposed strands (hexagonal stranding is shown in this example). Each strand is considered to be exposed to an ambient a.c. flux density produced by the total current in the conductor.

Fig. 3 — Stranded Conductor Concept.

It can be shown that for round strands, whose diameters are of the order of the skin depth or less, the eddy current losses per unit volume are given by:

Q = — x 1 0 " 1 6 watts per cubic centimeter (1) 32 p

49

where co angular frequency, second"x; d strand diameter, centimeters; p resistivity, ohm-centimeter; B local flux density (due to total current in the conductor), gauss.

For calculating total losses, B2 must be averaged over the conductor. For a hollow conductor with a wall which is thin compared to the bore, B2 = 1 / 3 B2

r av ' m Then, the sum of the eddy and conduction losses per unit length is:

l ^ + A ( B ^ M ) ^ x l 0 . 1 6

A 96p

where A = total metal area

^ 0.2 J2 Irms Since Bm = *- (3)

r (r - conductor radius)

^ . . . r^Ajs^xurq (4) LA 12 pr2 J

The quantity in parentheses is the a.c. resistance per centimeter length. The conduction and eddy losses vary with area and resistivity in opposite ways;

i.e., increasing the area and lowering the resistivity decreases the conduction loss but increases the eddy loss.

It is a paradoxical fact that, for a given strand diameter, adding strands will actually increase losses after a certain point. There are various ways to optimize a conductor design with respect to strand diameter and the number of strands. It can be shown that the minimum number of strands is involved if the eddy current losses are set equal to half the conduction losses. If this is done and the algebra performed, the a.c. resistance of the conductor turns out to be:

, - c o d x K T 9 , RflC = yji. ohms per centimeter

where r is the radius of the conductor. Figure 4 shows typical values for strand diameter and the number of strands as

a function of a.c. resistance for a hypothetical conductor made of wrapping transposed bundles of strands on a three-inch former as shown in figure 3. The middle curves (fig. 4) labeled " number of strand layers," refer to the hypothetical number of radial layers of strands; it gives one some feeling for the complexity of the trans­position required. For comparison the resistance of a solid three-inch tube, having a resistance ratio of 500, is 6 x 104 ohms per mile. Thus, it is seen that the resistance can be reduced by stranding, but the reduction by more than a factor of two or three requires many fine strands and a complex transposition scheme.

SHIELDING

If three cryogenic conductors are placed directly in a superinsulated pipe, the a.c. field caused by the phase currents will induce eddy currents in the pipe wall. It

50

can be shown that these losses, in a material such as stainless steel, would constitute a prohibitive refrigeration load. A high conductivity metal (several skin depths thick) has diamagnetic properties and can be used to shield the pipe wall from stray flux. Figure 5 shows two concepts of shielding: a single high conductivity cylinder surrounding all three conductors; and one coaxial shield around each conductor (the three shields must be electrically connected). In the case of the coaxial shield, the shielding currents are equal to the load currents. The shielding currents produce

\ \ \ \ M -- 20C \ M * 5 00 \

^

NUMBER OF STRANDS STRAND DIAMETER

vs. A-C RESISTANCE

M = RESISTANCE RATIO

^ ^

^ ^ ^

A-C RESISTANCE - I04 * OHMS/MILE

NUMBER OF STRAND LAYERS STRAND DIAMETER

vs A-C RESISTANCE

2 3 A-C RESISTANCE- I04x OHMS / MILE

N ^ M - 5 0 0

CONDUCTOR AREA vs.

A-C RESISTANCE

v 3-INCH BORE

N . M

^ X M * 200

•

* RESISTANCE RATIO 1

" ^ ~ ~

A-C RESISTANCE- 10* * OHMS/MILE

Fig. 4 — Strand Requirements Versus Resistance for Hollow-Bore Stranded Cable.

51

losses which are of the same order or less than the conductor losses and must be tolerated. The single shield is less lossy than three coaxial shields.

COOLANT CONDITIONS

In order to effectively remove the losses from a resistive cryogenic cable, the cryogenic fluid must be circulated and refrigerated. The flow rate depends on the allowable temperature rise in the fluid, which is bounded on the lower end by the freezin; point and on the upper end by a practical working pressure to suppress

TRIPLE SHIELDS

Fig. 5 — Shield Configurations.

boiling. Boiling is undesirable from the standpoint of dielectric properties as well as flow instabilities. For liquid hydrogen, the freezing point is 14°K and the vapor pressure at 30°K is eight atmospheres. These limits define a useful range. For liquid nitrogen, the corresponding limits are 63 °K and 100°K.Whether the full temperature rise is used depends on a number of factors in the design. The fluid must be removed from the cable, rerefrigerated at intervals, and then returned for recirculation.

Some cable configurations discussed later (see for example figure 9) consist of three hollow conductors in a single pipe, with a thick layer (e.g., one inch) of electrical insulation taped over the conductor. In such configurations, there is an intuitive tendency to use the three conductor bores as the " go " channel and the interstices

52

between the conductors as the " return " channel. This allows the system to be contained within a single pipe and promotes balance of the heat load between the " go " and "return" channels. However, depending upon the thermal conductivity of the elec­trical insulation, heat exchange will take place between the counter-flowing streams, thereby preventing the effective removal of the losses. Typical calculations of this effect for 10 mile or more lengths show that the thermal conductivity of the insulating layer must be nearly an order of magnitude lower than that of most nonmetals (and the fluids themselves) in order to achieve effective isolation of the two streams. These values approach the range of evacuated insulations. This factor adds some design complexity to the cable configuration.

REFRIGERATORS AND CRYOGENIC EQUIPMENT

The advantages of low resistivity or superconductivity in cryogenic cable are partially offset by the losses and costs associated with producing the low-temperature environment around the electrical conductors. The power required to produce refrigeration at low temperature is mostly determined by the fundamental thermo­dynamics of the refrigeration process. The minimum ratio of input power to refri­geration effect is governed by the heat rejection temperature and the desired refri­geration temperature. For the temperature levels most often considered for cryogenic cables, this results in the following ideal relationships for refrigeration when heat is rejected at 300°K. These ratios are not achievable in actual practice but are worth­while noting since they describe the absolute minimum in refrigeration penalty.

Refrigeration Refrigerator Input Power I Temperature Refrigeration Effect

77 °K 3/1 20 °K 14/1

4.2 °K 70/1

Three basic characteristics of refrigerators are important to the cryogenic cable system and, for that matter, to any application requiring low-temperature refrigeration. They are: — Ratio of refrigerator input power to the quantity of refrigeration; — Initial cost of the refrigerator; — Physical size of the refrigerator.

For power transmission application, reliability might be added to this list. The refrigerant used in the refrigerator may differ from the fluid employed as the

cable coolant. There is some freedom of choice of refrigerant, except at temperatures below the freezing point of hydrogen. If the coolant enters the cable in a subcooled liquid state, the refrigerant is not likely to be the same fluid as the coolant. Studies at the General Electric Research and Development Center indicate that neon is the desired refrigerant to provide subcooled liquid nitrogen, and helium is the refrigerant to provide subcooled liquid hydrogen. At liquid helium temperature, helium is the only possible refrigerant.

Refrigerators for cryogenic cable heat rejection consist of four basic components: — Turboexpanders; — Heat exchangers; — Compressors — Cooling towers.

53

A typical refrigeration cycle tor a nominal 20°K temperature is shown in figure 6. A cooling tower is required since heat rejection in most situations must be made to the atmosphere.

Ui±X U H *

100 F/410 PSIA -

100 F / 9 5 P S I A

93.4 K 4

9 4 ° F / 4 4 PSIA

HX

92*3 K / 4 7 PSIA

._J 92.3°K

27.6°K

3 0 ° K / 4 0 0 P S I A I 6 . 2 ° K / 5 0 P S I A

HX ■30°K

-+ LIQUID HYDROGEN |7°K

Fig. 6 — Helium Cycle Refrigerator for Cooling Liquid Hydrogen.

Table 4 shows a typical range of cable refrigeration requirements for three coolant temperature levels. These values are based on the estimated total heat load in the cable. In the General Electric studies, refrigeration processes have been selected for each of these conditions, and they result in the refrigerator characteristics described in table 4. The physical size of a refrigerator for a nominal 20°K temperature is shown in figure 7.

Temperature Level (°K)

77 20 4.2

Table 4 CRYOGENIC CABLE COOLANT REFRIGERATORS

Refrigeration Capacity

(kW)

2,000 300

10

Input Power/ Refrigeration

(kW/kW)

9.4 43.5

250

Refrigerator Plant Size

(Ft2)

8,000 5,400 2,400

54

The compressor and the drive motors are the major elements of the refrigerator costs in all of these systems. Therefore, approximate relative comparisons on refri­geration cost for cable systems using different coolants can be made on the basis of refrigeration input power. Cost of the refrigerator on a per watt load basis decreases significantly as the total load increases.

SYSTEM CONTROL AND STORAGE ROOM

■ UNDERGROUND CRYOGENIC CABLE - >

|MOT0R|«r 1 HEAT EXCHANGER

I COOLANT I VALVES

AND CONTROLS

4 feet _ i

Fig. 7 — Equipment Arrangement—300-kW Refrigerator for 20 °K.

M E C H A N I C A L DESIGN

The largest item of cryogenic equipment cost in a cryogenic cable system is the piping. The basic function of the cryogenic piping is to provide a thermal insulator to maintain the conductors at the design temperature. In addition, the piping may be required to act as a pressure vessel to contain the cryogen and to serve as the funda­mental structural support for the conductors, the electrical shields, and the insulation system.

Cryogenic piping has been in use for many years and is available from several commercial sources as a standard product. The pipe size and quantity required for cryogenic cable does, however, place new demands on this industry. The amount of

55

pipe required, for instance, for a twenty-mile line is almost an order of magnitude larger than the present yearly market for cryogenic piping.

The choice of design value for heat leak through the piping is a subject of consi­derable speculation. Very low values are possible, but they result in unreasonable costs and a requirement for difficult installation and maintenance. A practical design value for piping for liquid nitrogen or liquid hydrogen cable is about 250 microwatts per square centimeter. This results in a heat leak of five kilowatts per mile for a nominal inner pipe diameter of sixteen inches. Also, some allowance must be made for heat leak at field fabricated joints; this results in an additional heat load of about 3.5 kilowatts per mile.

Thermal contraction is an important factor in piping design for a cryogenic cable. Three basic schemes may be considered to allow for contraction: 1. Allow the inner pipe to freely contract and accommodate the contraction with

bellows in the outer pipe; 2. Use a low expansion coefficient material for the inner pipe and sustain a com-

pressive stress in the outer pipe; 3. Use contraction bellows in the inner pipe.

Scheme No. 1 requires that the outer pipe must contract with respect to its environment; this is an obvious disadvantage in an earth-buried installation. Scheme No. 2 requires relatively expensive and difficult-to-fabricate material fot the inner pipe and relatively large forces must be sustained. Scheme No. 3 appears to be the best solution, but it does result in some mechanical and electrical complications in the interior of the cable.

All of the cable system materials exposed to the cryogenic temperature will likewise be subjected to thermal contraction conditions. This includes the conductors and the electrical shields. Corrugated tubular conductors are a possible solution for this requirement, but they present manufacturing problems, increased material requirements, and an increase in cable resistance. The conductor configuration for a cryogenic cable may be of a flexible (reelable) form, like conventional cable, or in a much more rigid tubular configuration. Both methods have advantages and dis­advantages. The flexible cable may be pulled in long lengths through the pipe for installation, whereas the tubular conductor must be suitably joined at frequent intervals (approximately 40 to 60 feet).

Tubular conductors would probably have a short section of contraction bellows in every cable section. Flexible cable would require this provision at intervals of approximately 2,000 feet. Similar considerations apply to the electrical shields.

In the case of tubular conductors using cryogenic fluid and solid insulator spacers for the electrical insulation system, the spacers should be firmly attached to the conductors and the shields. The conductor and shield elements should therefore contract as a unit with respect to the inner cryogenic piping.

On the other hand, flexible conductors with an insulation wrap may be relatively independent of the piping, but they require solutions for the contraction problems associated with a built-up cross section of nonmetallic cylindrical former, electrical conductor, and organic insulation.

The "super-insulation" used to limit the heat leak into the cable requires a vacuum o f l 0 ~ 4 t o 10~5 torr. This vacuum is best produced under factory conditions which allow bake-out of the piping and sealing of the evacuated space at the factory.

The cost of the field installation of a cryogenic cable system will be a significant part of the total cost. Difficult installation situations, such as heavily traveled streets and highways or water crossings, may be easier with flexible cable configurations. In general, however, all configurations will probably suffer about equally from such conditions. The more important factors are the costs associated with joining cryogenic

56

piping, conductors, and shields. Maintaining quality in the field welds is always a difficult task and will be further emphasized because of the hermetic requirements and the need to produce good electrical conductivity.

Cryogenic cable concepts are quite easily put on paper, but those which are electrically practical and easy to install and have low loss are fewer in number. Auto­matic welding techniques are available for operations such as joining pipes and tubes, Significant effort must be devoted, however, to accommodating the use of such equipment. For instance, in a tubular conductor design employing bellows for thermal contraction, the physical arrangement of the components and the available space between the parts is a crucial factor in allowing the employment of these techniques.

The cost of assuring the integrity of the field joints may approach the cost of making the joint. X-ray techniques for inspecting pressure welds may be impractical because of interference from adjacent parts. Ultrasonic techniques may be more practical and spectographic methods will probably be required for leak checking. These factors of field installation cannot be ignored in design and cost studies. They must also be considered in determining the elapsed time for completing a cable instal­lation.

CONFIGURATIONS

Some results of applying the general ideas discussed above to specific cable configurations are illustrated by two basic configurations shown in figures 8 and 9. The configuration shown in figure 8 uses thin-walled tubes for both the conductor and the shield and free fluid as the dielectric. The second configuration, shown in figure 9, uses a stranded conductor on a hollow former, taped electrical insulation impregnated with fluid, and either a shield formed by taping the conductor on the outside of the insulation or a single large shield.

Fig. 8 — Three-phase Tubular Conductor Arrangement.

57

If the stranded conductor is used, it would be made in long lengths, reeled, and pulled trough the cryogenic piping essentially as is conventional, oil-paper, pipe-type cable. The hollow bore for the stranded conductor cable is used for direct cooling of the conductor. It can also be shown that a stranded conductor without the hollow bore would not be significantly smaller in diameter, for the same resistance, because of the strong influence of the surface flux density (inversely proportional to diameter) on the eddy current losses and the strand diameter.

OUTER CASING

SUPERINSULATION

INNER CASING AND

E.M. SHIELD

STRANDED CONDUCTORS

PERFORATED TUBULAR FORM(NONMETALLIC)

PROTECTIVE SHEATHING

SATURATED INSULATION

Fig. 9 — Three-phase Stranded Conductor Arrangement.

The tube version would presumably be welded together in the field; this precludes the use of taped insulation because of complexity of taping short lengths and making adequate splices in the electrical insulation in the field.

The configuration shown in figure 8 applies (with dimensional differences) to both liquid hydrogen and liquid nitrogen. The configuration in figure 9 has been sized for liquid hydrogen only. For purposes of comparison, all circuits are assumed to operate at 500 kV and have capacities of 3500 MVA, resulting in a phase current of 4040 am­peres. The cryogenic pipe is assumed filled with the appropriate fluid, all flowing in the same direction and, therefore, all capable of absorbing heat. A separate return pipe is assumed.

Various parameters applying to each of these configurations are given in table 5. It should be pointed out that many of the parameters chosen for the above comparison are somewhat arbitrary; other choices would lead to somewhat different values. For instance, the refrigerator spacing can be varied significantly with modest changes in the flow area. The voltage stresses represent the greatest unknown factor in sizing. Although the values assumed are about one-third of the breakdown stress measured in small samples, it may well turn out that this is not a sufficiently conservative assumption.

58

Table 5 CRYOGENIC CABLE PARAMETERS 500 kV, 3500 MVA, 4040 AMPERES

Coolant

Hydrogen Nitrogen Flexible Tubular Tubular

Conductor Bore Diameter, inches Strand Diameter, inches Conductor/Shield Wall Thickness, inches Conductor Resistance Ratio (to Copper at 300 °K) Maximum Electric Stress (volts per mil) Loss Tangent Shield Diameter, inches

Pipe Outside Diameter, inches Fluid Temperature Rise, °K Pressure Drop, Atmospheres Three-phase Losses, kW per mile

Conductor Shield Dielectric

*Heat Leak *Pump Power

Total Loss Refrigeration Multiplier Refrigeration Input, kW per mile Available Distance Between Refrigeration, miles

* Includes values for return pipe.

Other configurations are possible. P. Graneau has published a description of a liquid nitrogen system using vacuum as the dielectric medium [3]. He proposes a double circuit system, with one circuit used as the " go " nitrogen stream and the other for the "return ", with emergency capability of "go " and "return " in a single circuit. His conductor configuration is three solid tubes, suspended by spacers within a single eddy current shield. The conductor bores serve as the nitrogen flow passages.

Other configurations have also been studied in the course of the present investi­gation. These are summarized in figure 10. Space does not permit a detailed description of each, but a few comments may serve to illustrate the many conflicting factors which must be resolved.

Figure 10a, for example, shows a system with a return pipe inside the cryogenic pipe, thermally isolated by vacuum. Some saving in cryogenic piping cost may be realized in this way. However, the three shields must now contain the pressure of the cryogenic fluid. Since high-purity metal may not have sufficient strength, a com­posite tube of stainless steel lined with high-purity metal might be considered. Besides the inherent cost of manufacture, joints which contain expansion bellows between the lengths of composite tubing are particularly complex.

Figure \0b shows three flexible conductors in three separate pipes. This configu­ration has advantages in cable pulling and would confer some redundancy if a fourth conductor were to be added. However, it now appears that the piping cost for this configuration is significantly higher than for the three-in-one pipe configuration.

3 0.009

500 400

0.0001 17

(Single Shield)

21 16-28

7

11.5 5.8 3.7

13.7 3.6

38.3 42

1 610 16.5

3

0.030 500 250

6.6

26 16-28

7

29.0 13.6

20.9 9.3

72.8 42

3 060 53

3

0.24 8

250

6.9

28 65 - 85

7

202 104

23 16

345 10

3 450 17.5

59

Figure 10c is a rigid conductor version of figure 106, and figure lOd is a concentric configuration similar to a British superconductive design [4]. A separate fluid return circuit is implied for the configurations shown in figures 106, c, and d. The rigid conductor configurations apply both to hydrogen and nitrogen as coolants, with appropriate minor changes in dimensions. The flexible conductor has not been studied in detail for nitrogen, since the additional conductor metal required and the larger bore size for internal cooling make a practical, reelable configuration more difficult. Choice of the most promising configuration is not easy, since they all are complex. Study of many details of construction and cost are necessary for comparative evaluation.

VACUUM

OUTER CASING

SUPERINSULATION

INNER CASING

RETURN PIPE

-ELECTRICAL SHIELD

-TUBULAR CONDUCTOR

CRYOGENIC FLUID

CONDUCTORS

- P H A S E

- - PHASE

- P H A S E

A

B

C

CRYOGENIC FLUID

d. CONCENTRIC TUBULAR CONDUCTORS

Fig. 10 — Cable Configurations.

60

SUPERCONDUCTIVE CABLE CONFIGURATIONS

The so-called " hard " or " high-field " superconductors, such as niobium-titanium alloys or the compound niobium-tin, have reached a state of practical sophistication in applications for intens efleld d.c. magnetic coils. However, these materials are intrinsically lossy under a.c. conditions. Though the losses are small, the refrigeration penalty at 4.2 °K magnifies them to the point where they do not appear competitive with resistive conductors at 20 °K (see table 1). The full development of this thesis is lengthy and will not be attempted here; however, it is a generally accepted concept in the field.

The soft superconductors (such as lead and niobium) are not intrinsically lossy. Although they have low critical magnetic fields, the field strengths encountered in the cable application are low enough to permit the consideration of soft superconductors. Niobium appears to be the most likely candidate. In the soft superconductors, currents flow entirely in a surface layer only about 10~5 cm thick (in round numbers), and the critical current is related to the critical magnetic field through the classical laws of electromagnetism.

The configuration for resistive cable shown in figure 8 can easily be converted, conceptually, to one suitable for the application of superconductors. In this case, the coolant is liquid helium at 4.2°K and the conductor is niobium. The usual concept proposed is to coat, in some way, a thin layer of niobium on the high-purity metal tubes. The coating is on the outside of the conductor and on the inside of the shield. The high-purity metal tubes are retained to protect against current surges; if the critical current of the superconductor is exceeded, the current transfers temporarily to the high-purity metal substrate.

The cryogenic piping required for a liquid helium system is similar but more complicated than that required at higher temperature. The complication is introduced by the necessity for a cylindrical metal shield which is maintained at an intermediate temperature. A typical system might use a separate liquid nitrogen circuit to cool the shield. Most of the heat radiated through the superinsulation is trapped by the nitrogen shield and removed at a temperature where the refrigeration efficiency is good. The residual heat transfer to the helium is by radiation in vacuum and can be quite low (because of the T 4 dependence of radiant heat transfer).

Theoretically, the limit on the surface current in such a superconductive system is reached when the magnetic field at the surface (caused by the current) reaches the critical field of the superconductor. Niobium has complex superconductive properties, depending on its purity and state of cold work, and the apparent critical field varies markedly from sample to sample. Actually, an operating current limit well below that corresponding to the critical field, will be set by the level of a.c. losses in the niobium.

Although soft superconductors are intrinsically lossless at power frequencies, practical materials do have losses; these, again, depend on the purity and state of the cold work. A recently published British paper [4] quotes a.c. losses measured in commercial niobium and draws the conclusion that 375 amperes per centimeter (RMS) is a reasonable working limit for the surface current density (with a corresponding peak field of 660 oersteds).

Figure 11 [5] shows a summary of a.c. loss data taken on unalloyed niobium. These data support the British choice of 660 oersteds as the design field strength. Since the a.c. loss is proportional to the first power of frequency, it is here plotted in joules per square centimeter per cycle. The small cross at 660 oersteds corresponds to a surface loss of 10 microwatts per square centimeter at 60 hertz, which is shown later to be an acceptable level. The critical field of good niobium is 1300 oersteds at 4.2 °K and 920 at 6°K. (The latter is taken as the maximum allowable temperature

61

in the coolant.) Hence, some safety margin is built in by the above choice of the working field. The losses increase rapidly with increasing field; however, as will be shown, temporary over-currents could be handled in the heat capacity of the fluid.

10"

10"

UJ

o >-o e <> v c/)

UJ _ i =) o - 3

CO (/) °,

io-7

I0'8

IO"9

10 ■10

i i i I I I I I I i i i M i n i — T T T T T r

Nb

,m

IO"1

_...!_...L_LL1l.Iii _JL_L.1111Jil.. _...! L_l_Li_LLU IO1 IO2 IO3 IO4

PEAK SURFACE FIELD (Oe)

Fig. 11 — Reported a.c. Losses of Pure Niobium [5].

Using 375 amperes per centimeter (RMS) as the design current density, the three-inch tube pictured in figure 8 would have a current capacity of 7650 amperes. If one-third of the measured breakdown stress of liquid helium is used (150 volts per mil) the voltage capability is 300 kV.

Under these conditions, the three-phase line would have a capacity of 4000 MVA. Since the capacity of the hydrogen-cooled version is limited only by distance capa-

62

bility and refrigeration power, its capacity spans a similar range without exceeding reasonable limitations on either. Therefore, under the assumed voltage limit, there is not a great advantage in total capacity by using superconductors (with helium dielectric), although the current capability is somewhat better.

Assurance of low heat loads is of great concern in liquid helium systems because of the large refrigeration penalty and only about one or two degrees of temperature rise can be permitted, because of the reduced critical current nearer the critical temperature. The constraint on the temperature rise affects the helium flow required and, thus, the size of the line, through the flow area.

Assuming that the dielectric loss in liquid helium is negligible and that the a.c. superconductor losses are small, the dominant heat load to the liquid helium will be heat radiated from the 77°K nitrogen shield. This load can be calculated to be about 10 microwatts per square centimeter at an emissivity of 0.04, which is reasonably conservative. Achievement of this average level in practice depends on the kinds and number of spacers used and the method of joining factory assembled piping sections in the field. (Technology for helium piping of this quality does not now exist.) Although future practical experience may raise this number, for the present, it does not seem unreasonable to use it in comparisons.

As mentioned previously, the superconductor a.c. losses are of about the same order, and, as far as is known, the dielectric losses in free helium are negligible. Using these assumptions, the heat input to the liquid helium in the present example will be 265 watts per mile from thermal leak and 110 watts per mile from a.c. losses (at 10 microwatts per square centimeter); the latter figure seems reasonable but is somewhat tenuous depending on the metallurgical control of the niobium used.

If the liquid hydrogen-cooled version and the superconductive version are compared on the basis of refrigeration input per MVA transmitted, it is found that the liquid hydrogen values range from about 0.04 percent per mile at 2000 MVA to about 0.065 percent per mile at 5000 MVA, while the above helium value is 0.0047 percent at 4000 MVA. (Refrigeration ratios assumed were 40 and 500, respectively.)

The input power requirement is a good index to the size and cost of cryogenic refrigeration equipment. The superconductive cable appears to have a distinct advan­tage in the refrigeration requirement. However, this is substantially offset by a more complex piping system and the necessity for a composite conductor, the use of which intensifies the problems of contraction joints and field installation.

The British authors previously referred to have proposed an alternate configuration (fig. 12) for a superconductive cable design. The conductors are arranged concen­trically; each conductor contains an annular channel for the flow of liquid helium. The electrical insulation in this case is vacuum, the annular conductors being separated by spacers of suitable design. Since the conductors are separated by vacuum, "go " and "return " flow of the liquid helium can be maintained by using two conductors for the "go " stream and the third for the "return " stream without thermal contact between the streams. The concentric arrangement of the conductors produces zero net magnetic field outside and eliminates pipe losses. A liquid nitrogen shield is also shown schematically. The shield is split, thermally, so that "go" and "return " flow can also be maintained in the liquid nitrogen.

Recognizing that vacuum insulation is difficult to apply at high voltages, the authors propose that their system operate at 33 kV. The dimensions of their present design are such that the transmitted power is 750 MVA. The phase current for this power level is 13,000 amperes. The paper presents estimated costs for such a cable system. The authors conclude that it would be more costly than a conventional system at the proposed rating, but that the cost balance would swing in favor of superconductors at higher ratings.

It is worth noting that scaling up the rating at the same voltage would require

63

that the diameters of all of the conductors increase proportionally, and, at 3000 MVA, the overall diameter would be 66 inches if 33 kV is maintained. Hence, voltages over 100 kV are most certainly implied. This again raises the question of the successful development of high-voltage vacuum insulation.

HEAT SHIELD

THERMAL INSULATION LIQUID NITROGEN

DUCT

LIQUID HELIUM DUCTS

SUPER INSULATION

00025cm NIOBIUM CONDUCTORS BACKED WITH 0 25cm 99-999% ALUMINIUM

CORROSION-PROTECTED STEEL PIPE

10cm

SCALE

PHASES

yB Y, Y2 R

Fig. 12 — British Superconductive Cable Concept [4] (Reprinted by permission of the authors).

This brief comparison of superconductive systems with resistive cryogenic systems leads to no obvious conclusions with respect to their relative merits. The supercon­ductive systems have an edge in current capability but, with present knowledge, would have a corresponding deficiency in voltage capability. It seems that the outlook for a 500 kV superconductive system using liquid helium is less optimistic than that for the other two fluids.

64

The necessity for operating at low voltage creates a mismatch with the character­istics of future overhead systems; the importance of this mismatch has not been evaluated to date. There is strong reason to believe that superconductive systems will have substantially less refrigeration requirements; however, this advantage will be offset by a more complicated piping system, including a separate nitrogen circuit. The composite conductors required also present a problem. In the end, it appears that the final judgment again rests on details of construction and economics.

OVERLOAD CAPABILITY * | " i

A striking characteristic of the cryogenic cable system, which is not obvious at first glance, is the fact that it can readily withstand considerable overload. This comes about through the absorption of the extra power in the heat capacity of the cryogenic fluid. The configuration shown in figure 8 can be used as a typical example.

If the pipe is assumed completely filled with cryogenic fluid, the cable losses can be equated to the rate of the temperature rise of the fluid, assuming that no refrigeration is provided. In the liquid hydrogen case, the losses at a rating of 3500 MVA are 54 kilowatts per mile. The 21.5-inch pipe would contain 2.6 x 107 grams per mile of liquid hydrogen, whose specific heat at 20 °K is 9.4 joules per gram-degree K. The calculation shows that the rate of temperature rise of the fluid under these con­ditions is 1.25 hours per degree. In other words, the cable could carry a 40 percent overload (twice the losses) for 1.25 hours with a temperature increase of only one degree. A similar calculation with liquid nitrogen, assuming the same configuration is used, shows a temperature rise rate of 0.85 hour per degree.

Although liquid helium systems are commonly supposed to be sensitive to overload, the volumetric specific heat of helium is more than half that of liquid hydrogen. Figure 11 shows that near the critical field the a.c. losses are approximately 100 times that chosen at the design point in the example quoted. Under this load the temperature rise rate of the helium, if the pipe was completely filled, would be about four hours per degree. Also, even if the entire 7650-ampere load current is carried in the normal substrate, the temperature rise rate is 18 minutes per degree.

It should be emphasized that the above calculations were made for a case in which the entire cryogenic pipe is filled with the fluid. Configurations which contain less fluid will not yield such favorable results. Nevertheless, it appears that the outlook for overload capability is very good for all cryogenic systems.

COSTS

The eventual basis for the selection of a cryogenic cable system will be the lowest cost for the capacity needed. Figure 13 shows the range of cost and capacity for liquid hydrogen-cooled cable. These data were generated in a preliminary study several years ago [6]. The data generated in the present study support this general trend, although they have not been sufficiently verified for a detailed discussion at the present writing.

In the upper left-hand corner of the figure are average cost data for oil-paper, pipe-type cable installations [7], for both single and double circuit installations. Two curves are shown for each case, with the higher curve including the estimated cost of full reactive compensation.

Other data generated in the present study indicate that there is no large difference in cost between liquid hydrogen and liquid nitrogen installations. Although the present study has not examined superconductive systems in comparable detail, the

65

qualitative arguments presented in a previous section indicate no gross savings for superconductive a.c. systems. At the present state of knowledge, it can only be con­cluded that unit cost of transmission will be lower at high capacities, but more know­ledge and study are needed to separate the cryogenic systems from a cost standpoint and to place the present estimates on a firmer absolute basis.

$2.00

<: > LU Q_

tn en < _J _ i o Q

CO O O

230 h-345\

$1.00

230 I V345

\M K N N

^ — DOUBLE CIRCUIT

v500 kv -

R 500 kv

3</> CONVENTIONAL

3c/> LIQUID HYDROGEN SINGLE CIRCUIT

500 kv

1000 2000 3000 CAPACITY , MVA

4000 5000

Fig. 13 — Cost Comparison for Liquid Hydrogen-cooled Cable and Conventional Oil-filled Pipe-type Cable.

DIRECT-CURRENT SYSTEM

Direct-current systems have been frequently proposed using hard superconductors. Based on the previous discussion, it seems intuitively obvious that d.c. cryogenic systems will be much simpler and require much less refrigeration than a.c. systems. The present studies have not treated d.c. systems comprehensively, although it is easy to do so from the data now accumulated. Direct-current systems have been de-emphasized because the present concept of the utility of cryogenic cable is as an underground link in an otherwise a.c. system. The cost and space requirements for conversion equipment remain a significant negative influence. However, if a cryo­genic cable is a long, direct tie between a generator and a load center, or between two large systems, then a d.c. system may be competitive.

CONCLUSIONS

Cryogenic cable systems appear capable of handling large blocks of electric power with reasonable spacing between refrigeration stations. If the power capability is large enough, the cost in terms of dollars per MVA-mile is less than for conventional oil-paper cable.

66

The principal unknown technical factor is the obtainable dielectric strength and dielectric loss in long lengths at high voltage. Although small sample measurements are encouraging, these data do not extrapolate reliably, and larger scale testing will be required. The outlook is sufficiently attractive to warrant continuing study and research.

REFERENCES

[1] J. G. TRUMP, "Characteristics of High Voltage Vacuum-Insulated Systems," Proceedings of the Third International Symposium on Discharges and Electrical Insulation in Vacuum, Paris, France (Sept. 1968).

[2] J. T. S. LOOMS et ah, "Vacuum Insulation Between Very Cold Niobium Electrodes" British Journal of Applied Physics, Journal of Physics D, Ser. 2, 1 (1968).

[3] P. GRANEAU, " Economic Assessment of a Liquid Nitrogen Cooled Cable," Institute of Electrical and Electronics Engineers Paper No. 69, TP 95-PWR.

[4] E. C. ROGERS and D. R. EDWARDS, " Design for a 750 MVA Superconducting Power Cable," Electrical Review (Sept. 8, 1967).

[5] S. L. WIPF, " A-C Losses," Proceedings of the Brookhaven National Laboratory Summer Study on Superconducting Devices and Accelerators (1968).

[6] S. NEAL, " Cryogenic Transmission in the Power Industry of the Future," Cryogenic Engineering News, 3 (Aug. 1968), pp. 30-34.

[7] " Underground Power Transmission," Report to the Federal Power Commission by the Advisory Committee on Underground Transmission (April 1966).

DISCUSSION

A. SELLMAIER (West Germany) — What is the ratio between heat flux by supports and through the superinsulation ?

S.H. MINNICH — For liquid hydrogen and liquid nitrogen we have assumed superinsulated pipe of the kind now commonly sold in the U.S. It is evacuated and baked out in 40 foot lengths which are sealed at the factory. The end closures are re-entrant to form a long heat path. There are no other supports. The heat leak through the end closures is about 1/4 that through the superinsulation.

Obviously liquid helium piping constructed in this way would have a very high heat leak. I cannot quote a good number for well designed spacers. The numbers which I quote for helium heat leak neglect any spacer leak.

P. BURNIER (France) — Why did the author not consider beryllium as a cryoresistive conductor at LN2 level, as copper presents a resistivity ratio of 1/8, when beryllium has a ratio of 1/50 up to 1/100 under these conditions? The gain in resistivity might compensate for the higher price of the conductor.

S.H. MINNICH — Beryllium would be very attractive with a resistance ratio of this order at a reasonable price. We have not considered it because such metal does not now exist (except in a few laboratory samples).

We intend to do some cost calculations to see at what price such beryllium would be attractive. If the price seems attainable, some research may be started on this subject.

67

A TOTAL ENERGY TRANSFER SYSTEM FOR LIQUID NATURAL GAS (LNG) AND ELECTRIC POWER

A.V. PASTUHOV and F.E. RUCCIA Arthur D. Little, Inc., Cambridge (U.S.A.)

Un systeme a transfert d'energie total pour les gaz naturels liquefies (GNL) et le courant electrique

RESUME : On etudie aux points de vue technique et economique un systeme original consistant en transfert simultane par canalisations de gaz naturels liquefies (GNL) et de courant electrique; on examine divers facteurs tels que risolation thermique et electrique, ainsi que les methodes frigorifiques. Onpresente des conclusions preliminaires sur Veconomie possible d'un tel transfert d'energie integre et Von fait des previsions pour Vavenir le plus probable de ce systeme.

INTRODUCTION

Others have reported on the independent transmission of LNG by pipeline or the single transmission of electric power by either conventional conductors or supercon­ducting lines; this paper will attempt to present infomation about possible economi­cally attractive concepts whereby both sources of energy would be transported concurrently.

The growth in overseas shipment of LNG, from 150 x 109 cubic feet per day (150 MMCFD) in 1964 to almost 1 000 MMCFD projected for 1972, paralleling a similar spectacular increase in the generation of nuclear energy and the size of individual plants seems to warrant an evaluation of the concept of concurrent transmission.

Since LNG must be unloaded at a seaboard location, and because large nuclear power plants require large amounts of cooling water, in some instances at least, both energy sources can be transmitted along parallel paths from the same geographic locations. For example, Le Havre, France, already receives about 50 MMCFD for delivery 110 miles inland to Paris. It is conceivable that some day a nuclear power plant could be built on the Seine estuary because of overloaded water supplies with nuclear power plants located in Paris. Another example could be in the United States between Philadelphia and Bethlehem, Pennsylvania. LNG can be imported at Philadelphia and a nuclear power plant could be cooled by water from the Chesapeake Bay with both energy sources transmitted 43 miles inland to Bethlehem where they are required by the steel industry.

TECHNICAL DISCUSSION

There are several possible concurrent transmission means for LNG and electric power. Our evaluation will consider distances in the order of 50 or more miles only, rather than shorter distances.

Transmission of electric power at LNG temperature Conceptually, the simplest means would be to transmit the electric power at

LNG temperature since there is no need for sophisticated vacuum insulation and low temperature refrigeration.

If only the heat load due to resistive losses is evaluated, then, based on the above value of resistivity at LNG temperature, each 1 000 M VA (3 phase) of electric power generates 1 200 kW of heat per mile at LNG temperature as compared to 13 kW at liquid hydrogen temperature. A transmission line flowing the equivalent of 50 MMCF

71

of LNG per day has a heat of vaporization equivalent of about 5 200 kW which, for our case of 1 000 MVA, is equivalent to a little less than 4-1/2 miles before all the liquid is vaporized. A more practical approach would be to increase the cross-sectional

Resistivity of Pure Copper at Ambient Temperature and Cryogenic Temperatures

(°K) (Micro-ohm centimeters)

300 (ambient) 110 (LNG) 77 (LN2) 20 (LH2)

2 0.45 0.25 0.004

areas of the conductors for the LNG case from the area needed at liquid hydrogen temperature; for instance, it is conceivable to have 100 times more copper at LNG temperature than required at liquid hydrogen temperature to achieve a transmission line of equal electrical resistance; this approach would require about 1 000 HP of refrigeration every 10 miles in order to avoid LNG vapor formation. Consequently, technical considerations cannot alone select the best choice for electric power trans­mission at LNG temperature and an economic analysis trading refrigeration costs against high purity copper or aluminum costs must be made.

ELECTRICAL INSULATION

MULTILAYER INSULATION

3 PHASE - 3000mva 500 kv, 3500 amps

RESISTIVE CRYOGENIC CABLE

Fig. 1 — Concept of concurrent transmission of LNG and electric power.

Transmission of electrical power at liquid hydrogen temperature

The next temperature of interest is 20 °K (fig. 1) and [10, 11]. This temperature level can be achieved with liquid hydrogen as described by Minnich and Neal of General Electric, or by using gaseous helium as a refrigerant. The first concern with gaseous helium is whether or not its electrical breakdown is equivalent to liquid hydrogen. The system we considered operated with helium at 10 atm and a temperature

72

span of 20 to 30 °K;experimental data is lacking for these conditions, but extrapolation of experimental data [8] indicates that high pressure gaseous helium probably has an electrical breakdown value of the same order as that of LH2.

The second concern deals with the pressure drop; our finding indicates that for equal flow area and heat dissipation requirements, the liquid hydrogen system has a much lower pressure drop than pressurized helium. The low specific heat and low density of helium are responsible for this condition since the mass velocity is four times larger to dissipate the same resistive heat load (consequently, the pressure drop for helium is about 30 times higher than it is for liquid hydrogen.) Therefore, although the electrical breakdown of helium is satisfactory, the pressure drop consideration makes the use of helium unattractive.

We then evaluated the possibility of using vaporized LNG as an energy source for driving the refrigeration machinery. Using the General Electric model as an example, for 3 000 MVA the refrigeration power requirement would be 22 000 kW every 10 miles and assuming a combustion efficiency of 30 per cent (for 1 000 Btu/ft3 of gas) this corresponds to 6 MMCFD or for 50 miles to 30 MMCFD. For a total trans­mission rate of 10 times this amount, the power requirements for the refrigerator, therefore, would account for about 10 per cent of the total flow. The heat leak for an LNG transmission line of 300 MMCFD capacity with bulk insulation of 6 inches (such as polyurethane) would vaporize approximately 20 MMCFD of liquid (based on a line velocity of 10 ft/sec.)

The only LNG trades of the order used in this example are the ESSO-ENI trade between Brega, Lybia, and La Spezia, Italy, at 235 MMCFD and the Somalgaz-ERAP trade between Skikda, Algeria, and Marseilles, France, at 350 MMCFD. From the above example it can be concluded that, with the proper match between the LNG import size and the electric power generation, the concept of using vaporized LNG to drive the prime movers of the refrigeration system is feasible.

Transmission of electric power at superconducting conditions

Our next evaluation was to apply the concept to superconductive power trans­mission. Our analysis was based on the excellent work of Edwards, Swift, et al. [3, 12, 13, 15, 16]. The geometrical arrangement of the conductors, insulation and refrigeration streams for a non-optimized voltage is shown on figure 2. The arrangement has been scaled up from a 750 MVA system to a 3 000 MVA system for easier compar­ison with the General Electric cryogenic resistive concept. The major thermodynamic difference between non-superconducting and superconducting systems is the heat load level. When an optimum voltage of 155 kV is considered for the 3 000 MVA supercon­ducting system, the heat load to the helium refrigeration system is only 3.7 kW for 10 miles; assuming a ratio of 300 to 1 for the helium power consumption to refriger­ation, the total installed power would be 1100 kW. Again, if we consider vaporized LNG to power the compressor section of the helium refrigerator with the same previously assumed efficiency, the gas requirement is about 1/20 of that needed for resistive power transmission and must require the use of a better insulator like multi­layer insulation; for the case in point, a heat leak of about 150 to 200u, watts/cm2

will allow sufficient heat leak to generate the required gas. Therefore, from thermody­namic considerations, the general arrangement of figure 2 seems reasonable.

In figure 3, we have shown a typical arrangement of the equipment for a supercon­ducting power transmission line with concurrent LNG transmission; we have located all the heat exhangers in a single cold box and have installed a spare drive and com-

Note: The electrical industry defines 1 MVA as a million voltamperes while the gas industry defines 1 MMCF as a million cubic feet.

73

pressor assembly of such size as to carry the full load for any section, i.e., slightly in excess of 750 HP. We believe that the vacuum system should be a cryopumped system, thereby eliminating the need for diffusion pumps.

MULTILAYER INSULATION

ExgaJ LIQUID HELIUM RETURN (5°K)

E!nnJH3 LIQUID HELIUM GO (4°k)

3 PHASE-3,000 MVA 33kv - 52 500 amp (N0T OPTIMIZED)

Fig. 2 — Concept of concurrent transmission of LNG and supercon ducting electric power

-t--: VACUUM WALL BETWEEN

CRYO-PUMPED COMPARTMENTS'

- 1

1 I LI I I > . . .

HELIUM LOOP V

tf I I +

5T

ELECTRICAL CONDUCTORS

" REFRIGERATOR STATION

COMPRESSOR

ft ftftN ATURAL GAS PRIME MOVERS

VACUUM PUMP (FOR INITIAL J I PUMP DOWN)

POWER FROM j - ,

GENERATOR i

?5" 1- 5 k m - *- i POWER

i USE

AT "T5" S ^REFRIGERATOR STATIONS

Fig. 3 — Refrigeration equipment and prime mover arrangement for a superconductor power transmission system.

74

In figure 4, we have shown a schematic flow diagram of the refrigeration plant and its relation to the LNG transmission; of interest is the thermodynamic gain that can be achieved by using LNG as a heat sink for the helium refrigeration system. The Carnot work is improved by a factor of about 3, thereby reducing the horsepower requirement in the preceding example to about 370 kW.

Fig. 4 — Schematic flow diagram.

LNG pipeline

The design of the LNG transmission line has been based on a 9 per cent nickel line insulated with 6 inches of foam insulation; exansion bellows every 200 feet have been chosen rather than expansion loops which would be very difficult to accommo­date if parallel electrical power transmission is considered. In order to protect the LNG line from the soil environment (water), a 3 inch thick concrete conduit has been included. The following LNG transfer rates and corresponding line sizes were eva­luated:

MMCF/day

50 100 300 500

Transfer rate

MMCF/hour

2 4

12 20

GPM (1>

500 1 000 3 000 5 000

Line Size (in.)

6 8

12 14

Line sizes

Liquid Velocity ft./sec.

5.68 6.38 8.52

10.40

Water head loss in (ft.)/100 ft

3.36 2.98 3.17 3.85

(l)Gallons per minute Assuming an average head loss of 3 ft/100 ft., the total loss for an LNG line of 10 miles is approximately 20 atm.

75

ECONOMIC DISCUSSION

Conceptually, the simplest economic analysis which can be made is to assume that the installation costs for an LNG transmission line and an electric power transmission line are shared. The following cost estimates were made for various capacities of LNG and natural gas transmission (table 1).

The cost information for gas transmission lines were established from U. S. Federal Power Commission authorizations for 1967 [18]. This data was also used in establish­ing burial costs of various size lines from the known costs of the pipeline steel. An American Petroleum Institute 5LX42 steel was assumed for all line construction and consideration was given ASA Gas Transmission Code.

Table 1 COSTS IN $1 000 PER 10 MILES (16 km)*

Line Capacity (MMCFD)

A. LNG Line Size (in.)

50

6

100

8

180

10

300

12

500

14

Pipe Insulation Fittings/expansion joints Concrete conduit General construction

153 250 132 250 800

220 290 178 300

1 000

310 330 224 350

1 100

400 370 250 450

1 200

600 415 300 450

1 300

Totals 1 585 1 988 2 314 2 670 3 065

B. Natural gas line size (in.) 14 16 20 24 30

Total installed cost 480 550 780 1 100 1 630

(*) Line costs (exclusive of liquefaction costs).

In table 1 five items of cost are presented for establishing the installed costs of LNG transmission lines. The lines are sized on the basis of 300 psi pressure loops per 10 miles of length. Because the lowest temperature of the line is that corresponding to LNG, saturated at 1 atm ( — 259 °F), ASTM A353 (9 per cent nickel) steel was assumed as the most probable construction material. Schedule 10 was used for pipe wall thickness, and mill costs of $ 0.31 per pound were used for determining the line cost.

All lines were assumed to be covered with 6 inches of urethane foam insulation poured in place. Costs were computed on the basis of an average foam density of 3 pounds per cubic foot and on installed cost of $ 1 per pound. Current costs of the basic materials are from $0.40 to $0.50 per pound. In table 2 below are summarized the approximate heat leak per mile and the temperature rise of the liquid per 10 miles for the stated flow rates.

The line and foam insulation require some protective casing particularly in consid­eration of the fact that portions of the transmission lines might be installed in low­lands and swamp areas. A number of casings can provide the requisite protection. These include PVC plastic, steel and concrete. In the current analysis we assumed that the lines would be encased in concrete pipe, 3 inches wall thickness. The cost of this casing was based on $ 4.00 per cubic foot.

76

Table 2 L N G TEMPERATURE RISE RESULTING FROM LINE HEAT LEAK

Flow MMCFD

50 100 180 300 500

Line size (in.)

6 8

10 12 14

Heat leak (Btu/hr. mile)

122,000 143,000 164,000 185,000 200,000

Temperature rise per 10 mile (°F)

14.5 8.5 5.7 3.7 2.4

A steel line contracts approximately 3 inches per 100 feet of length when cooled from normal ambient to LNG temperatures. Corrugated transfer lines would appear to provide a solution to the contraction problem [1]; however, for this analysis we considered double expansion joints and anchor points at an approximate spacing of 200 feet. Preliminary cost information indicates that a 12-inch size expansion joint will cost about $ 1 000; other sizes, both larger and smaller, were estimated on the basis of $ 80 per inch of line size.

The final item of cost is that which includes rights of way, excavation, line burial, etc-These costs were determined frcm the costs associated with gas lines and were based on the LNG line diameter plus 18 inches for insulation and casing.

The previous technical discussion described three possible concepts of electrical power transmission; first, at LNG temperature, second, at 20°K, and third, at 4-5 °K. The economic considerations of each concept are discussed in the following paragraphs.

In the case of electrical power transmission at LNG temperature, our analysis indicates that it will always be less expensive to add copper than to add LNG refri­geration. If we take the GE model [11] at 3 000 MVA, we can make a gross estimate by eliminating hydrogen refrigeration and power charges and replacing these by an increase in the copper cost so to achieve the same heat load as at liquid hydrogen temperature. We assumed a cost of copper of $ 0.50 per pound which gave us a total materials cost of $ 1.90 per KVA mile, added to an installation cost of $0.33 per KVA mile, for a total of $ 2.23 per KVA mile. This total can be compared against $ 1.19 per KVA mile at liquid hydrogen temperature. Therefore, this method of transmission does not seem economically attractive in comparison to transmission in accordance with the GE model.

The next case for economic consideration is the liquid hydrogen case (GE model); in our technical discussion we concluded that using helium refrigeration at 20 °K was not as attractive as using liquid hydrogen as a primary refrigerant. The concurrent transmission of LNG may, however, offer some economic advantages otherwise not possible; for instance, the power charge (see note) for driving the hydrogen refrigerator is no longer required because the vaporized LNG would provide the energy require­ments. Another advantage would be to use LNG as a low temperature heat sink for the refrigerator system (however, in this example this factor was not evaluated). The following comparison shown in table 3 indicates the level of gross savings which might be achieved assuming no charge for burial of the LNG line.

Note: Power charge is defined as that incremental generator power required over and above the total transmitted power in order to operate the refrigeration system.

77

Table 3 COST IN MILLIONS ($) FOR 10 MILES

Line capacity (MMCFD) 50 100 500

LNG line materials cost .78 1.00 1.76 3 000 MVA line materials cost (without power charge) 11.80 11.80 11.80 3 000 MVA line installation cost 10.00 10.00 10.00

Total concurrent cost 22.58 22.80 23.56 Total independent costs* 37.38 37.79 38.86 Total independent costs for a

3 000 MVA line and a gas transmission line 36.28 36.35 37.43

13.70 13.55 13.87 Total cost savings (%) (38%) (37%) (37%)

* Includes a power charge of 250 $/kW—Figures based on the GE model estimates of 35.8 millions ($).

The last analysis we made was based on the superconducting line of the Norris and Swift, et al., 750 MVA Model; the liquid nitrogen was replaced with LNG (without change in geometry but with the normal LN2 return line flow reversed) allowing a flow in the order of 10 MMCFD for the same pressure drop. Only a relatively small increase in line size would allow a flow of 50 MMCFD.

Their estimated cost for a 3 000 MVA line optimized for voltage is S0.25/KVA mile, yet because British installation costs are small when compared to American costs, we have applied an additional $ 0.33/KVA mile for a total of $ 0.58/KVA mile.

The concurrent and independent installations of an LNG line of 50 MMCFD and a 3 000 MVA electrical power transmission have been estimated and the results presented in table 4.

Table 4 SUPERCONDUCTING TRANSMISSION LINE IN MILLION DOLLARS

1. Independent installation for 10 miles LNG Line (50 MMCFD) 1.58 3 000 MVA Electrical Power Line 17.40

Total Costs 18.98 2. Concurrent installation 17.4

In effect, the LNG line can be included at no additional cost to the system (we have assumed that the slight increase of line cost because of the 50 MMCFD flow and the increase in helium refrigeration requirements are approximately an equal total increase to the savings of using LNG as a heat sink for the helium system.

CONCLUSION

The two most economically attractive concepts for concurrent transmission of LNG and electric power appear to be the resistive model and the superconducting model; although the superconducting model seems to offer an unusual saving over the resistive model for equal conditions, that is 17.4 versus 22.6 million dollars. It must be recognized that the basic estimates, which the authors indeed are grateful for being

78

published, were prepared by different organizations on two different continents. Consequently, even though some adjustments were made, the authors urge caution to the potential users of the cost information.

Nevertheless, it appears that if the natural gas industry were to join the electric power industry, substantial savings could be made available to the ultimate customers.

REFERENCES

[1] K. ANDRESEN, F. DIAS and N.D. KENNEY, "Corrugated Metallic Cable Sheath. Paper 58 75 presented at the AIEE Winter General Meeting, New York (Feb. 2, 1958) 9 p.

[2] M. DIMENTBERG, "Better Economics Promise Eventual Use of LNG Lines." The Oil and Gas Journal (Sept. 18, 1967), pp. 96-102.

[3] D.R. EDWARDS and R.J. SLAUGHTER, "Superconducting Power Cables." Electrical Times (Aug. 3, 1967), 4 p.

[4] Pierre HERNE, "LNG Transportation by Pipeline: Possible Solutions, Future Prospec-tives." Paper 29B presented at First International LNG Conference, Chicago, Illinois (Apr. 7, 1968), 6 p.

[5] R.H. KROPSCHOT, B.W. BIRMINGHAM and D.B. MANN, Editors, "Technology of Liquid Helium." National Bureau of Standards Monograph III, (Oct. 1968), U.S. Department of Commerce, Washington.

[6] G. LAING, "Pre-insulated Pipelines." Natural Gas and L.P.G. (Aug. 1968), pp. 31-33. [7] J.S.T. LOOMS, R.J. MEATS and D.A. SWIFT, "Vacuum Insulation between Very Cold

Niobium Electrodes." Brit. J. Appl. Phys., 1, (1968), Ser. 2, pp. 377-379. [8] N. MATHES, "Cryogenic Cable Dielectrics." Paper presented at the 1968 NEMA Con­

ference in Los Angeles (Dec. 11, 1968), 12 p., 8 fig. [91 N. MATHES, " Dielectric Properties of Cryogenic Liquids." IEEE Transactions on Electrical

Insulation, E 1-2, No. 1 (Apr. 1967). [10] S.H. MINNICH, "Technical Aspects of Cryogenic Cable Design." Paper presented at the

1968 NEMA Conference in Los Angeles, (Dec. 11, 1968), 12 p., 4 fig. [11] S. NEAL, "Cryogenic Transmission of Large Blocks of Power." Cryogenic Engineering

News (Aug. 1968), pp. 30-34. [12] W.T. NORRIS and D.A. SWIFT, "Development Augur Design of Superconducting

Cables." Electrical World (Jul 24, 1967), pp. 50-53. [13] E.C. ROGERS and D.R. EDWARDS, "Design For a 750 MVA Superconducting Power

Cable." Electrical Review (Sept. 8, 1967). 4 p. [14] B. SCOTT, "Cryogenic Engineering." Prepared for the AEC, D. Van Nostrand Company,

Inc., Princeton (March 1959). [15] D.A. SWIFT, "Optimum Flow Conditions for Liquid Coolants in Superconducting

Power Cables." Cryogenics (Aug. 1968), pp. 238-243. [16] D.A. SWIFT, "Prospects For The Superconducting A.C. Power Cable." I.I.R. Xllth

International Congress of Refrigeration, Madrid (Sept. 1967), I, pp. 173-185. [17] "Pipeline for Gas and Electricity." Petroleum Times. (Aug. 2 , 1968). p. 142 [18] Natural Gas Construction Data. Gas Appliance Manufacturers Association, Inc.

(June 1968).

DISCUSSION

D.A. SWIFT (U.K.) — Have the authors considered the danger aspects of having inflammable gas and electrical plant in close proximity ? Would such a practice be permitted ?

A. V. PASTUHOV — In the event of a leak from the LNG system and the occurrence of a spark in the system a dangerous condition will not necessarily result; for instance, in the cases of either a superconducting assembly or a liquid hydrogen cooled assembly thermal insulation is by vacuum between the LNG and the colder electrical parts and therefore an oxidizer is not present to initiate an explosion or fire.

The General Electric model is based on liquid hydrogen and its application will require the same safety considerations as a system with LNG concurrently transmitted with the electrical power.

The greatest concerns should be 1) the amount of electrical stored energy in a 3,000-4,000 MVA in the event of a short and/or 2) a sudden vacuum failure.

79

BASIC PROBLEMS OF D C . POWER TRANSMISSION BY CRYOGENIC CABLES

H. VOIGT AEG- Telefunken, FrankfurtI Main ( West Germany) .

Problemes de base de la distribution de courant continu au moyen de lignes cryogeniques

RESUME : Au cours d'une analyse des problemes techniques et economiques de la distribu­tion souterraine de courant continu on a considere trois types de liaisons : (1) un type classique a conducteur de cuivre aux temperatures nor males de service; (2) une liaison basse temperature a conducteur a"aluminium refroidie par hydrogene liquide; (3) une liaison supraconductrice.

Vanalyse s'appuie sur des liaisons a un seul conducteur isolees pour une tension de reseau de 110 a 440 kV. Un traitement theorique indique qu'une cryoliaison de distribution a deux conducteurs d'aluminium fonctionnerait avec des pertes (y compris Venergie de refrigeration) plus basses que les pertes ohmiques d'un raccordement de courant a deux conducteurs classiques de diametre exterieur equivalent. La puissance transmise serait a peu pres la meme dans les deux cas. Deux cables supraconducteurs (toujours du meme diametre) permettraient une augmentation de 5 fois de la puissance transmise mais leurs pertes globales seraient ljfois plus grandes que celles de deux lignes classiques equivalentes. Des calculs s'etendant sur plusieurs Gigawatts de puissance transmise indiquent par exemple qu'une ligne a courant continu pour 5,5 GW/440 kV exigerait six paires de fils classiques en parallele ou cinq paires de fils aluminium a basse tem­perature en parallele ou bien une paire de fils supraconducteurs. Les rapports de pertes globales seraient de 2J (cuivre) : 1,7 (aluminium) : 1,0 (supraconducteur). Le calcul du cout d'instal­lation et d'entretien annuel d'une ligne de distribution de courant continu avec les differentes matieres mine aux resultats suivants : (a) les lignes a"aluminium refroidies par hydrogene liquide ri* off rent pas d'avantage economique important par rapport aux lignes classiques; (b) les fils supraconducteurs promettent une baisse des prix s'il s'agit de la distribution de 4 GW ou davantage. Le seuil de rentabilite pose la question generale de savoir si la distribution de courant continu sur le plan de plusieurs G W par le moyen de liaisons cryogeniques sera la meilleure methode de distribuer Venergie electrique aux regions tres peuplees a Vavenir.

Over several years now, we have become acquainted with the question being raised in publications and at conventions whether cables with low-temperature resistive conductors or with superconductors for the transmission of d.c. power in the multi-gigawatt range would be more advantageous than cables of conventional design [1, 2, 3, 4]. One of the most noteworthy projects has been published by Garwin and Matisoo [5] in which they described a superconductive d.c. line capable of trans­mitting 100 GW at 200 kV system voltage. In the near future however, there appears to be no demand arising for a cable line in that range of transmitted power, particu­larly not in Europe.

Generally it can be said that the considerable expenditure on refrigeration would render a cryogenic cable non-competitive with an equivalent pressurized paper-oil cable below a certain limit of transmitted power. In several studies (e.g. survey in [6]), conventional flexible power cables were compared with an underground cryogenic transmission line which is supposed to be built with pipeline techniques. This report presents some theoretical investigations on single-conductor d.c. cable models under the premise of comparable design, in order to find out at which amount of transmitted power the cost of a cryogenic alternative will break even with the cost of conventional cables.

What speaks for the application of cryogenic power cables is the reduction of conductor resistivity by several orders of magnitude at very low temperatures. But the factor by which the power density in a cryogenic design can be raised in comparison

83

to a conventional cable is much less than could be expected from the respective ratio of resistivity. If cables of equal outer diameter are considered, the conductor cross section which can be used in a cryogenic cable is relatively small, since a substantial amount of the total cable cross section must be sacrificed to a refrigerating duct and the thermal insulation. For a comparison, the effective power density might be expressed in respect to the overall cable diameter D, i.e.

P = 4 U 0 I / T C D 2 , (1)

where Uo denotes the voltage conductor-to-ground and I the rated current.

a) COPPER (b) ALUMINIUM (c) SUPERCONDUCTOR Tm = 343 °K Tm = 23 °K Tm = 5.5 °K

LIQUID HYDROGEN HELIUM

^ \ \ ELECTRICAL INSULATION ((((((((THERMAL INSULATION

Fig. 1 — Single-conductor cable models.

The technology aspects of cryogenic cables can be assessed quantitatively when they are considered in relation to present-day power cable techniques. So, three simplified models of single-conductor d.c. cables shall be compared with each other as represented in figure 1: on the left, a conventional design with Cu conductor; in the middle, a cable with a conductor of high-purity Al , cooled by liquid hydrogen; on the right, a superconductive cable with helium-cooled Nb/Ti wires. The mean operating temperatures are indicated to be 70°C, 23 °K and 5.5 °K respectively. The electrical insulation and protective sheath are supposed to be similarly designed for all cables. In the low-temperature Al and in the superconductive model, a thermal insulation (superinsulation, evacuated) is assumed to separate the conductor and its electrical insulation. The mean heatflux radially penetrating the thermal insulation is supposed to be 1.2 W/m2 in the aluminium, 2 W/m2 in the s.c. cable. Further data which are used for the theoretical treatment are listed in table 1.

For the cryogenic models the additional outer diameter of the thermal insulation corresponds to the conductor diameter in the conventional cable (6 cm). Hence, the overall cable diameter, too, would be the same for each cable at a given voltage.

Along a low-temperature transmission line a series of refrigerating stations (including the vacuum pumps for the thermal insulation) must be installed. One such station would cover a distance of around 4 km for the two cryogenic models (table 1).

Besides the transmitted power which can be handled by a cable line, the losses per unit length must be taken into account. For a conventional line of the length / and the ohmic resistance R, the specific loss is just N = PR// . The heat generated in the

84

Table 1 DATA OF THREE SINGLE-CONDUCTOR CABLE MODELS

CONDUCTOR

MEAN TEMPERATURE °K

CRYOGENIC FLUID

O.D.OF CONDUCTOR CM

O.D.OF THERMAL nM INSULATION O M

DIAMETER OF r M COOLING DUCT U M

THERMODYN. STATE OF CRYOGENIC FLUID

ENTERING LEAVING

MEAN VELOCITY OF M / c CRYOGENIC FLUID M / : >

DISTANCE BETWEEN ™ REFRIG. STATIONS rvi¥l

CU

343

-

6.0

-

-

-

-

-

AL

23

LIQUID HYDROGEN

3.0

6.0

1.7

20°K /14 ATM 2 6 ° K / 4 ATM

3.0

4.0

S.C.

5.5

SUPERCRIT. HELIUM

3.0

6.0

2.8

4.5°K/8ATM 6.5°K/2ATM

1-5

4.1

conductor is to be removed externally, i.e. by heat transfer to the soil or to an air or water cooling system in the cable trench. For a cryogenic line, again of the length /, the losses can be expressed by the equation

^ = (Qe+Qk+Qt)rn + nkcIll+Qt. (2)

Qe is the external heat influx per km through the thermal insulation. Qh designates the hydraulic losses in the cooling fluid per km. Q* is the ohmic loss per km. The factor m denotes the refrigerating ratio (watts input/watts load). For large hydrogen liquefying plants, m = 40 can be taken as a typical value. Helium refrigerating plants are expected to work with m = 400 for several kilowatts of refrigerating load at approximately 5 °K. The letter n stands for the number of cable terminals, e.g. n = 4 for one pair of single-conductor cables. kc designates the specific refrigerator input power required to cool one current lead. If the lead be optimized with regard to the electrical and thermal conductivity, kc can be brought down to 1 to 10 watts per ampere or even less. Of course, nkcl/l is the smaller the longer the line, and can be neglected in equation (2) for / > 100 km.

Equation (2) indicates that the bulk of the losses is represented by the power required to drive the refrigerating plant. This is especially significant for a super­conductive d.c. line which has no ohmic losses ( Q * = 0), but needs a relatively high ra-factor for the conductor to be kept at a few Kelvin-degrees only.

Figure 2 shows the transmitted power P 2 and the specific losses N 2 versus the system voltage, calculated for the three types of cables. Index 2 refers to a line with two single-conductor cables, operated in a symmetric d.c. system with an earthed mid-point. The diagram for P 2 reveals that one pair of low-temperature Al cables could transmit only about the same power as one pair of conventional Cu cables

85

respectively at 110 to 440 kV system voltage. On the other hand, one pair of s.c. cables would allow an increase in the transmitted power by a factor of 5.1 to 5.5.

The specific losses for Cu cables decrease slightly at higher voltages since the rated current must be reduced with an increased thickness of the electrical insulation.

MW J

8 _

6

4

1000 8 " 6~

4 _

2 -

100 H

fp2

•

y

V

1

^

- \

^*»

1

•""**

'

1

U

100 200 300 400 500 KV

KW/KMJ

OUU

200-

l U U

I

' N 2

U | i | f 0 100 200 300 400

U 500 KV

2 SINGLE-CONDUCTOR CABLES: CU L.T.AL S.C.

Fig. 2 — Transmitted D.C. power and total losses per km.

The losses of the Al and s.c. cables are not dependent on the voltage because of their internal refrigeration. In the voltage range given in the diagram, the Al cable losses turn out to be 1.7 to 1.9 times smaller, the s.c. cable losses 1.5 to 1.7 times higher than the Cu cable losses.

Table 2 HlGH-VOLTAGE D.C. TRANSMISSION LINES

CONDUCTOR

SYSTEM VOLTAGE KV

1 NUMBER OFSINGLE-1 CONDUCTOR CABLES

OUTER DIAMETER _Ril OF EACH CABLE C M

TRANSMITTED POWER GW

LOSSES KW/KM

EFFECTIVE M , l f / ~ . . o POWER DENSITY M W / C M 2

CU/343°K

440

12

10.4

5.6

516

5.5

880

12

12.3

10.3

444

7.2

AL/23°K

440

10

10.4

5.5

340

6.5

880

10

12.3

11.0

340

9.2

S.C./5.5°K

440

2

10.4

5.5

200

32.4

880

2

12.3

11.0

200

46.2

86

According to figure 2 however, one pair of s.c. cables can be made to convey 5.5 GW of d.c. power at 440 kV. If the same power is to be handled by low-temperature Al or conventional Cu cables, a certain number of such cables must be operated in parallel. The result of the respective calculations is shown in table 2.

With regard to mutual thermal influence, for the Cu cable transmission line the assumption is made that each pair of cables may be loaded with a current correspond­ing to 75 percent of that loss power which could be admitted without additional loaded cables laid close nearby. The Al and s.c. cables being cooled internally do not suffer under this restriction.

For transmitting 5.5 GW at 440 kV, either six pairs of Cu or five pairs of Al or one pair of s.c. cables would be necessary. Table 2 furthermore gives the capacity of each line at 880 kV system voltage: 10.3 GWfor the conventional technique, 11 GW for the cryogenic lines. In this case the outer diameter of each single-conductor cable would be 12.3 cm against 10.4 cm at 440 kV. The listed values of the losses indicate that the low-temperature Al and the s.c. line promise lower operational cost than the Cu cables. If the power density is calculated from equation (1), one gets for the Al cables approximately 1.2 times the value and fcr the s.c. cables 6 times the value found for the Cu cables. This involves the prospect of reducing the cable trench cost when cryogenic conductors are applied. However, savings in the cable trench and operational cost might be cancelled out by the cost for the cryogenic requirements. Therefore the total expenditure for each type of cable lines is to be estimated now.

At first, the installation cost are to be considered. Since comparison is made for the cables only, the expenditure for the inverter stations at both ends of a d.c. trans­mission line are disregarded.

The installation cost for a Cu cable line result from the partial cost for the con­ductor, electrical insulation and the requirements on the site (trench etc.)—expenses which are known within a certain margin from conventional techniques. For the low-temperature Al cables, the cost for refrigeration and thermal insulation is assumed to be 0.5 x 106 DM/km for one, 1.6 x 106 DM/km for five pairs of Al cables. These figures reflect the drop in plant cost with increasing refrigerating power. As for the

DM/KM

3-106

2-10

} AVERAGE INSTALLATION COST DM/KMf AVERAGE INSTALLATION COST

3 -10 6 -

6 GW

2-10

1-106

L.T.AL

S.C.

- > 1 r 2

TRANSMITTED POWER - i — | — i — | — i — | — i — | — ►

4 6 8 10 12 GW

880 KV

Fig. 3 — Average installation costs versus transmitted power.

87

conductor cost in the s.c. cables, the present price of stabilized Nb/Ti wires is used for calculation. (The quantity of single wires depends on the total current to be carried, i.e. on the transmitted power; the design of the model conductor is based on the condition of withstanding an overload current = twice rated current without quench­ing). The specific cryogenic expenditure is assumed to be 106 DM per kW of refriger­ation load for the helium plant and 5.5 x 105 DM/km for thermal insulation.

The result of the cost comparison is plotted versus the transmitted power in figure 3, on the left for the 440 kV, on the right for the 880 kV system. The curves for the Cu and the Al transmission lines are marked by a number of points which apply to 1 to 6 (or 1 to 5 respectively) pairs of cables in parallel. The Al line turns out to cause higher installation cost than the conventional line. The cost of the s.c. line is chiefly determined by the expenditure for the cryogenic installations; the slight increase with the transmitted power is solely caused by the superconductive material. The differing trends of cost for the Cu and the s.c. transmission line yield a cross-over at approximately 5 GW (440 kV) and 7.5 GW (880 kV). If one were to assume this estimate to lie within a margin of ± 10 percent, the break-even points would come to lie at 4.2 GW (440 kV) and 6.3 GW (880 kV) for the s.c. transmission line under the most favorable circumstances. With a similar allowance for errors, the Al line would still remain more expensive than the conventional line.

The three lines shall furthermore be compared in regard to their annual cost, again in dependence on the bulk power transmitted. The losses are estimated on the basis of 0.05 DM/kWh as well as 0.10 DM/kWh. Ten percent of the installation cost are credited to capital services and amortization. The hours of operation per year are taken as equivalent to eleven months of operation, during which period the refrigeration system would be required to maintain the conductor temperature. In order to make a fair comparison possible, operation at full load for 8000 hours per year is supposed for the cryogenic and for the conventional alternative.

The annual cost, calculated under the premises stated before, is shown in figure 4, both for 440 and 880 kV. In comparison to the conventional line, the low-temperature Al cables would effect several percent savings in the annual cost if O.lODM/kWh were taken as a cost basis. However, this is reversed for 0.05 DM/kWh. The s.c. line could not compete with the Cu line at too small a transmitted power, but for about

DM/KM

2 # 1 0 5 -

u ~~|

A V E R A G E

I

A N N U A L (

^

TRANSiy I

:OST

IITTED PO I

l_ U T.AL

J

.C.

NER

*

DM/KM + AVERAGE ANNUAL COST

6-10 5 -

4*10

2-10

0 1 2 3 4 5 6 GW

TRANSMITTED POWER ' I ' I ►

0 2 4 6 8 10 12 GW

440 KV 0.05 DM/KWH

880 KV

- - - 0.10 DM/KWH

Fig. 4 — Average annual cost versus transmitted power.

88

3 GW (440 kV) or 6 GW (880 kV) the break-even point in the annual cost would be reached.

Before the manufacture and installation of the first cryogenic high power cable can begin, extensive research and development activities are necessary. Quite a few problems are to be solved concerning for instance the thermal insulation. For the cable models studied in this report, 1 to 2 W/m2 have been used for the heat influx through the insulation. Values of this order and less can be guaranteed for special cryogenic applications on a laboratory scale. There still remains to find out however, what measures must be taken for maintaining a near-perfect heat insulation in the body of a cable which is to be built, handled and laid underground. Short-comings in heat insulating properties may force us to give up single-conductor cables in favor of designs enveloping two or more electrically insulated conductors in one common thermal insulation (possibly including radiation shields kept at intermediate temper­ature). But the bulk of such a design might preclude the conventional practices of transport and laying, so that a kind of pipe-line engineering must be resorted to. Cryogenic cable terminals (pot heads) also need a thorough developmental work since electrical insulation, refrigeration and current-carrying capacity must be closely matched. Finally some further efforts are called for in upgrading the availability of refrigeration and vacuum plant and components in order to satisfy the present-day standards of reliability for power generation and transmission plant.

The very great expenditure still necessary to develop cryogenic d.c. cables will be justified if electrical power in the multi-gigawatt range must be transmitted by cables because of operating contingencies which may arise in energy distributing systems of the future. Under these aspects, it is very difficult to give a safe forecast of the economic outlook for cryogenic cables. In this report, the costs of the super­conductive material and the refrigerating system components have been estimated according to present knowledge. If and when worthwhile applications for cryogenic cables should arise in the utility industry, these costs will be sure to come down. Then the economic break-even point of a s.c. line may be found for 2 or 3 GW of transmitted power. The trend to build ever greater blocks of generating capacity which is discernible particularly in nuclear stations, raises the problem of where to site these super power stations. Water-cooled power stations for instance cannot always be located close to the centre of energy consumption. From district to district an exchange of load must take place over high-power transmission lines of moderate length. Where considerations of appearance and right-of-way limitations speak against overhead lines in densely populated areas, underground transmission will be the logical alternative, eventually calling for the large-scale application of cryogenic cable systems, too.

REFERENCES

[1] P. DENZEL, Zukunftige Moglichkeiten der Uebertragung elektrischer Energie, ElektrU zitatswirtschaft, 67 (1968) 1, 1-5.

[2] S. NEAL, Cryogenic transmission in the power industry of the future, American Power Conference, 30th annual meeting, Chicago (Apr. 25, 1968).

[3] P. A. KLAUDY, Supraleitende Kabel, ETZ-A, 89 (1968) 14, 325-330. [4] E. MASSAR, Aspekte und Anwendung tiefer Temperaturen in der Elektrotechnik, ETZ-A, 89

(1968) 14, 335-339; [5] R. L. GARWIN and J. M ATISOO, Superconducting lines for the transmission of large amounts

of electrical power, Proc. IEEE, 55 (1967), 538-548. [6] Research on superconducting cables, The Engineer (Feb. 3, 1967), 196.

89

DISCUSSION

H. LONDON (U.K.) — Are your calculations for the superconducting cable based on type I or type II superconductors ? If they refer to type II, you have the problem of heating when there are changes in the load. It seems to me that the appro­priate way of operating a type II superconducting transmission line is to run it at constant current and to vary the voltage according to the load. Not only would the heating be avoided, but the power to be handled by the control gear would be less, because with the high magnetic fields available the magnetic energy stored in the cable is much larger than the energy of the electric field.

Conversion of variable voltage d.c. into constant voltage a.c. is certainly feasible, for instance by means of a motorgenerator set, but I should like to know how practi­cable it is.

H. VOIGT — The calculations are based on type II superconductors. Additional heating caused by changes in the load is a serious problem indeed. It has been consid­ered quantitatively by Garwin and Matisoo in their paper mentioned before. Generally, the operation of a superconducting d.c. line at constant current and variable voltage would be an expedient. But it still has to be found out whether this will be compatible with future utility operational practices.

90

TRANSPORT D'fiNERGIE ELECTRIQUE PAR CABLE SUPRACONDUCTEUR A COURANT CONTINU

G. DELILE Direction des Etudes et Recherches, Electricite de France, Clamart {France)

Electrical energy transmission via a direct current superconductor cable

SUMMARY: The report first deals with the general problem of the optimum dimensioning of a direct current superconducting electrical link as a function of two variable parameters: the capacity of the link and its length. In relation to this aspect, the author gives all the technical, technological and economical criteria adopted to determine its structure, the dimensions and the electrical properties of the cryo-links under study, and a justification of the selections made. Particularly studied is the dimensional incidence of overload currents defects in the cable.

The second part of the report sets out the main inherent difficulties in the industrial use of cryo-cables and determines the usage and the range of power requirements which may be advan­tageously applied.

1 — O B J E T DE L 'ETUDE

A partir des resultats obtenus dans l'etude de la refrigeration d'une liaison supra-conductrice [1], on etudie le probleme general du dimensionnement optimal d'un cable supraconducteur a courant continu en fonction de sa puissance nominale et en tenant compte des regimes de defaut auxquels il peut etre soumis. On suppose pour cela qu'une cryoliaison supraconductrice a courant continu serait utilisee comme le sont les liaisons a courant continu actuelles, c'est-a-dire pour relier deux points d'un meme reseau alternatif ou deux reseaux alternatifs differents. On determine une tension optimale de fonctionnement, et Ton evalue l'interet economique de telles installations.

2 — STRUCTURE ADOPTEE POUR LA CRYOLIAISON - ISOLATION ELECTRIQUE ET THERMIQUE

DES CRYOCABLES

Diverses geometries et dispositions peuvent etre envisagees pour les conducteurs d'une cryoliaison. On a adopte ici une structure coaxiale parce que cela semble conduire a une mise en oeuvre plus simple du cable. On a egalement suppose que chaque conducteur etait place dans une enceinte cryogenique separee ce qui permet de rendre negligeables les efforts electromagnetiques.

La liaison etudiee est done constitute par deux cables unipolaires paralleles refroidis par des refrigerateurs d'helium et d'azote regulierement repartis tout au long de son parcours comme il est indique sur la figure 2.

Chaque conducteur est forme par un tube cylindrique en aluminium tres pur recouvert d'un mince couche de supraconducteur (Nb3Sn) sur sa face externe. L'aluminium sert a la fois de materiau stabilisant et de support mecanique pour le supraconducteur. Le conducteur est maintenu a basse temperature par une circulation d'helium liquide a l'interieur du canal que constitue le tube.

Le conducteur a la temperature de 1'helium liquide est separe de 1'azote circulant sous pression par une enceinte sous vide destinee a assurer a la fois l'isolement thermique et l'isolement electrique du cable. Dans l'etude qui suit on a admis des pertes parietales par unite de surface au niveau de 1'helium de 0,1 W/m2 pour tenir compte des pertes par rayonnement et des pertes par conduction au travers de cales isolantes supportant le conducteur et de surface de contact aussi faible que possible.

91

La rigidite dielectrique du vide est certainement tres grande, mais peut etre reduite par l'etat de surface des conducteurs en presence, par leur distance, par la nature et la pression de gaz residuels dus au degazage des parois. L'influence de la temperature sur la tension disruptive du vide n'a encore ete que peu etudiee. Pour tenir compte de ses differents facteurs ainsi que de la presence indispensable de cales isolantes, on a admis comme valeur pour la tension disruptive du vide 150 kV/cm.

La distance entre les parois de l'enceinte sous vide est determined par la tension d'essai dielectrique entre phase et terre adoptee pour la liaison et qui a ete prise egale au double de la tension de service. On a done avec les notations indiquees sur la figure 1 :

^ L o g ^ = — - 1 0 - 6 m . (1) dx 15

ou U est la tension de service de la liaison en volts.

Helium liquide

Conducteur : couche supraconductrice

sur tube d'aluminium tres pur

Espace sous vide pour isolement

electrique et thermique

Tube en acier inoxydable (epaisseur 2mm)

Azote liquide

Tube en acier inoxydable (epaisseur 2mm)

Superisolat ion (epaisseur 2cm)

Tube en acier (epaisseur 5mm)

_ Les lettres "d" indiquent des diametres

Fig. 1 — Structure d'un cable supraconducteur

L'enveloppe cryogenique exterieure est constitute par une superisolation de 2 cm d'epaisseur qui isole l'ecran fluide d'azote liquide du milieu a la temperature ambiante. Les pertes parietales par unite de surface au niveau de l'azote qui sont dues a un phenomene combine de radiation et de conduction au travers du superisolant et des supports ont ete prises egales a 2 W/m2.

Les epaisseurs, les diametres et la nature des tubes constituant les differentes enceintes d'un cable sont indiques sur la figure 1.

3 — DlMENSIONNEMENT DES CONDUCTEURS ELECTRIQUES

Chaque conducteur de phase comporte un circuit supraconducteur et un circuit normalement conducteur place en parallele avec le premier, et destine a en assurer la stabilisation. Ces deux circuits doivent etre dimensionnes pour les regimes normaux d'utilisation et aussi en fonction des regimes perturbes.

92

d7 de

3.1 — DlMENSIONNEMENT DU CIRCUIT SUPRACONDUCTEUR COMPTE TENU DES SURINTEN-SITES DE DEFAUT

A chaque extremite de la liaison (fig. 2), on trouvera un poste de conversion comprenant un transformateur, un convertisseur fonctionnant soit en redresseur, soit en onduleur, et sur chaque conducteur de phase, une inductance de lissage.

I u

1

pf~l .ft. U : Tension entre poles

R -. Re'frige'rateur

33ra Transformateur

l r-yWP I — ^ \ Inductances

y de lissage

J Mmnp Convertisseur

Poste 1

--nnnnp-' I NGD^

Transformateur

Partie froide de la liaison

Fig. 2 — Structure de la cryoliaison

Convertisseur Poste 2

Un defaut se produisant en aval du poste 1, sur la liaison ou sur le r6seau qu'elle alimente, le courant va croitre brusquement. Sa croissance sera limitee d'une part par la presence des inductances de lissage, d'autre part par la variation des angles de retard a rallumage des redresseurs.

On peut, par l'emploi de systemes de protection particuliers rendre le reglage de ces angles pratiquement instantane. Dans ces conditions des etudes theoriques et experimentales developpees a E.D. F., [2] ont montre que la surintensite est maximale lorsque le defaut se produit sur la liaison elle-meme et lorsqu'il a lieu au debut d'une commutation.

Dans ce qui suit, nous nous placerons dans ces conditions, tout en remarquant que Papparition d'un defaut sur la partie froide de la liaison tres protegee par de multiples enceintes thermiques, est tres peu probable et qu'en consequence le defaut le plus dangereux a prendre en compte est un court-circuit aux bornes d'un des convertisseurs.

La duree et 1'amplitude de la surintensite de defaut vont alors dependre de la valeur des inductances de lissage. Pratiquement, comme pour les liaisons a courant continu classiques, et ceci afin d'assurer une regulation convenable du courant dans la liaison et surtout pour eviter tout risque de defaut dans les onduleurs, il sera necessaire de dimensionner convenablement ces inductances de lissage, pour limiter la vitesse de croissance du courant a une valeur qu'il semble raisonnable de prendre egale a 200 I„ par seconde. (Ce resultat sera obtenu en donnant au rapport : tension de fonctionnement U sur inductance totale de la liaison L, la valeur U/L ^ 200IM).

Dans ces conditions, on montre que la valeur maximale du courant de defaut est egale a Im = 1,8 I„, et que la duree de la surintensite est de l'ordre de 10 milli-secondes.

93

L'echauffement admissible pour les circuits electriques depend essentiellement des exigences d'exploitation. Si, quelques secondes apres la disparition de la surin­tensite de defaut, la liaison doit pouvoir etre remise en service, ce que nous suppo-serons ici, il faut que l'echaufTement des conducteurs reste faible.

Pour cela, il est tout d'abord necessaire que les circuits supraconducteurs ne transitent pas, c'est-a-dire que, compte tenu de l'accroissement du champ magnetique et de la temperature resultant de la surintensite, le courant dans la liaison reste a tout moment inferieur au courant critique des circuits supraconducteurs.

La croissance et la decroissance du courant etant des phenomenes lents, l'echauf­fement des circuits pendant la surintensite peut etre calcule en admettant que le comportement des materiaux et les processus de pertes correspondent au modele phenomenologique propose par Bean.

L'obligation de pouvoir remettre la liaison en service quelques secondes apres disparition de la surintensite se traduit alors par deux conditions [3] :

— la premiere indique que le courant critique a 4,2 °K de la liaison doit etre de Ic > 2 h (2)

— la seconde indique que le rapport \Jdx, du courant nominal In de la liaison au diametre dl du cylindre supraconducteur constituant un circuit de phase doit etre tel que :

\Jdx ^ 1,15.106. (3)

3.2 — DlMENSIONNEMENT DU CIRCUIT STABILISANT

La section du circuit stabilisant, selon les travaux de Cornish et Williams [4] est donnee par la formule :

S > ^ 20>A

ou — pM est la resistivite a 4,2 °K du metal stabilisant que nous supposerons ici etre

de Taluminium pM ~ 2.10" n Q.m. — O est le flux thermique pouvant etre evacue par l'helium de refroidissement,

la temperature du metal stabilisant restant inferieure a la temperature critique du supraconducteur. Pratiquement, le flux O est independant des conditions de l'ecoulement, et garde une valeur sensiblement constante et egale a 3 000 W/m2

lorsque la temperature du metal stabilisant passe de 4-5 °K (l'ebullition se fait alors par ebullition nuclee) a 10°K (le refroidissement se fait alors avec films de vapeur d'helium a la surface du conducteur).

— A est le perimetre mouille de metal stabilisant en contact avec l'helium de refri­geration.

Pour un conducteur stabilisant de forme tubulaire de diametre interieur d0 et servant de canal de refrigeration (fig. 1), A = nd0

et SM = - ( d ? - d g ) > 1,1. K T 1 5 - ^ (4) 4 d0

Dans la structure adoptee, le circuit stabilisant sert de canal de circulation pour l'helium liquide. Son epaisseur depend done aussi, de la pression de l'helium et du diametre d0. Pour Taluminium, et pour les pressions d'helium correspondant aux

94

FIGURE 3 - Tension optimale de fonctionnement pour une cryoliaison a courant continu en fonction de la puissance.

2000 4000 6000 8000 Puissance de la cryoliaison (MW)

0,9

0,8

0,7

0,6

0,5

0,3

0,2

0,1

t Cout specifique T minimal (F/kW.km)

\ \ \ \ ^

FIG

^—

URE 4 _ Cout sp< cryoliaison a ligne a^rienn courant con

Scifique minimal courant continu e classique a ' Mnu.

"—— « = i o

OCz 5

1 IX' s 1

fe. 8000

Puissance (MVA)

95

optima economiques (pression inferieure a 2,5.105 pascals) cette condition de resis­tance mecanique conduit a la formule :

^^^lO-'do (5)

Cependant pour des raisons technologiques, on a adopte les limites suivantes :

d0 ^ 5.10" 3m (6) et

^ ^ ^ 5 . 1 0 - 4 m (7)

4 — ASPECT ECONOMIQUE

Pour une liaison de puissance P, l'ensemble des formules (1), (2)...(7) permet de determiner en fonction d'un seul parametre U, les dimensions des circuits electriques. Les dimensions des canaux de circulation des fluides cryogeniques et en consequence les couts de refrigeration par circulation d'helium liquide et d'azote liquide se deduisent immSdiatement des resultats d'une autre etude publi6e par ailleurs [1].

On dispose done de tous les elements du cout de la cryoliaison mais avec une grande incertitude sur le prix des mises en oeuvre des differents composants.

Pour en tenir compte, on a introduit un facteur multiplicatif a, variable entre 1 et 10, qui permet a partir du prix des materiaux utilises pris a l'etat brut, d'evaluer le cout des elements du cryocable apres elaboration et montage, et ainsi de determiner une zone probable de cout de transport par cryocables.

Le calcul montre qu'il existe pour une puissance transitee donnee, une tension de fonctionnement optimale qui est independante de la longueur de la liaison, qui croit avec la puissance transitee mais ne varie pratiquement pas avec la valeur du coeffi­cient a (fig. 3).

La comparaison entre le transport d'energie electrique en courant continu et en courant alternatif est delicate, car si le premier est d'un prix de revient plus eleve, il presente des interets techniques qui pour certaines applications peuvent le rendre avantageux. On s'est done limite ici, a comparer les liaisons courant continu supra-conductrices aux liaisons courant continu classiques.

La figure 4 montre comment varie en fonction de la puissance installee les couts specifiques minima d'une liaison supraconductrice et d'une ligne aerienne courant continu. Ces couts prennent en compte les investissements propres aux cryocables ou aux lignes ainsi que le cout des pertes, mais non le prix des stations de conversion et les frais de genie civil ou d'acquisition de terrain.

On voit que pour des puissances inferieures a 5 000 MW, il est vraisemblable qu'une ligne de transport classique serait plus avantageuse qu'une liaison supracon­ductrice. II faut remarquer neanmoins qu'aux puissances elevees une cryoliaison aurait une tension de service tres inferieure a celle d'une liaison classique, ce qui pourrait reduire notablement le cout des stations de conversion.

D'autre part, il apparait qu'une cryoliaison supraconductrice serait moins onereuse qu'une liaison classique par cable a courant continu puisque le cout d'une telle liaison evalue dans les conditions indiquees ci-dessus se situe approximativement a 2 F/kW.km (pour des puissances d'environ 700 MW).

96

REFERENCES

[1] A. M. SCHWAB, G. DELILE et Y. JEGOU, Etude generale des problemes lies a la refrigeration d'une liaison electrique supraconductrice a courant continu. Commission I, Londres, Annexe 1969-1 Bull I.I.F. pp. 163-170.

[2] Surintensites dans les liaisons a courant continu — Note interne E. D.F . [3] Dimensionnement electrique d'un cable supraconducteur a courant continu — Note

interne E .D.F . [4] D.N. CORNISH and J .E.C. WILLIAMS. Phys. Letters (1965), 16, p. 18.

DISCUSSION

N. KURTI (U.K.) — Selon la figure 4 un cable supraconducteur coute moins qu'une ligne aerienne pour les hautes puissances; c'est-a-dire qu'il coutera, a fortiori, beaucoup moins cher qu'un cable classique souterrain. Est-ce exact?

G. DELILE — II resulte de notre etude que meme pour les puissances de l'ordre de 1 000 MW, le cout du transport (investissement et prix des pertes) par cable supra­conducteur a courant continu semble etre largement plus faible que le cout du trans­port par cable a courant continu classique.

II faut cependant noter que nos conclusions reposent sur des hypotheses concernant les prix, qu'il est a l'heure actuelle impossible de verifier.

97

MAQUETTE DE CRYOLIAISON HYPERCONDUCTRICE EN COURANT CONTINU

M. AUPOIX, F. MOISSON

Compagnie Generale d'Electricite, Laboratoires de Marcoussis

et

E. CARBONELL Centre d'Etudes Cryogeniques de la Societe L'Air Liquide, Sassenage {France)

A model for a direct current hyper conducting cryo-cable

SUMMARY: This report describes a model of a cryo-cable which was jointly realised by the Research Centre of the Compagnie Generale d'Electricite, Cables de Lyon and the Centre d'Etudes Cryogeniques de U Air Liquide. This model, of20 m in length, was tested for a nominal power o /80 MW with a nominal current of A 000 A.

This high-purity aluminium cable was designed for a constant length operating at low temperature. The operating temperature was 20 °K under a pressure of 20 bars of helium gas. The interior of the cable housing was superinsulated and it consisted of two cable lengths linked together with a compensating bellows against heat contraction. The temperature of the cable was maintained by a closed circuit system carrying cooled helium.

With this short-length model cable, the major part of the available cooling capacity was used to cool the leads. The ends of the cable were so designed as to minimise the heat flux in their lower sections, thus reducing as far as possible the refrigerating capacity requirements; this was done to avoid disturbing the thermal measurements on the cable.

The arrangement of temperature points and voltage take-offs along the length of the model enabled one to control the operation of the cable, current flow and the refrigeration of the model cryo-cable.

From the technical point of view this model cable represents an approach towards realising an industrial type cryo-cable.

Les cryoliaisons en courant continu ou alternatif sont actuellement l'objet dans le monde de plusieurs programmes de recherche et de developpement soit dans le domaine des hyperconducteurs avec l'emploi de metaux de haute purete dont la conductivity electrique est tres elevee soit dans le domaine des supraconducteurs.

Ces liaisons permettraient de transporter des puissances unitaires plus elevees que dans les cables classiques; en outre, une economie de puissance consommee peut etre realisee vis-a-vis de ces derniers malgre le faible rendement des cryogene-rateurs par suite de la valeur minime de l'ensemble des pertes a basse temperature.

L'etude de cryoliaisons a ete menee par le Centre de Recherches de la Compagnie Generale d'Electricite associe aux Cables de Lyon, avec la participation du Centre d'Etudes Cryogeniques de la Societe « L'Air Liquide ».

1 — CABLE (photographie n° 1)

Le procede de realisation retenu consiste a tirer le cable conducteur reposant a l'interieur d'une enceinte a meme temperature mise sous pression du fluide refri­gerant. L'helium sous 20 bars a ete adopte, ce qui permet d'eviter les difficultes d'une circulation sous double phase que presenterait une liaison en terrain accidente, et n'est pas dangereux. L'hydrogene aurait permis d'assurer une fiabilite fortement accrue pour I'isolant electrique rubanne impregne du fluide refrigerant mais n'a pas ete retenu pour des raisons de securite.

99

Photo 1 — Cryocable et enceintes associees.

Les dimensions adoptees pour le cable correspondent a la possibility de maintien en froid d'une cryoliaison de grande longueur avec un diametre de 7 cm pour la circulation du fluide cryogenique a l'interieur du cable. Le cable a ete realise aux « Cables de Lyon » au moyen d'aluminium contenant environ 30 p.p.m. d'impuretes; la purete de ce metal permet d'assurer un « recuit» partiel a la temperature ambiante; la resistivite du metal recuit est divisee par 1 200 lorsqu'il est porte de T ambiante a 4,2°K; a la temperature de fonctionnement (20°K), la resistivite apres mise en oeuvre vaut environ 1/280 fois celle du cuivre a la temperature ambiante (fig. 1).

Les contraintes mecaniques pouvant apparaitre dans le conducteur lors de la mise en froid peuvent etre prejudiciables pour la resistivite electrique du cable et pour la resistance mecanique de I'isolant electrique; dans le procede retenu, la longueur du cable reste invariante lors de la mise en froid sans creation de contraintes importantes parce que les nappes cylindriques formees de conducteurs en helices sont aptes a une reduction notable de leur diametre.

La structure du cable permet aussi d'enrouler de grandes longueurs sur un touret et de diminuer notablement le nombre des jonctions electriques.

Le cable etant de dimensions industrielles est experimente avec une densite de courant reduite 2A/mm2 avec une section voisine de 20 cm2.

Le cable et les conducteurs de transition comportent une isolation electrique mise sous pression d'helium, prevue pour une tension de service de 20 kV (essai en surtension a 40 kV). L'isolation est constitute par un rubannage de mylar (de 25 / j , pour une epaisseur totale de 3 mm).

2 — CONDUCTEURS DE TRANSITION (photographie n° 2)

Les pertes occasionnees par l'entree et la sortie des conducteurs aux deux extre-mites seraient negligeables vis-a-vis des pertes « en ligne » dans une liaison de plusieurs kilometres, mais sont preponderates sur le troncon court experimente.

Les conducteurs de transition comprennent une somme d'echangeurs a tempe­ratures intermediaires 20°K, 55 °K, 77 °K separes par des troncons conducteurs

100

10

1 )

8

9

AO |

- t l

1 il.ml

•i—+■ 1 T

i l l x I I I ^ 1 i j , *»^

/ /_

7 /

1 7 1/ / / ;7

/ /

M / •*7

/ / 1 l\ \ \

Hi 1

T °H 2 3 4 5- 6 7 8 9 10 20 30 AO SO 60 80 100 200 300 400

FIG. 1 - RESISTIVITE DE L» ALUMINIUM A99 EN T0NCTI0N DE LA TEMPERATURE

+ racial r e c u i t aprks niise en oeuvre

sans echanges thermiques avec le fluide refrigerant. La longueur et la section de ces derniers ont ete calculees pour minimiser le flux thermique atteignant les echangeurs.

: Staoeittt* 30* % JgD fears)

Enceinte t* asabianie Support isela&t | K'

/>- '*! * U-

r ^ ^ s ,. < - ■ <■

Photo 2 — Montage des enceintes et de la superisolation thermique.

Photo 3 — Conducteurs en transition 4 000 a 20 °C—20 °K.

102

3 — ENCEINTES ET ISOLATION THERMIQUE (photographie n° 3)

L'enceinte froide sous pression qui contient le cable est formee de 2 troncons rectilignes de 8 m dont la contraction thermique est compensee par des soufflets; cette longueur serait voisine de 20 metres pour une liaison industrielle.

La jonction entre les 2 troncons de l'enveloppe cryogenique permet d'isoler les volumes sous vide de chaque troncon; dans une liaison industrielle, ceci limiterait les consequences facheuses d'une fuite de Tune quelconque des enceintes.

Dans la gamme des temperatures voisines de 20°K, 1'isolation thermique de l'enceinte contenant le cable peut s'effectuer grace a un matelas de superisolant et a des supports isolants : la superisolation consiste en quelques dizaines de feuilles de « mylar » (polyterephtalate d'ethylene) aluminise, superposees avec un tassement faible, qui reduisent la transmission de chaleur lorsqu'elles sont placees sous un vide meilleur que 10 ~4 Torr.

4 — MAINTIEN ENJPROID

La circulation et le maintien en froid de l'helium sous 20 bars sont assures par 2 cryogenerateurs Philips PGH 105 fournissant des frigories a 20 °K et 55 °K, ce qui a conditionne la structure des conducteurs de transition.

Photo 4 — Vue generate de la maquette de cryocables hyperconducteurs.

5 PERSPECTIVES D'APPLICATION A UNE CRYOLIAISON INDUSTRIELLE

Les materiaux et les procedes adoptes pour la realisation de ce troncon experi­mental refletent la volonte de resoudre les problemes poses en tenant compte des imperatifs industriels lies a l'execution d'une cryoliaison reelle.

103

Une telle cryoliaison pourrait etre constitute par exemple de deux elements paralleles analogues au troncon experimental; pour une densite de courant 4 A/mm2

dans le troncon de cable realise et sous une tension de l'ordre de ±125 kV, la puis­sance transportee serait de 2 000 MW.

Les etudes et Fomentation de la realisation nous ont permis la mise au point de procedes de fabrication; la constitution des enceintes sous forme d'elements standards montes en usine facilite leur assemblage sur le site.

Les connaissances acquises dans cette etude pourront etre utilisees dans l'analyse et l'etude des cryoliaisons supraconductrices; elles representent, en outre, un acquis important dans le domaine des applications de la supraconductivite a l'electro-technique.

DISCUSSION

A. SELLMAIER (West Germany) — What is the heat influx per meter of the cable?

M. AUPOIX —0,5 W/m(effet Joule pour un courant de 4 000 Amperes) + 0,3W/m de pertes thermiques (heat influx) (superisolation sous vide et supports mecaniques de l'enceinte froide).

P.H. MELVILLE (R.U.) — Je voudrais demander pourquoi vous avez considere un systeme seulement pour courant continu et non pas aussi pour les courants alter-natifs.

M. AUPOIX—En alternatif, reflet de peau oblige a fractionner les conducteurs en fils fins ou en feuillards qu'il est necessaire de transposer; la profondeur de penetration du courant alternatif vaut en effet 8 = v /p/7iu/ ~ 0.5 mm pour l'aluminium que nous avons utilise.

Outre la complexity de fabrication de cables a conducteurs transposes, qui imposait la realisation d'une machine couteuse, leurs perspectives d'application sont limitees par suite d'une augmentation de la resistance du cable vis-a-vis de celle d'un cable de meme section en courant continu, lorsque I'epaisseur totale des conducteurs devient superieure a 20 5 ^ 1 cm; ce phenomene du a la creation de courants induits dans les brins conducteurs d'epaisseur non nulle (au moins 0,05 mm = 0,1x5) entraine une limitation du courant alternatif et par consequent de la puissance transportee qui reste inferieure au seuil de rentabilite d'un cryocable. Cet obstacle n'est pas rencontre en courant continu.

La maquette etudiee a constitue egalement une etape technologique necessaire vers les solutions supraconductrices en courant continu et alternatif.

104

SURVEY PAPER

A.C. LOSSES IN SUPERCONDUCTORS

B.B. GOODMAN The British Oxygen Company Research Centre, London (United Kingdom)

Les pertes dans les supraconducteurs en regime alternatif

RESUME: Afin d'etudier convenablement Vutilisation des supraconducteurs dans des appareils a courant alternatif, que ce soit aux frequences industrielles (cables, machines, transforma-teurs), oil a des frequences allant de 1 Hz (synchrotron supraconducteur) jusqu'a 109 Hz (cavites guide-ondes supraconducteurs a haut facteur de surtension), il est indispensable de bien comprendre les processus qui entrainent les pertes energetiques dans les supraconducteurs en regime alternatif.

On passe en revue les mecanismes principaux qui conduisent a de telles pertes, en tenant compte des quelques cas d'interet pratique, tout en soulignant la difference entre les phenomenes de surface et de volume.

A partir de la physique des pertes en regime alternatif on deduit des relations entre la grandeur de telles pertes et d'autres caracteristiques electriques d'un dispositif donne, ce qui pourrait en limiter le rendement global.

A la lumiere de cette revue il est possible non seulement de choisir le materiau supracon­ducteur qui conviendrait le mieux pour une application donnee, mais surtout de prevoir quelles ameliorations du rendement en regime alternatif peuvent etre apportees par rexistence de mate-riaux mieux etudies.

1. INTRODUCTION

Most of the large projects in superconducting engineering which are now under way involve large DC superconducting magnets, notably as research tools for high-energy or other branches of physics, or else, as in the case of the IRDC motor, as part of a DC machine. The effort being devoted to AC applications of superconductivity is much smaller, to a large extent because of the AC losses which occur in high field superconductors.

Such losses could be important:

a) In any applications of superconductors at industrial frequencies, whether 50-60 Hz in power networks or, for example, 400 Hz as in aircraft equipment;

b) In any DC application in which transients are important, e.g. the DC super­conducting cable [1];

c) In special applications, particularly for research purposes, with frequencies ranging from 1 Hz for the superconducting synchrotron [2] to several GHz for the superconducting linear accelerator [3].

A recent review by Wipf [4], which cites over a hundred references, shows that since the early measurements on a type II superconductor by Kamper [5] in 1962, much attention has been devoted to the question of AC losses. Nevertheless, in spite of the definite progress which has been made, engineers and physicists still do not have a clear enough understanding to be able to say with any certainty whether or not AC losses are likely always to prevent superconductors from being employed in one or other of the various aspects of AC power engineering for which they have been considered. From a purely historical point of view factors affecting this relatively slow progress have included the lack of familiarity of electrical engineers with the language being developed by physicists and the lack of involvement of physicists with the engineers' approach.

105

Furthermore, in normal conductors the problem of calculating AC losses is, by comparison with the superconducting case, far simpler. In principle it merely involves solving a set of linear equations (Maxwell's equations and Ohm's law).

In contrast to this simple behaviour, the electromagnetic behaviour of a super­conductor:

a) Can be non-linear; b) Can depend on the magnetic history of the specimen; c) Depends on the temperature of the specimen, and d) Depends critically on the exact metallurgical condition of both the surface and the

inside of the specimen.

2. THE APPROACH TO LOSSLESS BEHAVIOUR IN THE MEISSNER STATE

The AC losses in a superconductor depend critically on the way in which it is penetrated by a magnetic field. If the superconductor is cooled through its transition temperature in zero field, and if a weak magnetic field is subsequently applied, the specimen remains in the pure superconducting or Meissner state; the magnetic field in the specimen, b, then obeys the London equation

b = A,2 V 2 b, (1)

where the penetration depth, X, is typically of the order of 0.1 um. Equation (1) implies that a magnetic field is excluded from all but a surface layer of depth of the order of X,.

When a superconductor is in the Meissner state described by equation (1) AC losses are vanishingly small at all frequencies below the microwave range ( < 100 MHz) [6].

However, the almost complete flux exclusion described by equation (1) does not persist up to indefinitely large fields. In type I superconductors normal, flux carrying, tubular or laminar regions penetrate the specimen, running from one part of its surface to another, when the field at the surface exceeds the critical value Hc [7]. In type II superconductors quantized vortex lines, each carrying a flux of approxi­mately 2+ 10" 1 5 weber, similarly enter the specimen, thus setting up the mixed state, when the field at its surface exceeds the lower critical field H c l [8]. When flux has entered a superconductor of either type in this way its AC behaviour is no longer lossless and even if the external field is subsequently removed, flux usually remains trapped in the specimen.

In order to approach the ideally lossless AC behaviour characteristic of the Meissner state (e.g. in a superconducting AC cable or linear accelerator) it is therefore necessary;

a) To cool the superconductor through its transition temperature in as low a magnetic field as possible, in order to avoid trapping magnetic flux. Magnetic fields which might be trapped include the earth's field, fields due to currents of thermoelectric origin in a non-isothermal environment and the field due to a fault current in one phase of an AC cable;

b) to ensure that the maximum operating field is lower than Hc or H c l , as the case may be, by a safe margin. Relatively pure niobium is preferred because its value of H c l is higher than that of H c or H c l for any other superconductor;

c) to ensure that surface irregularities which could locally accentuate the field produced during the operation of the device and thus trap flux [9], are absent.

106

In niobium at 50 Hz and 4.2 K the AC losses have generally been found [4] (Fig. 1) to be less than the maximum permissible value of 0.1 W m - 2 provided the amplitude of the surface field is less than about 6 x 104 A m " 1 , or only about 50% of H c l at 4.2 K (11 x 104 A m~*). More recently Easson and Hlawiczka have reported [10] losses of less than 0.1 W m~ 2 at a surface field of 11 x 104 A m - 1 . However it is not yet known whether it is economically feasible to produce a niobium sheathing for a

CO in

10"

I 0 " b ^

10" LU

O CVJ E

ft 10" O

10" f 9

10' ■10

I0"1

I I I I Mill i i i Mini I i i i I a b

Nb

,m

10 I I 1 l l l l l l I I I 1 Mill I I I I Mill

1 I02 I03 I04

PEAK SURFACE FIELD (Oe)

l/i C B <u OH

B 3 o

X

6

© N X

. o

13s o <u x

1 cti O

GO

107

commercial AC cable for which the low loss region would extend to such a high field; nor is it known whether this value of the field represents the limit of what might be technically possible, for example given an appropriate treatment of the surface of the niobium.

3 . A C LOSSES IN THE MIXED STATE

So far we have principally been concerned with the ideally resistanceless Meissner state, which is only of practical interest in relatively low field devices. In the operation of a high field device vortex lines are nucleated at the surface of the type II super­conductor by the application of a field exceeding H c l , thus driving it into the mixed state.

Figure 2 shows how the magnetic induction, b, varies with the distance r from the centre of an isolated vortex line. At the centre of the vortex line b(0) = 2jioHcl; b(r) decays to zero over a distance of the order of X, varying approximately as exp (-r/X)forr>>X.

£ l h

r/x Fig. 2 — Variation of the magnetic induction b(r) as a function of the distance r from the

centre of an isolated vortex line.

The total flux in an isolated vortex line is equal to the so-called fluxoid quantum,

2nrdrb(r) = 4>0 . (2)

<|>o-2x 10 1 5 weber:

i: A vortex line requires energy to be created, the principal contributions to the energy

density per unit volume being the magnetic field energy (b2/2u.0) and the kinetic energy density of the current which circulates around the vortex line (X2 (curl b)2/2|i0). Suppose now that a vortex line, whose field is bl9 is placed in a region where there already existed a field b 2 , due to some other source. The latter could be either the surface penetration field, given by equation (1), or the field of another vortex line.

108

The energy of the system can now be evaluated by putting b = bx + b 2 into the expressions for the contributions due to the energy density given above. The existence of the cross product terms in bx .b2 and in (curl b^.Ccurl b2) shows that there exists an interaction energy between the vortex and the field b2 . This energy corres­ponds to a repulsive force either:

a) Between the London penetration field and any vortex lying in its vicinity, i.e. within a distance of the order of X of the surface of the specimen, or

b) Between two parallel vortices within a distance of the order of X of each other.

In a perfectly homogeneous specimen vortices tend to form a regular triangular array [11]. However in type II superconductors of practical interest strong pin­ning forces are exerted on vortex lines. This arises because, owing to the existence of inhomogeneities and other extended defects in the specimen, X is position dependent; the existence of the term in A,2(curl b)2/2n0 for the self energy of a vortex line suffices to show that the self energy of a vortex line is also position dependent, i.e. that there exists a mechanism for pinning,

As the external magnetic field H c at the surface of a type II superconductor with strong flux pinning is gradually increased, there arises a situation illustrated in figure 3. The total magnetic field, b total, is the sum of the London penetration field, bL, given by equation (1), and the fields b v of all the individual vortex lines, whose positions are indicated by dotted lines. The non-uniform distribution of vortex lines around any given vortex line (greatly exaggerated in figure 3 for the sake of clarity) ensures that there exists a net repulsive force on that line which can just overcome its own pinning force.

Fig. 3 — Spatial variation of btotal, the total local magnetic induction in a type II super­conductor showing strong flux pinning, as a function of x, the distance from the surface of the specimen; B is the average value of btotal. Ha is the external field.

Figure 3 also shows that, on a macroscopic scale, there is a discontinuity J I 0 A H between the external field |X0HC and the mean internal field near the surface

B(O) = j i 0 ( H e - A H ) . (3)

109

The quantity AH is clearly connected both with the nature of the vortex distribution near the surface of the specimen, and with the presence of the London penetration

• field bL- In reversibly behaved specimens AH is equal in magnitude to the magneti­sation and can be calculated theoretically [8], but the factors which determine AH in irreversibly behaved specimens have not been exhaustively studied.

Well inside the specimen the mean density of vortex lines N, (over a distance large compared with both X and with the vortex line spacing) is related to the mean field by B

B = ^ N , (4) so that a spatial variation in N corresponds not only to the existence of a net force acting on a given vortex line, but also to an average transport current density J:

J = jx0 1 curl B = (</>o/m) CU1"1 N . (5)

The maximum value of J which can be achieved in a given specimen (at given values of the temperature and of B) is therefore a measure of the pinning force which must be overcome in order to displace the vortex line array.

While J can be enormous by conventional engineering standards, viz. 103-104 A m m " 2 ,

it is rather small compared with the current density of the order of H c l fk ~ 106 A mm ~ 2

which circulates either in the London penetration layer in the Meissner state, or around an isolated vortex line in the mixed state. This observation corresponds to the fact that in figure 3 dB/dx is much less than the maximum values of either dbL/dx dbv/dx.

While much still remains to be done to put flux pinning on a quantitative micros­copic basis, there is now strong evidence [4] in favour of the critical state model, originally proposed by Bean [12] and London [13]. According to this model vortex lines only move when an attempt is made to exceed a certain critical current Jc, or critical value of curl N, i.e. they only move when the maximum available pinning force is exceeded.

While the critical fields H c l and H c 2 are thermodynamic quantities, and are sensibly the same for all specimens of the same composition and temperature, Jc is, like yield strength or magnetic coercivity, an extended defect controlled property, which varies widely among specimens of similar composition.

[THICKNESS OR ! DIAMETER OF

-[-SPECIMEN

MAGNITUDE OF FIELD

Fig. 4 — Spatial variation of the magnetic induction inside a cylinder or slab parallel to an external field He, illustrating the four types of possible regime according to value of He in relation to Hcl, Hp and Hc2.

110

Figure 4 illustrates the penetration of a cylinder or slab of superconductor by a field, parallel to its surface and distinguishes 4 regimes, according to the peak value of the external field. On the scale of this diagram the structure in the curve representing btotai in figure 3 is not perceptible.

Four different types of regime can be distinguished, according to the maximum value of the external field, He:

I H e < H c l

II H c l < H e < H p

III H p < H c < H c 2

IV H c 2 < H e

In regime I, discussed in section 2, the losses are essentially zero if the specimen surface is sufficiently perfect. In regime IV the specimen is in the normal state, except possibly for a thin surface layer in fields up to H c 3 = 1.69 Hc 2 , but this is of no practical engineering interest, so we are left with regimes II and III.

In regime III, for which the external field is large enough for the entire specimen to be penetrated by flux, rather little information has so far been obtained. This is to some extent due to the fact that for the size of specimen usually studied (0.1 to 1 mm in thickness) the heat losses per unit surface area are sometimes sufficient to cause thermal run away.

Measurements in regime II have been carried out:

a) by calorimetric determination of the energy dissipated; b) by electrical determination of the energy dissipated; c) by tracing magnetization curves, and d) by observing the damped mechanical oscillations of the specimen in a magnetic

field.

In order to relate the measurements to the critical state model it is mathematically convenient to assume that Jc is independent of B, although it is known that this is not exact, and more general expressions for JC(B) have been studied [14-16, 20].

Furthermore, in order to take account of the current carrying properties of the surface of the specimen, we shall assume that a surface current density can vary in the range from — AH to + AH without modifying the value of B immediately inside the specimen. If it is assumed, for simplicity, that AH is independent of field then, for a plane surface exposed to a field of amplitude Hw, the energy dissipated per unit area and per cycle may be written [15, 16].

W = (2^0/3 Jc) (Hm - AH)2 (Hm+2AH) . (6)

An important feature of this equation is that the energy lost per cycle is independent of the frequency, a result which remains true if the critical state hypothesis holds, no matter how the bulk and surface critical current density depend on field.

Figure 5 shows the results of several measurements on various niobium alloys, recently reviewed by Wipf [4] which, except for fields greater than about 3 kilo oersted (2.5 x l 0 5 A m _ 1 ) , all refer to regime II. The scatter in the results probably arises not only from genuine differences between the specimens, but possibly also from diffe­rences between the magnetic histories^ of the various specimens before the measu-ments were made. Nevertheless the results are qualitatively in agreement with equation (6) if Jc is taken to be 300 to 600 A m m - 2 , but the departure from a simple propor­tionality to H m

3 at low fields is more gradual than indicated by the simple equation (6), and so only the order of magnitude of AH can be given for the measurements, i.e. of the order of a few hundred oersted (104-105 A m~ 1 ) .

I l l

ICT3 b -I I I I I I I I 1 1 I I M i l l ~~1 I I La! IT

/ ,' ~ / / ~

/ < > c Z

Nb-ALLOYS

10 I 0 2 103

PEAK SURFACE FIELD (Oe)

"1 I I 111

I04

Fig. 5 — AC loss per unit area and per cycle for various niobium alloys, after Wipf [4].

112

More careful attention to the quantity AH has been paid by Ullmaier and Gauster [17], whose measurements on niobium-25% zirconium are reproduced in figure 6. Taking, for this alloy at 4 K, [i0Hc = 0.16 T, uH c 2 = 8 T, it may be estimated [8] that uH c l = 0.015 T, which is appreciably smaller than the low field result |i0AH = 0.05 T.

500

4 0 0

to a 200 & UJ

o Z. o * 5 Q O

4 0 0

2 0 0

C4

D

1

*

o

A

1 1

9 $ * ° * 3 I 41% M

O I 2 3 4 EXTERNAL FIELD (KILO-OERSTEDS)

Fig. 6 — Measurements of AH at 4.2 K in niobium-25% zirconium by Ullmaier and Gauster [17]. The two graphs refer to different specimens; for each specimen the different symbols refer to slightly different experimental conditions.

As mentioned above, no complete explanation can yet be offered for a value of AH/Hc l greater than unity. Park [19] has considered the contribution to AH arising from the de Gennes-Saint James superconducting sheath, but this calculation is not applicable at low fields. A further possible mechanism is that vortices within a distance of the order of X of the specimen surface, which are attracted towards the surface by their image vortices, are particularly easily pinned by extended defects in the specimen, so that locally a high value of curl N or Jc can occur.

If it were possible to achieve such a high value of AH/Hc l in niobium, the current carrying capacity of a niobium AC cable would be correspondingly increased, but recent measurements [10] are not very encouraging in this respect.

If AC losses cannot be much further reduced by increasing AH, which is certainly the case for fields greater than about 1 T, there is one line of approach which is begin-

113

ning to receive attention, and which concerns regime III. As originally pointed out by London [13], the AC losses per unit volume of conductor in this regime decrease with decreasing conductor size. The recent development of composite superconductors [18] in which tens of niobium-titanium filaments, each 10-20 urn in diameter, are embedded in a normal matrix, is a step in this direction, even though a major reason for adopting such fine filaments was to render flux jumping energetically improbable.

A particular difficulty with such composite superconductors is that circulating currents of very long time constant can be set up. Thus flux can only enter or leave the normal substrate by diffusing from either end of the wire, the corresponding time constant being given by

x = l2IDn2

where: / is the length of the wire and

D = (p/ja0) is the electromagnetic diffusivity of the normal substrate of resistivity p. Taking, as a typical value for OFHCcopper a t 4 K , p ~ 10~1 0 ohm m, and taking

/ = 1 km, x ~ 109 sec ~ 30 years. For AC applications at a frequency / one necessary condition for losses to be

small is that /x<^ 1, and it is hoped to achieve this:-

a) by employing a high resistivity substrate, such as cupronickel, and b) by twisting the whole conductor with a pitch p < /. With such an arrangement the

circulating currents which could be induced by an external magnetic field would have a time constant only of the order of/?2/D7i2. With/7 = 1 cm x ~ 10~4 sec for a cupronickel substrate.

Although little work has yet been done to confirm the soundness of these ideas, initial results are very promising and the way may possibly be opened at least for very low frequency (1 Hz) applications.

4. CONCLUSION

a) The last few years have seen a better understanding of the fundamentals of AC losses, particularly through the use of the critical state model. This model has been particularly important in showing how the effects on the AC losses of either the specimen geometry or of the superconductor properties (Jc, AH) can be separated. Work is now in hand on generalizing this approach to various types of field dependence of the bulk and surface current densities.

b) Work on the relation between the material properties Jc and AH and the metal­lurgical structure of promising alloys will undoubtedly be pursued, the factors determining AH in low fields being of particular interest.

c) Probably the most promising lines of approach in attempting to reduce AC losses in superconducting devices will be concerned with new geometrical arrange­ments of conductor, e.g. by twisting or transposing. A further posibility is the use of polyphase composite, with a normal metal of high resistivity ensuring low eddy current losses and a normal metal of low resistivity fulfilling the protection function.

ACKNOWLEDGEMENTS

1 would like particularly to thank Dr. S.L. Wipf for allowing me to use before publication material from his forthcoming review article.

114

REFERENCES

[1] R.L. GARWIN and J. MATISOO, Proc. IEEE, 55 (1967), p. 534. [2] P.F. SMITH and J.D. LEWIN, Nuclear Instruments and Methods, 52 (1967), p. 298. [3] T.I. SMITH, Advances in Cryogenic Engineering, 13 (1968), p. 102. [4] S.L. WIPF, Proceedings of the 1968 Brookhaven Summer Study, to be published. [5] R. A. KAMPER, Physics Letters, 2 (1962), 290. [6] C.G. KUPER. An Introduction to the Theory of Superconductivity, Oxford (1968),

pp. 58-61. [7] C.G. KUPER, idem, Ch. 1. [8] B.B. GOODMAN, Reports on Progress in Physics, 29 (1966), p. 445. [9] T.A. BUCHHOLD, Cryogenics, 3 (1963), p. 141.

[10] R.M. EASSON and P. HLAWICZKA, Brit. J. Appl. Phys., {J. Phys. D), Ser. 2, 1 (1968), p. 1477.

[11] H. TRAUBLE and U. ESSMANN, / . Appl. Phys., 39 (1968), p. 4052. [12] C.P. BEAN. Phys. Rev. Letters, 8 (1962), p. 250.

C.P. BEAN. Rev. Mod. Phys., 36 (1964), p. 31. [13] H. LONDON. Physics Letters, 6 (1963), p. 162. [14] I.M. GREEN and P. HLAWICZKA, Proc. IEEE, 114 (1967), p. 1329. [15] W.I. DUNN and P. HLAWICZKA, Brit. J. Appl. Phys, Ser. 2, 1 (1968), p. 1469. [16] G. FOURNET and A. MAILFERT, (private communication, 1968). [17] H.A. ULLMAIER and W.F. GAUSTER, / . Appl. Phys., 37 (1966), p. 4519. [18] P.F. SMITH, Proceedings of the 1968 Brookhaven Summer Study, to be published. [19] J.G. PARK, Phys. Rev. Letters, 15 (1965), p. 352. [20] R. HANCOX, Proc. I.E.E., 113 (1966), p. 1221.

DISCUSSION

T. BUCHHOLD (West Germany) — An interesting effect was observed with cylindrical Nb samples which were carefully machined and polished. If such a sample was cooled down quickly the losses were small. However, if the cooling down was slow, e.g. overnight cooling, the losses could be more than 100 times higher. If the sample was warmed up and then cooled down quickly the losses were small again. The losses were independent of the cooling rate if the sample was annealed.

P.H. MELVILLE (U.K.) — Perhaps I can throw a little light on the mysterious AH. In the results of Easson and Hlawiczka and also Schwechzer on low K type II materials, the change in flux on field reversal does not occur until near H c l . In increa­sing field the flux density in a hysteretic material is less than the free energy minimum as determined by an equivalent reversible curve. Thus on field reversal we do not expect B to change until the value of B at the surface is less than that given by the reversible curve. For small Bthis change occurs near H c l in which case 2 AH is the difference between the peak field and H c l . The situation becomes more complicated for large B and in high K superconductors.

115

COMPOSITE CONDUCTORS FOR A SUPERCONDUCTING AC. POWER TRANSMISSION CABLE

M.T. TAYLOR Materials Division, Central Electricity Research Laboratories

Leatherhead {United Kingdom)

Conducteurs composites pour uite ligne supraccn due trice a courant alternatif

RESUME : Les constructions preferees des liaisons supraconductrices s'appuient sur Vemploi de conducteurs tubulaires. Ces tubes ont une pellicule supraconductrice sur les deux surfaces ce qui empeche Vechauffement de tout metal normalement conducteur par courants de Foucault. Le conducteur doit: (1) porter le courant normal de fonctionnement avec une perte inferieure a 10 |aW cm - 2 ; (2) porter pour une seconde un courant de defaut sept fois plus grand que celui du fonctionnement

normal. La perte doit etre infer ieure a 60 mW cm - 2 ; on prefere une perte de 50 mW cm - 2 ; (3)\empecher Vauto-destruction en cas de panne du systeme cryogenique; (4) comprendre une securite structurale et contenir le refrigerant.

Pour pouvoir repondre a ces fonctions le conducteur aura deux ou plusieurs composantes. Niobium et matieres dures du type II telles que les alliages Nb-Ti ou Nb-Zr sont des composantes supraconductrices convenables, tandis que le cuivre et Valuminium ont ete consideres pour les fonctions 2-4. II semble que le conducteur le plus satisfaisant consiste en un tube de cuivre revetu d'une couche de supraconducteur de type II (epaisseur d'environ 80 y,m), laquelle est recouverte d'une couche de niobium (epaisseur d'environ 10 \im).

1. INTRODUCTION

There have been several proposals for superconducting a.c. power transmission cables. Detailed consideration has been given to a 750 M.V.A. 3 phase cable [1] and indications are that high power cables ~ 2000 M.V.A. could be competitive with conventional underground cables [2]. The most favoured designs are those based on tubular conductors. There are several possible arrangements to form a 3 phase cable [3]. Whatever the configuration the conductor has to serve the following func­tions: (a) Carry the normal operating current with a power loss below a specified level*

| i.e. lOuAVcm"2. (b) Carry a fault current equal to seven times full load current and limit the heat

produced so that the cable is still operable. The size of this heat pulse imposes severe restraints on the cryogenic design of a cable. A loss of 50 mW c m " 2 would allow the cryogenic engineer complete freedom of choice and reduce the quantity of expensive helium required. A loss of 600 mW c m - 2 is barely acceptable.

(c) Protect against the self-destruction of the conductors in the event of a failure of | the cryogenic envelope during operation of the cable.

(d) Provide structural integrity and contain the refrigerant, helium. Since vacuum is a possible dielectric, the tubular conductors may have to support both bursting and collapsing stresses.

Table 1 lists the operational parameters for a proposed 750 M.V.A. 3 phase cable. The discussion in this paper of the material requirements of conductor components will be carried out using the data in this table.

In order to perform the functions listed above it is obvious the conductor will have to be a composite with at least two components. Niobium is almost universally

119

Table 1

PARAMETERS FOR A 750 MVA-3 PHASE 50 Hz CABLE

Parameter Normal Operation Fault Condition

1. Refrigerant Pressure, atmospheres Temperature, °K

Liquid or supercritical helium In range 3-10 In range 4-6.5

2. Electrical O (a) Current/phase, A (b) Interphase voltage, kV (c) Current density A mm"

rms value peak value

(d) Surface field hs, rms value A m" l

rms value Oe hp, peak value A m~1

peak value Oe (e) Loss

Preferred Upper Limit

13,000 33

40 56.6

4 104

500 5.66 104

715

<1 uAV cm"2

lOuAVcnr2

91,000 210

280 396

2.8 105

3500 3.96 105

5000

^ 5 0 m W c m - 2

600 mW cm"2

O Parameters 2 (c)-(e) refer to the smallest diameter conductor only (10 cm diameter), larger diameter conductors are less severly loaded (oc radius-1).

considered for function (a), but it cannot fulfil any of the other functions. To carry the fault current three possibilities have been considered; aluminium, copper and a hard type II superconductor. Of these, both aluminium and copper are suitable for the protection function.

The conductors are thus conceived as copper or aluminium tubes with a super­conducting skin which can be on either the inside or the outside surface of the tube to prevent eddy currents in the normal metal.

2. NORMAL OPERATION

In order to operate a superconductor under a.c. conditions and obtain essentially zero loss a type I or a type II superconductor operating below Hci should be used. Since the cable may be operated at the relatively high temperature of 6.5 °K, all known type I superconductors are eliminated leaving Nb as the only candidate. Pure Nb [4] at 6.5°K has HCi equal to 915 Oe (7.3 104 A m " 1 ) and H C 2 equal to 1450 Oe (11.5 104 A m - 1 ) . From table 1 it is obvious that the niobium will become normal under fault conditions. Losses are observed in Nb even when operating in fields less than Hci [5]. However if care is taken to produce sufficiently smooth surfaces, it should be possible to operate with a surface current density of 40 A mm ~* and keep the loss within the specified limit. The critical fields of Nb are structure sensitive, cold working tends to increase H C 2 and lower HCi and is therefore a point to be watched.

The minimum thickness of the niobium layer is obviously determined by the penetration depth [4], which is approximately 0.05 urn at 6.5 °K therefore a thickness in excess of 0.2 |im is required. The maximum allowed thickness can be determined from the loss in the niobium layer under fault condition, see 3.5.

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3. FAULT CONDITION

Since the fault current cannot be carried by the niobium it has to be backed by a low resistivity material or a hard type II superconductor. The important material parameters for the copper and the aluminium are the resistance ratio, R, between room temperature and the operating temperature and the magneto-resistance.

3.1 Assuming the conductor can be treated as a semi-infinite plane slab, then the loss, P, in W m ~ 2 is given by

P = H0co8bs2/2 (1)

where 8 = {2pRT/^i0coR}^ is the skin depth, p R T is the resistivity at room temperature, |X0 is the permeability of free space and co the angular frequency. Equation 1 is a good approximation provided the conductor thickness, t, is very much smaller than its radius and greater than 1.2 5.

In table 2 are shown the values of R necessary for copper and aluminium to meet both the preferred and upper limit heat loss. It is obvious that commercially available copper and aluminium are barely acceptable.

Table 2 MATERIALS PARAMETERS FOR FAULT OPERATION

Materials parameter required for

Material Preferred loss of 50 mW cm - 2 Upper limit loss of 600 mW cm"2

Copper R = 83 103 R = 575 {R(0) = 2000} (a) Aluminium R = 134 103 R = 930 (R(0) = 2500} (fl) Hard Type II Material Jc = 5 105 A cm"2 Jc = 4 104 A cm"2

(a) Value of R(0) required to make R = R(H) for H = 5 kOe.

3.2 The effect of magneto-resistance is important for both copper and aluminium even for the low fields occurring in a cable. The resistance ratio in the field, R(H), is less than its value in zero field R(O), and can be estimated within 20% from the relations [6]:

RAL(0) = RA1(H) \l + J^™^Ll (2) K I 1.8 + 1.6H.+0.53H.2J

and

Rcu(0) = RCu(H) [ l + .279 — ^ — 1 (3) L H* + 1.4j

in these expressions H* = R(0) H 10" 6 where H is the magnetic field in Oe. To calculate the effect of magneto-resistance on the a.c. loss is difficult. Equations (2) and (3) have been used to estimate the value of R(0) to make R(H) for H = 5 kOe (3.96 105 A m _ 1 ) equal to the values of R determined in 3.1, see table 2. This gives an estimate of the kind of purity required for the copper and aluminium. Commercial quantities of copper with R = 2000 are not available. Aluminium is available with resistance ratios of several thousands, and although this is likely to be reduced during the fabrication of a composite with niobium, it is however just acceptable to fulfil the fault function.

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3.3 The loss produced in a semi-infinite slab of hard type II superconductor is given by the relation [5]

P = H 0 C O ( H P - A H ) 3 / 3 T I J C (4)

where Jc the critical current density and AH are material parameters. For the fields present during a fault the factor AH is only a few 100 Oe and will be ignored. The values of Jc necessary to meet the preferred and upper limit heat losses are listed in table 2.

Values of Jc in excess of 5 105 A c m - 2 have been achieved by many commercial materials at 4.2°K and 5 kOe, for example N b - 4 4 % Ti [7], N b - 2 5 % Zr [8], Nb - 40% Zr - 1 0 % Ti [9]. However data at 6.5°K is sparse, but commercial Nb - Ti[7] has a Jc of 2.3 105 A cm" 2 and heat treated Nb - 25% Zr [8] a value of 3.7 105 A cm" 2. The compound Nb3Sn is not being considered because of its extreme brittleness. Thus the reduction of the losses in the fault condition to the preferred values seems probable with a hard Type II layer.

3.4 At 50 Hz, the appearance of a fault current produces very large rates of field change, which can cause flux instabilities such that the material ceases to superconduct. The maximum field a superconductor can shield, IIM, has been calculated by Swartz and Bean [10] who assume adiabatic conditions, and a simple critical state model with Jc independent of field but decreasing linearly to zero at T*. They obtained the following relation:

hM = {5n l O ^ T j - T ^ A m " 1 (5)

where C is the specific heat in J °K " 1 m " 3. Using data on Nb - 25 % Zr for (T* - T)8

and only the T3 term in the specific heat then AM = 185 A m m " 1 at 6.5°K and 128 A m m " 1 at 4.2°K. Since these values are less than the peak fault field of 396 A m m " 1 there may be serious instability problems. Experimental data [12] on Nb — 25% Zr at field sweep rates comparable with, but smaller than, those to be found in a cable gave values of hu between 319 and 350 A m m " 1 at 4.2°K. The theoretical estimates of hu are probably pessimistic because they do not take into account the effects of Hci and AH. From equation 5 it is obvious that a supercon­ductor with a large Tc is required and operation at the higher temperatures is indicated because the specific heat increases rapidly with temperature.

The thickness of the hard type II layer, A, is determined by hp and Jc. If Jc = 5 105 A c m " 2 and hp = 3.96 105 A m " \ since A = hpIJc then A = 80 urn.

3.5 A further matter that requires examination particularly if a hard type II superconductor is used, is the loss produced in the non-superconducting Nb layer. A simple estimate can be given by assuming the Nb layer is completely normal and very thin, the loss in such a layer, when backed by a material of similar resistivity, is given by.

P = |x0co ah] a <t 5 (6)

Note that P is independent of the resistivity of the layer. Thus for hs = 2.8 105 A m - 1

and a = 10 urn, then P is 30 mW c m - 2 and is comparable with the loss in a hard type II material. Although this is an overestimate it indicates that the Nb layer must be thin ~ 10 urn and its losses cannot be completely ignored. If Al is used for carrying the fault current the thickness of the Nb layer is determined by manufacturing con­siderations.

122

3.6 Thus to give the engineers flexibility, a conductor consisting of a hard type II layer to carry the fault current should be developed; a superconductor such as Nb 25% Zr appears suitable because it is ductile and has a large Tc, it must have a thin layer of Nb to carry the normal operational current. The question of instabilities must be very seriously examined.

4. PROTECTION

The problem of preventing the self destruction of the conductors in the event of some damage to the cryogenic envelope, which results in an excessive heat input, has not previously been discussed, although the problem has been examined for large superconducting magnets [13], there are a very large number of unknown quantities. In the following analysis it is assumed for simplicity that the heat entering the cable is completely absorbed by the refrigerant, whose temperature locally exceeds the critical temperature of the superconductor. Another assumption is that all the Joule heating in the conductors will be absorbed by the conductor itself.

If the maximum temperature to which the conductor can be allowed to rise is Tm and xw is the time taken to discover the failure and break the current, then the condition on the surface current density, hs is

h, < L\i0(OT -1U F(Tm)

where (7)

F2(TJ = {p(TJ}* C„dT

S and / is the thickness of the normal metal. Equation 7 was derived by assuming the Joule heating was always equal to that for a semi-infinite slab, which implies O l . 2 5(Tm).

10 20 30 4 0 50 TEMPERATURE. Tm°K

Fig. 1 — Relation between surface current density and maximum temperature for protection of conductors against cryogenic envelope failure.

123

The simplest and maybe the cheapest method of detecting damage to the cryogenic envelope is to monitor the pressure rise in the refrigerant. Since the damage envisaged is localised, any increase in pressure will propagate with the velocity of sound. The velocity of sound in helium at the contemplated temperatures and pressures varies from approximately 200 m s - 1 to 600 m s " 1 . Considering the worst case (damage exactly half way between refrigerator stations [2] and the slowest velocity) gives a value of 30 sees for TW.

In figure 1, hs is presented for both copper and aluminium as a function of Tm; it was assumed that t = 1.2 5(Tm) to keep the quantity of backing material to a mini­mum. It should be noted that the allowed value of hs is insensitive to resistance ratio and copper is superior to aluminium because of its larger density. In order to limit thermally induced stresses Tm should be less than 50°K for both aluminium and copper.

Cryogenic failure occurring simultaneously with the onset of the fault condition does not impose any further restrictions because the circuit will be broken in less than one second. Probably the most serious problem associated with cryogenic failure —even the minor failure considered here—is that of pressure rise in the refrigerant. This is also difficult to calculate. Both copper and aluminium could fulfil the protection function, but copper is preferred because it has a higher mechanical yield stress.

5. STRUCTURAL

The major stresses on the conductor are due to the pressure of the refrigerant, and to thermal contraction.

On grounds of thermal contraction copper is to be preferred since the total contraction from room to liquid helium temperature is 0.33% compared with 0.44% for aluminium. Some means of compensating for this large thermal contraction must be found, i.e. bellow sections etc.

Table 3 STRUCTURAL COMPONENT PARAMETERS

Material

Copper Aluminium Stainless steel

E kg mm 2 -

14 103

7 103

21 103

Yield Stress

kg mm - 2

12 3.5

70

//D (14

Bursting

0.006 0.020 0.001

atmospheres)

Collapse

0.013 0.017 0.011

To provide the mechanical strength it would be convenient not to have to provide a strengthening component to the conductor or add backing material in excess of the electrical requirements. In table 3, the Young's Modulus, yield strength of copper, aluminium and stainless steel are listed along with the thickness to diameter ratios required for a pressure of 14 atmospheres. This pressure was used to allow a safety margin over the maximum possible pressure of 10 atmospheres. If the conductor has a diameter of 100 mm then the copper would have to be 1.3 mm thick which is approximately the skin depth for copper with R = 200 required for protection, see figure 1. The aluminium would have to be 2 mm thick which is also approximately the thickness required for protection. However, if the conductors have to be stabilized against collapse there is little point in strengthening the aluminium with stainless steel.

124

6. CONCLUSION

A conductor consisting of an aluminium tube (2 mm wall thickness) with a niobium coating (25-50 um thick) would be just acceptable for a superconducting power cables the aluminium would have to keep its purity during processing because a resistance ratio of approximately 2500 : 1 is required.

It appears a conductor consisting of a thin niobium coating (~10 jam thick) over a hard type II superconducting layer (~ 80 urn thick) on a copper tube (1.5 mm wall thickness) would be the most satisfactory. Since the copper performs only the protec­tion and structural functions it therefore does not need to be particularly pure. The alloy Nb-25% Zr appears to be a suitable candidate for the hard type II superconductor. This conductor might impose some limitations on the allowed surface current densities because of magnetic instabilities in the hard type II layer.

ACKNOWLEDGEMENT. The work presented in this paper was carried out at the Central Electricity Research Laboratories and is published by permission of the Central Electricity Generating Board.

REFERENCES

[1] E.C. ROGERS and D.R. EDWARDS, Electrical Review, 181, 348 (1967). [2] D.A. SWIFT, I.I.R. 12th International Congress of Refrigeration, Madrid (1967)

I, pp. 173-185. [3] D.R. EDWARDS and R.J. SLAUGHTER, Electrical Times, 152, 166 (1967); also D.N.H.

CAIRNS et al, Commission I, London, Annex 1969-1 Bull. I.I.R., pp. 145-162. [4] D.K. FINNEMORE, T.F. STROMBERG and C.A. SWENSON, Phys., 149, 231, (1966). [5] S.L. WIPF, Proc. Brookhaven. Summer Study (1968). [6] R.J. CORRUCCINI, N.B.S. Tech. Note 218 (1964). [7] R. HAMPSHIRE, J. SUTTON and M.T. TAYLOR, Commission I, London, Annex 1969-1

Bull. I.I.R., pp. 251-257. [8] R.D. CUMMINGS and W.N. LATHAM, J.A.P., 36, 2971 (1965). [9] T. Doi et al, Cryogenics, 8, 290 (1968).

[10] RS. SWARTZ and C.P. BEAN, J.A.P., 39, 4991 (1968). [11] A.E1. BINDARI and M.M. LITVAK, J.A.P., 2913 (1963). [12] S.L. WIPF and M.S. LUBELL, Physics Letters, 16, 103 (1965). ]13] B.J. MADDOCK and G.B. JAMES, Proc. I.E.E., 115, 543 (1968).

DISCUSSION

S.H. MINNICH (U.S.A.) — The assumption that the resistive losses are all absorbed in the metal of the conductor seems unnecessarily pessimistic. The specific heat of liquid helium is such that the circuit could stay in the resistive state for several seconds with a temperature rise in the helium limited to 1 °K. The liquid should be considered in such analysis to avoid over-pessimism.

M.T. TAYLOR — This pessimistic assumption is reasonable for a protection analysis because the magnitude of the heat inleak due to a cryogenic envelope failure cannot be known and must be absorbed by the refrigerant.

P. BARRET (France) — What time would it take to restore normal operating conditions of the cable after a fault lasting 1 sec. and rising to 7 times the nominal current ?

M. T. TAYLOR — None, the cable will be designed so that such a fault will only increase the temperature of the refrigerant by a maximum of 1 °C, so that the cable will still be fully operational.

125

A "SUPERCONDUCTING LINK" FOR LABORATORY TESTS ON CONDUCTORS FOR SUPERCONDUCTING CABLES

E.C. ROGERS, E.C. CAVE and R. GRIGSBY Central Research and Engineering Division

British Insulated Callender's Cables Ltd., London {United Kingdom)

Un « raccordement supraconducteur » pour essais en laboratoire de cables supracon-ducteurs

RESUME : Le « raccordement supraconducteur » a ete construit au cours d'une etude de la praticabilite des systemes supraconducteurs pour le transport de courant alternatif et cela en repondant aux exigences suivantes :

1. Fournir un banc d' essais pour la mesure sur des systemes conducteurs de dimensions normales des pertes en courant alternatif et des fuites thermiques;

2. Fournir Vexperience de construction pratique et a"utilisation d'un systeme cryogenique a echelle reduite fonctionnant a la temperature de Vhelium liquide;

3. Fournir la demonstration en laboratoire de la capacite d'un systeme supraconducteur pour transporter le courant alternatif.

Le raccordement comprend une paire coaxiale de conducteurs de niobium de diametres 3,5 et 5,5 cm, epais de 0,1 mm et longs de 2,7 m. Ils se terminent a chaque bout sur des transfor-mateurs de courant a rapports d'enroulement de 400:1. Transformateurs et conducteurs sont refroidis a 4,2 °K a Vaide de conduits remplis a"helium liquide.

On decrit la construction du raccordement et Vexperience de son fonctionnement jusqu'd present. La capacite de transport de courant est maintenant limitee a 2 000 A par les methodes de raccordement des feuilles minces de niobium et le systeme doit etre modifie pour augmenter sa capacite de transport de courant a 10 000 A environ et pour ameliorer la methode de mesure des pertes en courant alternatif.

INTRODUCTION

The Superconducting Link was constructed as part of a study of the feasibility of a.c. superconducting power transmission systems, and was designed to meet the following requirements:

1. To provide a test-bed for measurements of a.c. losses and thermal inleak on possible conductor systems of full-scale dimensions;

2. To provide practical design and operating experience of a pilot scale cryogenic system operating at liquid helium temperature;

3. To provide a laboratory demonstration of the current-carrying capability of an a.c. superconducting system.

An overall view of the Link is given in figure 1. Basically, it consists of a single-phase conductor system, in the form of a coaxial pair of niobium conductors of 3.5 and 5.5 cm diameter, and of 2.7 m length. The conductors are terminated at each end by superconducting current transformers of 400:1 turns ratio. The trans­formers were designed to permit currents of up to 5,000 A to be passed through the conductors under virtually short-circuit conditions. Both the transformers and the conductors are cooled to 4.2°K by means of ducts filled with liquid helium. The helium system is thermally insulated by a vacuum space and by a radiation shield maintained at liquid nitrogen temperature, which is in turn insulated from the vacuum casing by a second vacuum space. The vertical manifolds at each end of the Link accommodate the filling and boil-off lines to the cooling ducts, and all the electrical leads.

127

Fig. 1 — Overall view of the Superconducting Link

THE CRYOGENIC SYSTEM

In figures 2 and 3 schematic diagrams are given of a transverse section through the conductor system and of a longitudinal section through one of the terminations (in the latter diagram, the liquid nitrogen shield and the vacuum casing have been omitted for clarity).

The helium system consists of three separate liquid helium containers: 1. The Inner Duct, of 0.75 1 capacity, which cools the inner conductor; 2. The Outer Duct, of 2.5 1 capacity, which cools the outer conductor; 3. The Transformer Duct, of 5.01 capacity, which cools the transformers. Each

transformer is enclosed in a cylindrical container, and the two containers are interconnected by a central tube.

The entire helium system is supported within the liquid nitrogen duct by load-bearing spacers located one at each end of the outer duct. These spacers each consist of two concentric nylon rings, separated by a 1 cm gap, and bound together by three equally-spaced radial bindings of 50 close-wound turns of 0.025 cm diameter nylon monofilament. By using nylon in tension in this way, the 20 kg weight of the helium system can be supported by a total cross-section of nylon of only 0.30 cm2, so mini­mising the heat inleak caused by the spacers. This is about 22 mW.

The liquid nitrogen duct is supported within the vacuum casing by three equally-spaced Tufnol blocks at each end. The transformer containers are surrounded by

128

copper radiation shields, whih are bolted to the liquid nitrogen duct, so that the helium system is fully shielded from room temperature radiation, except for that

LIQUID NITG03EN

(inches)

Fig. 2 — Transverse cross-section of the Superconducting Link.

TBANSFQgMEg DUCT

HELIUM PIU./&OIL-QPF

LINES [7777] UpUlO/ GA5SOU5 HELIUM

OUTEC DUCT

TQgQiPAL TgAN5PQgM£g VrYCAST' COSE SEALS CONDUCTQgS

Fig. 3 — Longitudinal section of termination.

entering through the holes in the end shields, which permit the space between the helium system and the liquid nitrogen shield to be evacuated, and through which the helium filling and boil-off lines pass.

129

The filling lines, consisting of \" (0.634 cm) o.d., 3/16" (0.476 cm) i.d. nylon tube are fitted at one end to each of the three liquid helium ducts, and at the opposite ends of the inner and outer ducts, identical tubes serve as gas boil-off lines. The boil-off lines from each transformer container consist of two f" (1.59 cm) o.d., 15/32" (1.19 cm) i.d. nylon tubes; these also accommodate the transformer leads, which are thus gas-cooled. To reduce heat inleak due to absorption of room temperature radiation, the filling lines are each lapped with aluminised mylar tape. At each end of the nylon tubes, demountable joints are made by forcing them over close-fitting brass stub tubes lubricated with silicone vacuum grease. These joints have been found to withstand repeated thermal cycling, and to remain vacuum-tight when liquid helium is transferred.

As shown in figure 3, the filling and boil-off lines to the inner and outer ducts pass through brass U-tubes let into the transformer containers, which therefore intercept the heat inleak down these lines. The purpose of this arrangement was to thermally isolate the inner and outer ducts, so that the conductor losses and the thermal inleak could be measured calorimetrically by monitoring the rate of helium boil-off from these ducts.

In order to thermally isolate the inner and outer ducts, it was originally intended that the space between these ducts, and also the space between the inner duct and the central tube linking the transformer containers, should form part of the vacuum space. This meant, however, that a large number of internal joints, which were inaccessible except by complete dismantling of the helium system, were required to be vacuum tight. Particular difficulty was experienced with the joints where the niobium conductors pass through the end wall of the transformer containers (see fig. 3), which are sealed with Stycast G.T. 2850 resin. After repeated attempts a satisfactory vacuum (about 10"5 torr) was attained at room temperature, but these joints invariably opened up when the system was cooled. The helium system was therefore modified by the fitting of the "bellows seals", shown in figure 3. The vacuum space was thereby divided into two regions:

1. The outer space, between the helium system and the liquid nitrogen shield, and between the shield and the vacuum casing, which provides thermal isolation and so must be maintained at high vacuum (10~5 torr or better), but in which all joints are accessible;

2. The inner space, within the helium system. Loss of vacuum in this space is less serious in that it does not prevent cool-down, though it does result in thermal coupling between the ducts, and so prevents separate measurement of the a.c. losses in the two conductors. Provision was made for this space to be separately pumped via a \" (0.634 cm) o.d., 3/16" (0.476 cm) i.d. nylon tube.

The "bellows seals" successfully overcame the problem of vacuum leaks, and a vacuum of 5 x 10"6 torr or better can now be maintained in the outer vacuum space at all stages of cool-down.

CONDUCTORS AND TRANSFORMERS

The conductors are of as-rolled niobium foil, of 0.010 cm thickness. The inner conductor consists of a single strip, wrapped longitudinally around the inner duct, with a spot-welded seam. The outer conductor consists of four overlapping strips, laid longitudinally within the outer duct, and held in place by the Tufnol spacers that separate the two conductors, and which incorporate compression rings of nylon

130

tube. Recent a.c. loss measurements made at B.I.C.C. * have shown that the conduc­tor thickness is in fact greater than is necessary, and that niobium sheet of only 0.0025 cm thickness can carry current densities of up to 500 A per cm width (corres­ponding to 5,500 A in the conductors) with acceptable a.c. losses.

Fig. 4 — Termination, partly dismantled to show superconducting transformer.

A toroidal search coil, wound on a Tufnol former, is threaded over the inner conductor, and the voltage developed in the coil, which is measured with a valve-voltmeter, gives a direct indication of the current in the conductors.

The terminating transformers are also toroidal, with cores wound form grain-oriented silicon iron tape of 0.013" (0.033 cm) thickness. The turns ratio is 400:1. The primary winding consists of 400 turns of enamelled 0.010" (0.025 cm) diameter as-drawn niobium wire, while the secondary winding consists of a single turn, made up of 16 strips, each 0.5 cm wide, of 0.025 cm thick as-rolled niobium sheet, all connected in parallel. The leads to the primary windings, which pass through the helium boil-off tubes, each consists of a bundle of fifty 28 S.W.G. enamelled copper wires. This sub-division of the leads results in good heat exchange with the effluent gas. The joints between the secondary strips and the conductors (see fig. 3) were made by first nickel-plating the niobium, and then soldering with Wood's metal. Figure 4 shows one end of the Link partially dismantled to reveal the transformer.

* Measurements made by R. Grigsby as part of a joint study with the Ministry of Tech­nology.

131

A protection system is provided which automatically disconnects power from the Link if a superconducting-to-normal transition occurs in either a conductor or a transformer. A schematic diagram of the system is shown in figure 5. A small voltage transformer is connection across the primary winding of each of the terminating transformers, and the added output voltages are balanced against the amplified search coil voltage (which is proportional to the current in the conductors). Any change of impedance in the superconducting circuit, either in a transformer or in a conductor, disturbs this balance, and the resulting out-of-balance voltage is used to fire a thyratron, which in turn operates a cut-out, and disconnects the power supply, in an overall time of about 40 ms.

COOL-DOWN PROCEDURE

An automatic topping-up system is used to maintain the liquid level in the liquid nitrogen duct. The level sensors consist of resistors wound with copper wire. In normal operation, the boil-off rate from the duct is about 0.5 1/hr, corresponding to a heat inleak of 20 W, or 1.45 mW/cm2.

The helium system contains several Wood's metal joints which are mechanically weak, and so slow and uniform cooling is desirable. Cooling is therefore commenced by blowing cold nitrogen gas through the ducts, the gas being generated by energising a heater immersed in a liquid nitrogen container. This stage of cool-down is in progress in figure 1. The temperatures of the gas emerging from the three ducts are monitored with thermocouples, and the flow rates are adjusted to give uniform cooling. After the ducts have been cooled to below 100°K they are purged with pure helium gas which is pre-cooled by passing it through a heat-exchanger coil immersed in liquid nitrogen. The ducts are then filled in turn with liquid helium, starting with the transformer duct.

Initially, considerable difficulty was experienced in filling the ducts with liquid helium, as a result of a large apparent heat inleak, which was present only when helium was being transferred. It was then realised that this was caused by the thermal coupling between the ducts, resulting from the poor vacuum in the inner vacuum space (see fig. 3). When liquid helium was transferred into one duct, the other two ducts were also cooled, by conduction, and helium gas, at room temperature, was drawn into these ducts from the gas recovery system. The high heat capacity of this gas was responsible for the large apparent heat inleak. Heat exchangers were fitted to the inlet ports, so that the gas drawn into the ducts entered at liquid nitrogen temperature, and the difficulty in transferring liquid helium was then overcome.

PERFORMANCE

When the system was first operated, the critical current was 2,080 A, corresponding to a volume density in the inner conductor of 18,900 A/cm2. The transition was completely reversible and stable. However, this current is considerably less than the critical currents of either the transformers, which are known, from separate tests, to be capable of supplying 5,000 A, or of the conductors, which should have a critical current of about 13,000 A. The relatively low measured critical current is limited by the joints between the conductors and the transformers, which were made by nickel-plating the niobium, and soldering with Wood's metal. It seems likely that repeated thermal cycling, during the early unsuccessful cool-down attempts, has caused some deterioration of these joints. It is also possible that they are inadequately cooled, since they are not directly immersed in liquid helium, but are cooled by conduction through the brass transformer containers (see fig. 3).

132

The total heat inleak to the helium system is approximately 3 W. This is about an order of magnitude greater than the design value, and the excess is thought to be caused by room temperature radiation reaching the helium system through the holes in the end-plates of the radiation shield, despite the fitting of baffles to prevent this. Because of the thermal coupling between the helium ducts, resulting from the poor vacuum in the inner vacuum space, measurement of the a.c. losses in the con­ductors has not yet been possible.

50H* I

itOO-i

SUPERCONDUCTING. CIRCUIT AT U *

cou.

VMM yoirriFrfg

■AW* —

UNIT

EMS

U

LOAD

Fig. 5 — Schematic diagram of protection circuit.

Separate tests have shown that a more robust joint in niobium strip can be made by spot-welding, and, with the joint fully immersed in liquid helium, critical currents of up to 98% of that of the unjointed material have been measured. It is now proposed to rebuild the conductor system, using this type of joint, and also to improve the transformers, so that currents of up to about 10,000 A can be attained. This corres­ponds to a current density in the inner conductor (assuming a thickness of 0.00254 cm) of 357,000 A/cm2. For comparison, the current density in the 3 in2 copper conductor of a 400 kV conventional oil-filled cable is 83 A/cm2. The possibility of using an electrical method of loss measurement is being examined, since this would eliminate the need for sub-divisiorr of the liquid helium system, and so would greatly simplify the construction.

ACKNOWLEDGEMENTS

Acknowledgements are due to Dr. A.L. Williams, Director of Research & Engi­neering, B.I.C.C., for permission to publish this paper. The authors also wish to thank numerous colleagues, both past and present, who have rendered valuable assistance with the work.

133

A COMPARISON OF SUPERCONDUCTING AND CRYO-RESISTIVE POWER CABLES

E.C. ROGERS Central Research and Engineering Division

British Insulated Callender's Cables Ltd., London (United Kingdom)

Comparaison des lignes electriques supraconductrices et cryo-resistantes

RESUME : On etudie Vamelioration du transfert de courant et la diminution des pertes par refroidissement interieur sous pression des cables conducteurs de temperatures superieures a la temperature ambiante a 4 °K, et Von examine le cas de Vutilisation des cables cryogeniques. On cite lesfrais d'achat et d'exploitation pour lignes supraconductrices a conducteurs de niobium refroidis par helium liquide ou supracritique et pour lignes cryo-resistantes a conducteurs d*aluminium refroidis par hydrogene ou azote liquide.

On conclut que les lignes supraconductricespourraient faire concurrence aux lignes classiques pour des puissances de 1500 MVA et davantage, et que les lignes cryo-resistantes ne supportent la comparaison a aucun niveau de courant ni avec les lignes classiques ni avec les lignes supra­conductrices.

The need to transmit progressively higher levels of power, and difficulty in obtaining wayleave for overhead lines, is resulting in an increasing demand for the underground-ing of major power routes. However, even at the power levels required at the present time, the heat generated by the cables, which is typically 165 kW/km for a 275 kV 750 MVA single circuit, can be dissipated naturally only in exceptionally favourable circumstances. Recourse must therefore be made to multiple circuits, laid in widely spaced trenches, with forced cooling of the soil by water flowing in buried pipes.

As power levels increase still further, it will be necessary to cool the conductors internally by the forced circulation of oil or possibly water. A possible extension of this approach is to refrigerate the conductors so that advantage is taken of the increased conductivity of metals at low temperature. In the limit, if the temperature is reduced to a few degrees above absolute zero, use may be made of the phenomenon of super­conductivity.

The purpose of refrigerating the conductors of a power cable is therefore primarily to increase the current rating per circuit and to make it independent of the thermal characteristics of the soil in which the cable is installed. In addition, it is possible that the transmission efficiency may be improved, and, if so, the reduced running costs would provide a further incentive.

Clearly, the choice of the most suitable operating temperature for a cryo-cable is a crucial problem and the purpose of this paper is to compare the relative merits of cryo-resistive cables, in which aluminium conductors are cooled with either liquid nitrogen, or with liquid hydrogen, with superconducting cables, in which niobium conductors are cooled to 4°K with liquid helium, and also to determine whether any one of these possible types of cryo-cable is likely to be economically competitive with conventional high-voltage, oil-filled cables. The ultimate choice between these possible systems must depend on the comparison of the capital and running costs of fully engineered designs. However, in the sections that follow, it will be shown that a useful preliminary indication of their relative merits can be obtained by considering a conductor of given size, and comparing the current ratings and transmission losses for each of the various coolants. A comparison will then be made of the capital and running costs of possible cable designs.

135

COMPARISON OF CURRENT RATINGS

A tubular conductor is considered, of overall diameter D cm, and of thickness determined by the operating temperature. With cryo-resistive cables, the current rating of the conductor is limited by the coolant flow rate, and this is in turn limited by the maximum permissible coolant pressure, which, for an assumed spacing between cooling stations of 10 km, has been taken as 20 atm (2.07 MN/m2). Single-phase cooling is assumed, and, for turbulent flow, the coolant velocity Y is given by:

l0.092Lao'V'2J

where AP is the pressure drop, L is the length, d is the internal diameter, and a and T| are the coolant density and viscosity at the mean operating temperature. The current rating, I, is then given by I2 = %d2 oc VCy(T0 -Tf)/4R, where R is the 50 Hz resistance, Cv is the mean specific heat of the coolant, and Tt and T 0 are the inlet and outlet temperatures. For a conductor of given thickness, and for a given coolant, I is proportional to d1'8. All assumed coolant parameters are listed in table I. With the liquid gases, a static pressure is needed to suppress boiling at the outlet temperature. The assumed outlet temperatures and pressures have been chosen to give the maximum rate of heat extraction, and therefore the maximum current rating.

Table I ASSUMED COOLANT PARAMETERS

Coolant

Inlet temperature Outlet temperature

Inlet pressure (atm) Outlet pressure (atm)

Mean density (g/cm3) Mean viscosity (cP) Mean specific heat (J/g°K)

Water

25 °C 85 °C

20 0

0.99 0.50 4.2

Oil

25 °C 85 °C

20 0

0.88 4.0 1.8

Liquid Nitrogen

77 °K 106 °K

20 10

0.74 0.10 2.1

Liquid Hydrogen

20 °K 27.5 °K

20 4

0.071 0.0125 6.3

With a plain aluminium tube, the resistance/unit length is a minimum when the wall thickness is nX/2 and is given by 0.92 p/X, where p is the resistivity and X is the 50 Hz skin depth. The conductor resistance cannot be further reduced by increasing the thickness beyond this limit (although a greater thickness may be required for mechanical reasons). However, effective use can be made of a greater cross-section of metal if the conductor is stranded, and the layers are transposed, so that each strand is subject to the same flux linkage. The strands then all have the same 50 Hz impedance, and the current divides equally between them. Eddy current losses occur within the strands, and decrease with decreasing strand diameter, though a lower limit to the strand size that can be used is clearly set by practicability. For a given strand size, there is an optimum number of layers which gives the minimum effective resistance per unit length. An expression for the effective resistance of a stranded conductor is given by Wilkinson [1].

If the thickness of a tubular conductor is small compared with its diameter, its resistance can be expressed in the form of ohms per square, i.e. per unit length and per unit peripheral width. Values of resistance per unit square for an aluminium conductor cooled with either oil, water, liquid nitrogen or liquid hydrogen, are com-

136

pared in table II. The assumed conductor thicknesses are all optimum values, (i.e. those giving minimum resistance), with the exception of the liquid nitrogen cooled stranded conductor. With this conductor, the resistance could be further reduced by increasing the number of strands, but the advantage would be largely offset by the greater diameter and cost. It will be noted that, because of the skin effect and eddy current losses, the gain in reduced 50 Hz resistance that can be obtained by refrigerating an aluminium conductor is considerably less than might be supposed from the d.c. resistivity ratio.

Table II ASSUMED CONDUCTOR PARAMETERS

Conductor Material

Coolant

Mean Temperature

Resistivity (Q-cm)

Construction

Optimum Thickness (cm)

Resistance per square (oil)

Oil/Water

55 °C

3.0 x 10-6

Plain Tube

1.93

2.25

Aluminium

Liquid Nitrogen

91 °K

3.2 x 10-7

Plain 0.159 cm Tube Diameter

Strands

0.633 1.90*

0.73 0.25

Liquid Hydrogen

24 °K

1.5 x 10~9

Plain 0.025 cm Tube Diameter

Strands

0.043 0.10

0.050 0.031

Niobium

Liquid Helium

5°K

— Plain Tube

0.0025

3.4 x 10~5

* Not optimum value. See text.

For a superconducting cable cooled with liquid helium, niobium is the preferred conductor material. The coolant pressure drop in a superconducting cable is likely to be small (typically 0.2 atm, or 0.02 MN/m2, with cooling stations spaced 10 km apart), and is not a factor limiting the current rating, as it is with cryo-resistive cables. If it is assumed that fault currents can be limited by a terminal device, (e.g.

OVfRflll Dl f lMfTfRM

Fig. 1 — Comparison of conductor current ratings

137

a "resonant link" [2]), then the current rating is determined by the acceptable level of a.c. losses. Measurements made at BICC (to be published) have shown that a niobium conductor of 0.0025 cm thickness can carry a peripheral current density of 500 A/cm, with an a.c. loss, at a mean operating temperature of 5°K, of 8.5 uW/cm2. This corresponds to an effective resistance per square of only 3.4 x 10~11 Q, which is three orders of magnitude less than the value for a stranded aluminium conductor cooled with liquid hydrogen (see table II). The niobium conductor would be supported on an aluminium backing tube, and would have a current rating of 500 TtD amps and an a.c. loss of 2.67 D W/km, where D is the outer diameter in cm.

In figure 1, current ratings are given for aluminium conductors cooled with either oil, water, liquid nitrogen or liquid hydrogen, and also for a superconducting niobium conductor cooled with liquid helium. The constants assumed are those given in tables I and II. For conductor diameters of up to 14 cm, superconducting niobium gives the highest current rating. Above this limit, the rating for a liquid hydrogen cooled stranded aluminium conductor is higher. Liquid nitrogen cooling is only slightly better than water cooling, though both show a considerable gain compared with oil. With liquid nitrogen cooling, the lower resistance of a stranded conductor, as compared with a plain tube, is largely offset by its greater diameter.

REFRIGERATION COSTS

It would thus appear that, at sufficiently high current ratings, a liquid hydrogen cooled aluminium conductor could compare favourably with a liquid helium cooled niobium conductor. However, it remains to consider the price that has to be paid for these high current ratings in terms of refrigeration capital and running costs. In order to estimate these costs, the following assumptions have been made: (1) A 3-phase single circuit is considered, with refrigerators installed at 10 km intervals.

Refrigerators are duplicated to safeguard against failure. (2) Refrigerator capital and running costs are based on figures given by Kurti [3].

The capital cost can be expressed in the form of C = C0 Q" where Q is the refriger­ation capacity and C0 and n are constants. If C is in £x 103, and Q is in kW, then, for 4.5, 20 and 80°K refrigeration, values of C0 are respectively 120, 29.0 and 2.5, whilst corresponding values of n are 0.52, 0.63 and 0.68. To calculate the running costs, the refrigerator power inputs required to extract 1 W of heat at 4.5, 20 and 80°K are assumed to be 500, 33 and 6 W respectively. The cost of electricity is taken as 0.75 d/kWh.

(3) With stranded aluminium conductors cooled with liquid nitrogen or liquid hydrogen, the refrigerator load consists of the conductor losses only, and the heat inleak and dielectric losses are by comparison negligible. The running cost is taken as theftcost of the power required to drive the refrigerator.

(4) With superconducting niobium conductors cooled with liquid helium, the heat inleak and dielectric losses are not negligible, and the refrigerator load is taken as three times the total conductor losses (this is approximately correct for the cable design considered in the next section). The running cost is taken as the cost of the power required to drive the refrigerator, plus the cost of helium gas make-up, taken as 100 s.c.f. per hour per kW capacity [4], at 6 d/s.c.f.

In figure 2, capital and 10 year running costs of refrigeration are given in the form of £ per A per 10 km length for stranded aluminium conductors (see table II) cooled with liquid nitrogen or liquid hydrogen, and for superconducting niobium conductors cooled with liquid helium. The 10 year running cost is a convenient way of "capi-

138

talising" the losses. These figures show that, despite the high current rating of liquid hydrogen cooled aluminium conductors, the capital and running costs are in fact higher than for liquid nitrogen cooled conductors, and that both types of cryoresistive conductor compare very unfavourably with superconducting niobium conductors.

Q i

uxH

CAPITAL C06T» 10 yrAK fluNNJNc; coZT■-

Q Nfc KxiJud-- c~4«J -Jk 1H> © Slv-~d«4 HI - " IN © • ■ ■ ■ ' ■ » *

J&.

£) JO

»o ao 30 ko

CURRENT KflTlNG M )

Fig. 2 — Refrigeration costs of cryo-conductors.

The refrigeration costs of cryo-resistive conductors can be reduced by reducing the current rating, but this increases the conductor size for a given rating, and so increases the capital cost of the cable itself. For a given cable geometry, there will be an optimum current density, giving the minimum overall capital and running costs. In order to compare overall costs of cryo-resistive and superconducting cables, a specific cable geometry will now be considered.

COMPARISON OF OVERALL CABLE COSTS

Cryo-cables operating at 132 kV are considered, with the three phase conductors arranged in trefoil, and contained within the same cryogenic envelope. The cross-section of the cryo-resistive version is shown in figure 3 (a), and of the superconducting version in figures 3 (b). The basic differences, apart from the conductor material and coolants, are: (1) With the superconducting design, a liquid nitrogen cooled radiation shield is

needed to reduce heat inleak to the conductor system, but is not needed with the cryo-resistive design.

(2) With the superconducting design, the conductors consist ot three coaxial pairs, whereas with the cryo-resistive design the outer conductors are omitted.

In the latter design a uniform current distribution is ensured by the spiralling and transposition of the strands, and the outer coolant duct serves as a common screen. The losses in this duct are taken as equal to the losses in one conductor. With both designs, the dielectric consists of lapped polypropylene tapes impregnated with the coolant. Assumed dielectric properties are given in table III. The conductor con-

139

Table III ASSUMED PROPERTIES OF LAPPED POLYPROPYLENE DIELECTRIC

Coolant

Working stress (kV/cm)

Dielectric constant

Tan 8

Liquid Nitrogen

150

2.35

4 x 10~5

Liquid Hydrogen

150

2.35

2 x 10~5

Liquid Helium

100

2.35

2 x 10-5

HfAT SHIELD

5UPf/?IN5Uifl7lON I c THICK

O00JL5 OH hk> CONDUCTORS BBCKVfD

WITH AlUMlNlUn

JXZUURIC

COOLflWT

CORROSiON PROTECT!D

(cOSUPfRCONPUCTINk

( 5 « THICK)

iki CORROSION-PffOttCTH)

5T£I i Pipr CRY0-ftfS!STlV£

Fig. 3 — Schematic cross-sections of cryo-cables.

Table IV ASSUMED UNIT COSTS

Item Cost

Niobium Aluminium conductors and ducts (manufactured):

High purity (p20°K = 1.5 x lO-^Q-cm) Commercial (99.6%)

Polypropylene dielectric Superinsulation, per m2 per cm thickness Corrosion-protected steel tube, o.d. D cm Liquid nitrogen Liquid hydrogen Liquid helium

£ 50/kg

£0.512/kg £0.315/kg £ 0.581/kg £ 3 £0 .16D/m £ 0.02/1 £ 0.50/1 £ 1.25/1

Fixed Costs {per 10 km length) Contraction joints Factory assembly and inspection Buildings and plant Open country installation Stop joints Instrumentation

£ x 103

50 100 60

215 35 10

structions are as given in table II , and the refrigeration capital and running costs were calculated as previously described. All additional assumed unit costs are listed in table IV.

Capital and 10 year running costs have been calculated for power levels of 1500, 3000 and 4500 MVA. With superconducting cables, a fixed peripheral current density of 500 A/cm is assumed, and this determines the conductor sizes. With cryo-resistive

140

Table V

PERCENTAGE BREAKDOWN OF CAPITAL COSTS OF 3000 MVA 132 kV CRYO-CABLES

Cable Type

Coolant

Conductors and duct Dielectric Cryogenic envelope LHe filling LH2 filling LN filling Duplicate refrigerators Contraction joints Factory assembly and inspection Buildings and plant Open country installation Stop joints Instrumentation

Costs per MVA/km (£) Capital cost 10 year running cost

Total

Cry o-Resistive

Liquid Nitrogen

% 29

6 18

— —

2 18 3 6 3

12 2 1

122 73

195

Liquid Hydrogen

% 10 3 9

— 17

— 41

2 4 2 9 2 1

159 77

236

Super -conducting

Liquid Helium

% 17

3 17 24

— 2

19 2 4 2 8 1 1

71 37

108

cables, costs were evaluated, at each power level, for a range of conductor diameters (and therefore current densities) up to a maximum of 24 cm, and the diameter selected was that giving the lowest value of overall capital cost plus 10 year running cost. At power levels of 3000 and 4500 MVA, it was found that a lower overall cost could be obtained by the use of double circuits. Results are plotted in figure 4, and compared with corresponding values for 275 kV conventional oil-filled cables. A percentage breakdown of the costs for the 3000 MVA designs is given in table V.

CONCLUSIONS

The costs given in figure 4 and table V confirm the conclusions reached from the consideration of conductor refrigeration costs, namely, that with cryo-resistive cables, cooling with liquid nitrogen results in lower overall costs than cooling with liquid hydrogen, but that the cost of superconducting cables is substantially lower than that of either type of cryo-resistive cable, over the whole power range considered.

It must be emphasized that the cable designs considered are not necessarily the best attainable and that, because of the many assumptions involved, no great accuracy is claimed for the absolute values of cost. The primary purpose of the work has been to compare the relative merits of superconducting and cryo-resistive cables, and, with the latter, of different coolants, by considering the same basic conductor geometry, and by making mutually consistent assumptions for each. Nevertheless, the margin between the costs of conventional 275 kV oil-filled cables and of super-conducting cables would appear to be sufficiently large to justify the conclusion that the latter could be economically attractive at power levels of 1500 MVA or greater, and that these cables could provide a possible solution to the problem of transmitting the very high power levels that will be required in the not-so-distant future.

141

<4

I

AS*

aoo-

tfo-

too-

XO-

10

it)

U)

cgyo-gfsiSTivf X

U) ^075»f7o^C—

o > \ ^ ^

tf«* »rx Jt facKef i

indicate ru> oF circuit-*

i DO 4000

w

0)

3000

J ^ C O O I J ^ C (a)

LN mm INC. %1

supr^coNPticri^/c

9

i»oco

TRflHSMISSION CHPflClTV ( M V * )

Fig. 4 — Comparison of costs of cryo-cables.

ACKNOWLEDGEMENTS

Acknowledgements are due to Dr. A.L. Williams Director of Research and Engineering, British Insulated Callender's Cables Limited, for permission to publish this paper.

REFERENCES

[11 K.J.R. WILKINSON, Proc. l.E.E., 113, 9, p. 1509 (1966). 12] K. M. JONES, "The Economics of the Reliability of Electricity Supply", l.E.E. Conference

publication No. 34, (Part 2, discussions). [3j N. KURTI, New Scientist, p. 604 (7th December 1967). !4j A.R. WINTERS and W A. SNOW, Adv Cryog. Eng., 11, p. 116 (1966).

DISCUSSION

S, H. MINNICH (U.S.A.) — Have you considered dielectric losses in your helium cable analysis?

E .C ROGERS —Yes.

142

S.H. MINNICH (U.S.A.) — Your loss tangent is much lower than anything we have measured. Are they based on measured data?

E.C. ROGERS — Yes, the assumed values of dielectric loss tangent (given in Table III) are based on results reported by M. J. Chant {Cryogenics, 7,6, p. 351,1967).

S.H. MINNICH — You have assumed nitrogen shielded pipe in the helium design. Did you take piping costs into account? Our studies show piping costs to be large—more than 35% of the total. Was the cost of the shielded pipe taken higher than that of pipe for resistive cables ?

E. C. ROGERS — Differences in the piping costs of cryo-resistive and supercon­ducting cables were taken into account by separately costing the components.

J. JEANMONOD (U.S.A.) — In the comparison of costs shown in figure 4, only a single circuit is suggested for a superconducting cable up to a capacity of 4500 MVA. I would like to know if this is acceptable in view of the emphasis these days on system reliability. Would it not be preferable to use two or more circuits of smaller transmis­sion capacity ?

I would also like to offer for Mr. Rogers' comment a different philosophy on the design of cryogenic cables which we at Simplex are using. Throughout this paper is the underlying concept of cooling cables to reduce resistive losses. However, these losses contribute only a few per cent to the unit capital cost of a cable sytem expressed in £/MVA.km. Merely to attempt a reduction in this small percentage is to miss the point of developing cryogenic cables. Lower capital investment made possible by reductions in the cost of cable construction, civil engineering for trenches, etc., far outweighs any advantages to be gained from a simple reduction in resistive losses.

When these points are considered, it is possible to design a superior cryo-resistive cable to that considered in this paper. We have developed a liquid nitrogen cooled cable employing vacuum insulation which is more attractive economically than the superconducting cable presented here. The liquid nitrogen cable also has other advan­tages. In particular, it is designed to have very high short-term overload and cyclic ratings, so that under certain conditions it can even approach the cost of overhead lines. In addition, we have the possibility of deferred uprating. Refrigerators can be installed in future years as the load increases, since cooling is primarily necessary to handle I2R losses. Superconducting cables cannot take advantage of this deferred capital expenditure since, for these cables, the full refrigeration capacity has to be installed at the beginning to cope with heat inleak.

P. H. BURNIER (France) — I feel surprised not to see the different cryoresistive materials at their optimal temperature (except in the case mentioned by Dr Pastuhov with combination of electrical power transport and LNG transport). This optimal temperature may easily be seen by comparing the efficiency curve of the refrigerator to the resistivity curve of the cryoresistive material. This clearly shows that it is useless to envisage the cooling of Cu or Al to LN2 temperature. Only beryllium is worth cooling in the range 60 to 80 °K.

E. C. ROGERS — In reply to Mr. Jeanmonod, and as stated in the paper, I consider that the primary purpose of refrigerating the conductors of a power cable is to increase the current rating per circuit, and thereby to reduce unit capital costs. The reduced running costs are only a secondary advantage. At a power level of 3000 MVA, unit capital costs for the cable designs described are £l22/MVA/km for a double circuit of the liquid nitrogen cooled cryo-resistive version, and £71 and £96/MVA/km respec-

143

tively for single and double circuits of the superconducting version. These unit costs may be compared with the value of £130 MVA/km for the Simplex 345 kV liquid-nitrogen cooled cable. (S.B. Afshartous, P. Graneau, and J. Jeanmonod, I.E.E.E. Paper, No 69, TP 95-PWR). The possibilities of deferred uprating, and of high short-term and cyclic loadings apply equally to the cable design considered in the paper, and are not peculiar to the Simplex design. Our differing conclusions regarding the econ-nomic merits of liquid-nitrogen cooled cables would therefore appear to stem primarily from a differing basis of comparison, namely, with pipe-type conventional cables as installed in the U.S.A., and with direct-buried cables as installed in the U.K., rather than from a difference in design philosophy. The unit costs per MVA/km (continuous rating) of 345 kV pipe-type cables and of 275 kV direct-buried cables are £410 and £144 respectively.

In the Simplex design, the liquid nitrogen is used only as a coolant, and not as a dielectric impregnant. It may be of interest to note that, in such a design, approximately the same current rating could be attained by operating the cable at above ambient temperature, and cooling the conductors with water instead of liquid nitrogen. (See fig. 1 of present paper.)

I would therefore agree with Mr. Burnier's conclusion that there is little to be gained by cooling aluminium or copper to liquid nitrogen temperatures. However, I must again emphasize that the attainment of high current ratings and low unit costs is more important than improving the efficiency. Also, to asses the latter for a.c. transmission, the refrigerator efficiency should be compared with the a.c. resistance per square, and not with the resistivity. Unfortunately, the use of beryllium must be ruled out at the present time because of its high cost.

There are obvious difficulties in using water for the internal cooling of very-high-voltage cables, but, though different in kind, these are not necessarily any more formi­dable than those encountered in developing cryo-resistive cables. It is by comparison with advanced cable designs of this type, operating at above-ambient temperature and internally cooled with either oil, water or compressed gas, rather than by comparison with present-day conventional cables, that the merits of cryo-cables must eventually be assessed.

GENERAL DISCUSSION ON TERMINOLOGY DEFINITION

D. R. EDWARDS (U. K.) — We are already using different terminology on the two sides of the Channel. In Great Britain we use the terms " superconducting cables ", and "cryo-resistive cables". In France the terms used are "superconducting" and "hyper-conducting". I would suggest that the time is opportune for settling the terminology to be used.

N. KURTI (U. K.) — Members of this Conference might be interested to know that a small exploratory meeting to discuss cryogenic terminology is to be held early in May 1969, under the auspices of the Comite d'Etude des Termes Techniques Francais.

Industrialists and University people from several countries (France, Netherlands, Switzerland, U.K. and U.S.A.) are expected to attend.

144

ELECTRICAL CONSIDERATIONS FOR AN A.C. SUPERCONDUCTING CABLE

D.N.H. CAIRNS, R.H. MINORS, W.T. NORRIS and D.A. SWIFT Central Electricity Research Laboratories

Leatherhead {United Kingdom)

Considerations electriques pour les cables supraconducteurs a courant^alternatif

RESUME : On choisit le conducteur pour donner de faibles pertes en regime normal et en court-circuit. Une fois le materiau choisi, il existe une valeur limite du courant admissible, determinee par des considerations thermiques. Le dielectrique est choisi en principe pour sa bonne resistance a la tension de choc, bien que les pertes en courant alternatif aient une certaine importance. La contraction mecanique et la disposition des canaux de circulation du fluide cryo-genique ont une influence sur le choix du dielectrique et du conducteur, et Von examine diverses methodes de construction des cables utilisant des materiaux sur lesquels il existe un petit nombre de donnees experimentales.

Une disposition symetrique du cable est a preconiser afin de produire des champs electriques et magnetiques autant que possible homogenes. On examine les resultats obtenues en ce qui concerne la variation de tension, Vimpedance caracteristique et Vequilibre de courant entre les phases et la terre pour trois dispositions de cables, employant des materiaux susceptibles d'etre utilises pour un cable a courant alternatif. Les cables se caracterisent par une variation de tension elevee et une tres faible impedance caracteristique.

1. INTRODUCTION

With conventional high power cables, joule heating losses necessitates operation at high voltages (i.e. 275 kV and above). The limitations on power handling capacity and length, and the difficulties in manufacture and installation stems mainly from this practice. With superconducting cores it thus seems reasonable to examine high current cable designs first.

Because the optimum voltage of a superconducting cable was not known, a possible future generator voltage was chosen for study since this could allow connections to load centres without the use of transformers. It was further argued that by consider­ing a specific case, the more important design parameters, constructional problems and areas requiring research would be more quickly identified. The experience gained from such an exercise could then be used to assess and determine more favourable specifications.

This work supported by the C.E.G.B. has to date included studies of outline designs and cost estimates (Rogers and Edwards [13], Edwards and Slaughter [4], Norris and Swift [12], Swift [17]), refrigeration and coolant flow systems (Edney et al. [3], Cairns et al. [2], Swift [18]), conductor construction (Taylor [20]) and prelimi­nary dielectric experiments (Looms et al. [11], Swift [19]). The present paper com­plements this work by discussing electrical aspects in more detail. Earlier work by other groups is discussed by Edwards and Slaughter [4].

2. SPECIFICATIONS

The specification is based on that for a conventional cable. The designs are concerned with a 750 MVA 33 kV three phase cable which has

an impulse strength of 190 kV phase-to-earth and a fault rating of seven times full load current for 1 second. The cable must be operational immediately after one fault and must be capable of carrying two faults in close succession.

145

3. DESIGN CRITERIA

Because of the high cost of the cryogenic envelope, helium and its refrigerator, the heat entering the coolant and the cable cross section must be kept to a minimum. Therefore all phases will be within the same heat shield, the dielectric will be at a low temperature and coolant streams must be arranged on a 'go' and 'return' basis. The conductor circumference is determined mainly by a.c. loss considerations. During overloads, conduction may revert from the superconducting to the normal state.

The minimum cost arrangement will be a compromise solution of many conflicting requirements which cannot yet be realized because of lack of information. Therefore in order to proceed with initial designs, it was decided to assume values or limits for certain parameters. In this respect a good starting point is to decide upon a tolerable heat input to the helium W T , and from there determine the corresponding optimum working conditions. The component heat inputs are conductor loss W c , dielectric loss WD, thermal inleak WH and viscous loss W v . It is useful to express these losses in terms of conductor surface areas per metre length of cable as follows,

W c = SK1 y\fn Watts WD = SK2 8 tan 5 Watts WH = SK3 Watts

where n,Kl,K2 and K3 are constants, \|/ is the peripheral current density (A m~*), 8 is the permittivity and 8 the loss angle of the dielectric, S is the circumference of the conductors. Under optimum flow conditions (Swift 1968a) W v ~ 0.15(WV + WD + WH) and so the optimum current density becomes

[ ( K 2 s t a n 5 + K 3 ) / K 1 ( n - 1 ) ] 1/n

Therefore

Wc = °^™I and WD = 0.87 fciW-W„ n \ n )

Present data suggests K2 = 200 and K3 = 6 x 10"2 . (Rogers [14]). Thus, since S = 3I/\|/, the design criteria for the conductor and dielectric are

determined. An economically acceptable value for WT is about 0.1 Wm" 1 .

4. CONDUCTORS

A possible superconductor is niobium for which existing data for sheet of industrial finish suggest n = 6.8 and K{ = 6 x 10~3 4 (Rogers 1969), yielding v|/ = 40 A mm" \ As this current density is a factor of 2 below the critical value, development may achieve a superior performance having significant economic consequences.

With the present specification, overloads produce major problems since the resistivity of niobium (p = 5 x 10" 9 Qm) is too great for it to be allowed to carry much current. Possibilities for a parallel conductor are aluminium (p = 1.5 x 1 0 " l l Qm), copper (p = 4x 1 0 " x l Qm) and a high field superconductor (Jc ~ 1010 A m~2). It is important to note that normal conductors must be sited in field free regions to avoid eddy current heating. A triple sandwich of niobium—high field superconductor— copper would certainly provide an elegant solution to this problem, but manufacture may be difficult to realize.

146

A parallel normal conductor will also assist stabilization since it would be thermally adequate provided not more than 0.1 % of conductor length had reverted to normal conduction. Instrumentation could identify and locate such an event.

5. DIELECTRICS

The dielectric loss should be < 14 mW m~ 2 giving limits on s tan 5 and leakage current of 5 x 10" 5 and 1 uA m~2, with polarization the more stringent condition. To achieve compactness, an impulse stress > 15 MV m~ * is desirable. Other require­ments are: adequate mechanical strength, matched thermal contraction, and low thermal conductance (~ 2 x 10"x Wm~* K~ *).

The media being assessed are reviewed below. The first three options will also incorporate solid dielectric spacers, the designs of which will be discussed briefly later. Two points are worth emphasising. The conductor size will be appreciably greater than most experimental electrodes and deleterious area effects can occur. Also, quoted breakdown values are often the result of a short duration experiment in which conditioning values are discounted, whereas here the first breakdown is important in a time-span of years.

5.1 Vacuum. The attractions are low cost, loss and thermal conductance. Room temperature studies suggest adequate strength but variability occurs. However, low temperature working may be beneficial. Spacers will probably reduce the strength; best room temperature designs achieve only 8 MV m " 1 (Shannon et al. [15]). Highly refined mechanical polishing techniques do not seem practicable. However, glow discharge conditioning may be feasible and thin dielectric coatings (e.g. 1 0 " 4 m epoxy) have proved successful with rough finishes (Jedynak [9]).

The degree of vacuum will be a compromise between the density effect (Germain and Rohrback [6]) and ionization restrictions imposed by the magnetron effect. Present estimates suggest p ~ 1 0 " 4 N m " 2 .

5.2 Helium. This achieves an efficient use of space. Unfortunately the breakdown strength is questionable although, as yet, there is insufficient data to make fair assess­ments.

Data on the liquid relate only to atmospheric pressure and small gaps. Its weak­ness is thought to be due to the formation of bubbles; hot spots, stress points and impurities are thus important. A guide to magnitudes is that nucleate boiling starts at 2 0 W m ~ 2 , field enhancements of 10-20 occur with asperities of 1 0 " 7 m (Alpert [1]), and particle diameters > 2 x 10" 8 m may (Krasucki [10]) produce break­down at these projections with average fields of 15 M V m " 1 . With well prepared electrodes and good purification, an a.c. stress of 30 MV m " 1 has been achieved on a 1 mm gap (Goldschwartz and Blaisse [7]). However, in general, results are variable and extrapolation to larger gaps is difficult. The theory of bubble formation predicts that a coolant pressure of 0.4 MN m " 2 would improve the strength but danger may occur near the critical temperature.

With gaseous helium, there is no information at 4°K and so room temperature data is taken as a guide. Assuming number density to be the important parameter, the equivalent pressure is a few hundred atmospheres. D.C. field strengths of 20 MV m " 1 have been measured (Trump [21]) at 8 M N m ~ 2 (gap 10 mm); with other compressed gases, impulse values are 20% greater than d.c. strengths (Howard [8]) but spacers can cause an equal reduction (Skipper & McNeall [16]).

5.3 Solidified Gases. Since the strengths of liquid H 2 , N 2 and A are significantly greater than He, it is pertinent to consider their usefulness when solidified. Such solids are soft around the melting points but are hard and brittle at low temperatures.

147

The only electrical data so far gives a d.c. strength for argon of 20 MV m~ 1 on a 0.5 mm gap (Gallagher [5]); nitrogen and hydrogen could yield higher values. With the former, danger lies in possible cracking during the phase change on cooling. An interesting possibility is a nitrogen-argon alloy (~ 60% N2) which avoids this problem.

5.4 Conventional Materials. In general they are too lossy (e.g. 8 tan 8: paper 10" 3, epoxy resin 2x 10"4 , glasses 5 x 10"3), but some polymers (e.g. polyethylene, poly­propylene, p.t.f.e.) appear possible (8 tan 5 10"5). The problem here is that they contract by 2-3% on cooldown compared with 0.3% for relevant metals, but loading with low contraction fillers may help. However, even if matched conditions could be produced without a significant increase in loss, axial contraction complicates designs.

6. ASSEMBLIES

To avoid stress and current concentrations, designs are based on a symmetrical arrangement of tubes. Three basic assemblies are shown in figures 1, 2 and 3. The maximum peripheral current density is taken as 40 A m m - 1 and the maximum impulse stress as 15 MV m ~1 and 25 MV m ~1 for vacuum and helium respectively.

H E L I U M CO

N I O B I U M NEUTRAL CONDUCTOR

N I O B I U M PHASE CONDUCTOR

Fig. 1 — Twin tube conductor assembly.

V A C U U M

H E L I U M CO

N I O B I U M PHASE CONDUCTOR

N I O B I U M N E U T R A L CONDUCTOR

Fig. 2 — Multi-tube conductor assembly.

148

The twin-tube arrangement is probably the simplest to manufacture and install but will be of larger overall diameter than the other versions. Each phase contains a pair of concentric tubes. The outer tube is the neutral return and currents flow such that no magnetic field exists outside each tubular assembly. The superconductor is on the outer side of the inner tube and inner side of the neutral.

Fig. 3 — All coaxial conductor assembly.

The multi-tube version is aimed at achieving compactness by replacing the indi­vidual neutrals by a common outer tube. A triangular lattice array of four conductors per phase attempts to achieve a balanced magnetic field system of minimum cross-section.

N E U T R A L i c

E B= l7 -9kV

E y - 1 7 - l kV H l y = l2 -5kA

Fig. 4 — Circuit and vector diagrams for a 40 km length of the all coaxial design.

The all-coaxial version has one superconductor for the red, two for each of the yellow and blue phases and one neutral. The screening of the normal conductor produces a current flow pattern that is essentially two phase and neutral (fig. 4) viz: the current in Yx is equal and opposite to that in the red conductor whilst that in Y2 balances that in B x . The neutral current is opposed by that in B 2 .

149

7. ELECTROMAGNETIC FORCES

At any point within the conductor assembly, the magnetic pressure is given by B2/2|i0 N m~2, B(T) is the resultant magnetic field. The resultant force pulsates at 100 Hz and is a maximum during overloads.

In coaxial structures there is a pinching force on the inner conductor and a bursting one on the outer. The maximum pressure is about 0.2 MN m~2. Since this is only 10-20% of the coolant pressure, no serious problems are foreseen. With multitube designs, there is a sideways force on the tubes of up to 5 x 1 0 4 N m _ 1 . To avoid large beam deflections, shearing stresses at the supports and deformation of the tubes, it will be necessary to have spacers at 50 mm intervals. This can cause serious design problems.

8. SPACERS

The amount of gap filled with solid will be about 1% and 10% for coaxial and multitube designs respectively. The problem of differential contraction is limited to the radial direction and stud designs are easier to match than discs. On the other hand, electrical problems may be less with the latter. Since the electrical weakness is at the triple point of spacer-electrode-main dielectric, the probability of failure will be a function of junction perimeter. Thus, for a given bearing surface, discs can provide a factor of two improvement over studs.

As breakdown voltage varies with gap length to a power less than unity, multigap spacers may be beneficial. However, the metal sheds will have to be superconducting to avoid excessive heating.

With coaxial structures, the mechanical stress at the edge of the spacer is alleviated if Young's Modulus for the solid dielectric is less than that for conductor material (~ 7 x l 0 1 0 N m " 2 ) .

The material best suited to provide these and other requirements is probably a filled epoxy resin.

9. ELECTRICAL CHARACTERISTICS

The surge impedance (i.e. V W Q of these designs is about 10 Q. As this is only a third that of conventional cables and an order of magnitude less than overhead lines or busbar sections, superconducting cables are probably overdesigned with the present impulse specification.

Because of the high currents, the inductive voltage-drop is significant. For example, with the twin-tube and multi-tube designs, the voltage regulation is about 0.3% per kilometre. This is nearly an order of magnitude greater than that of conventional cables; then it is the opposite problem of large charging currents. With the all-coaxial arrangement, the situation is more complex because of imbalance of circuit inductances which can cause considerable neutral current (i.e. ~ 5 kA). The corresponding current and voltage phase diagrams for a 0.9 p.f. load terminating a 40 km cable fed from a generator of 14% reactance is given in figure 4. An unusual feature is that the results depend upon phase sequence. With the twin-tube and multi-tube designs, the problems of unequal phase regulation and normal—load neutral current do not occur. In the all-coaxial arrangement the effect can be alleviated by using more tubes, e.g. a six conductor system having the red phase also split with the extra tube outside B 2 . This arrangement also has the added advantage of being the most compact yet devised.

150

Present indications suggest that the problems with all designs would be much alleviated by increasing the working voltage and reducing the impulse ratio.

10. CONCLUSIONS

The best available superconductor is niobium for which present indications suggest that a.c. losses are not too serious. Overload currents are troublesome in that they either produce large heat pulses or necessitate the development of a triple composite conductor. However, shunt switching techniques may overcome this problem.

The dielectrics most suitable for very low temperatures are helium and vacuum with epoxy resin spacers, but reliability needs to be improved. As impulse strengths are probably not much greater than a.c. values, surge diverters could be usefully employed.

Because of electromagnetic forces, current and voltage stress considerations, coaxial conductor assemblies are favoured. Since the surge impedance is low, the transmission coefficient is a lot below that for conventional cables; significant benefits would accrue from changing the voltage specification accordingly.

For cables of 1000 MVA and above, the voltage regulation at 33 kV can be appreciable. However, this and related problems can be alleviated by doubling the operating voltage and halving the standard impulse ratio.

ACKNOWLEDGEMENTS

This paper is published by permission of the Central Electricity Generating Board. The work is in part based upon and in part developed from work done by BICC on a design contract from the Research & Development Department of the C.E.G.B.

REFERENCES

[1] D. ALPERT, Proc. Int. Conf. High Voltages in Vacuum M.I. T. (1964). [2] D .N.H. CAIRNS, D. A. SWIFT, K. EDNEY and A. J. STEEL. Commission I, London, Annex

1969-1 Bull. I.I.R., pp. 155-162. [3] K. EDNEY, M. FOX and G. GILBERT, Cryogenics (Dec. 1967). [4] D.R. EDWARDS and R.J. SLAUGHTER, Elect. Times (3 Aug. 1967). [5] T.J. GALLAGHER, Ann. Rep. Conf. Elect. Insul. (1967). [6] C. GERMAIN and F. ROHRBACK, Proc. VI Int. Conf. Ionization Phenomena in Gases, Paris

(1963). [7] J .M. GOLDSCHWARTZ and B.S. BLAISSE, Brit. J. Appl. Phys., 17 (1966). [8] P.R. HOWARD, Proc. I.E.E. 104A (1957). [9] L. JEDYNAK, / . Appl. Phys. 35 (June 1964).

[101 Z. KRASUCKI, ERA Report No. 5157 (1966). [11] J.S.T. LOOMS, R.J. MEATS, and D.A. SWIFT Brit. J. Appl. Phys., 2, 1 (1967). [12] W.T. NORRIS and D.A. SWIFT, Elect. World (24 July 1967). [13] E.C. ROGERS, D.R. EDWARDS, Elect. Rev. (8 Sept. 1967). [14] E.C. ROGERS, 1969, Private communication (BICC report on C.E.G.B. contract). [15] J.P. SHANNON, S.F. PHILP and J.G. TRUMP, / . Vac. Sc. and Tech., 2 (1965). [16] D.J. SKIPPER, P.I. MCNEALL, Proc. I.E.E., 112 (1965). [17] D.A. SWIFT, I.I.R.-XII Int. Cong, of Refrign, Madrid (1967)1, pp. 173-185. [18] D.A. SWIFT, Cryogenics (Aug. 1968). [19] D.A. SWIFT, Vacuum, 18, No. 11 (1968). [20] M.T. TAYLOR, Commission I, London, Annex 1969-1 Bull. I.I.R., pp. 119-125. [21] J.G. TRUMP, Gas. Dis. and Elect. Supply Ind. Butterworths (1962).

151

REFRIGERATION AND CIRCULATION OF HELIUM IN SUPERCONDUCTING POWER CABLES

D.N.H. CAIRNS, D.A. SWIFT Central Electricity Research Laboratories, Leatherhead

K. E D N E Y

British Insulated Calender's Cables Ltd., London

and

A.J . STEEL

British Oxygen Cryoproducts, Morden (United Kingdom)

Le refroidissement et la circulation de Theliumdans les cables electriques supraconducteurs

RESUME : Le maintien du conducteur d'un cable a Vetat supraconducteur implique la circu­lation de tres grandes quantites d'helium. De nombreuses dispositions relatives au refroidissement et a la circulation de Vhelium ont ete proposees pour des temperatures s'etendant de 3,5 °K a 6,5 °K. Cette bande de temperature semble en effet la plus convenable en ce qui concerne le niobium, qui est le materiau le plus prometteur parmi ceux qui peuvent etre utilises pour les cables a courant alternatif. On examine la consommation d'energie et les avantages resultant de Vexploitation deplusieurs systernes. On pourrait faire circuler Vhelium en employant le com-presseur du circuit frigorifique, mais on emploie de preference une pompe a basse temperature. On examine la fonction d'une telle pompe et d'autres problemes d'exploitation tels que le refroidissement, le stockage de Vhelium et les effets des surcharges.

1. INTRODUCTION

Helium must be circulated within a superconducting power cable in order for the temperature of the conductor assembly to be lowered to and maintained at a few degrees above absolute zero. The heat entering the coolant will need to be removed by refrigerators spaced at intervals along the cable route.

Since these refrigerators are likely to feature significantly in the cost and operation of such a scheme, it is important that they be specifically designed to match the cable conditions and be reliable. Also, reasonably efficient and dependable coolant cir­culators will need to be used. This paper, therefore, examines these and other related issues with regard to a recent proposal for an a.c. cable (Edwards and Slaughter [2]).

2. CABLE REQUIREMENTS

For economic reasons, it is desirable to contain all three phases within a single compact cryogenic envelope. Under normal load conditions heat will enter the coolant from thermal inleak, dissipation in the conductor and dielectric, and viscous dissi­pation within the coolant channels. For cables of about 1000 MVA capacity, the sum of these first three components is estimated to be about 100 W k m - 1 . With optimum coolant flow conditions, the viscous loss in the cable will be around 15 W km ~1 resulting from a pressure drop along the coolant channel of less than 0.3 MN m~ 2. The refrigerator size depends upon the coolant scheme but will probably be of a few kilowatts. The time taken for the coolant to flow between refrigerators will be about a day.

With a niobium superconductor at a normal working current density of 40 kA m~ 1 , the maximum permissible temperature is about 6.7 K. Because refrigerator costs

155

increase significantly as the output temperature is reduced much below 4.5 K, a lower limit of 3.5 K is assumed.

Since transmission systems must also sustain overloads for short periods of time (e.g. 7 x F. L. C. for 1 second), the niobium may be driven normal temporarily. Some possibilities for dealing with this problem are: (a) use an aluminium backing material (Rogers and Edwards [4]), (b) use a type II superconducting backing material (Norris and Swift [3]), and (c) shunt switch on the generator side of the cable termination (Swift [5]). The pulse of heat entering the coolant with these three schemes has been estimated to be respectively 3 x 105, 3 x 103 and < 4 x 104 J m" 3 per event.

To overcome topographical and pressure drop problems, it is desirable to avoid two-phase flow. It is useful to define 3 alternative ranges of coolant condition:

(i) 'liquid' (p ~ 0.3 MN m " 2 , T = 3.5 - > 5 K ) (ii) 'supercritical' (p ~ 0.4 MN m " 2 , T = 5.2 -► 6.7 K)

(iii) 'gaseous' (/? ~ 1 MN m~2, T = 4.2 -> 6.7 K)

The pressures are determined by maximizing specific heat x density for the 'liquid' scheme (Swift [6]) and (specific heat)3 x (density)2 for the 'supercritical' version (Edney [1]).

LOAD

FIG 2 BASIC_SUPE_RC_RJ_TICAL

REFRIGERATOR

/ \ TURBINE I

^ J • TVALVF

L 0 AD w FIG I BASIC LIQUID REFRIGERATOR

156

For a 'gaseous' system, 1 MN m 2 represents a compromise between excessive mechanical loads and too large a change in density along the cable.

3. BASIC FLOW DIAGRAM

The basic refrigeration cycle depends upon a two-stage Claude cycle using gas bearing expansion turbines and liquid nitrogen precooling. The cycle employed for a 'liquid' scheme refrigerator is shown in figure 1; Figure 2 shows the low temperature end simplification possible when a 'supercritical' scheme is considered. The load arrangements vary with the cooling requirements of the cable, but the following are the basic options that have been studied (see fig. 3).

REFRIGFRATOR

CABLE J W W W W W L

(a) REFRIGERATOR FEEDS CABLE DIRECTLY. AND FLOW IS CIRCULATED BY MAIN COMPRESSOR.

REFRIGERATOR

A EXTRA COOLING STAGE (TURBINt" OR J-T VALVEi

A/VWAAA-

A / W V N A A

CABLE

- A / W W W W \ A -

HEA1 tXCHANGER .OR LIQUID TANK)

M>> REFRIGERATOR FEEDS CABLE THROUGH AN EXTRA STAGE OF COOLING

HEAT EXCHANGER OR LIQUID TANK>

- A / W W W W V ^ CABLE

SEPARATE CABLE COOLANT SYSTEM WITH ITS OWN CIRCULATOR

FJG_J THR ESTRANGEMENTS OFJHE CABLE AS THE REFRIGERATOR LOAD

(A) Direct cable refrigeration; (B) Liquid tank and/or heat exchanger with direct cable refrigeration; (C) Liquid tank and/or heat exchanger with a closed cycle cable coolant flow.

157

Since liquid nitrogen will probably be available at refrigerator sites, there is only small advantage in replacing the N 2 boiler with a third stage of turbine cooling. A turbine expansion ratio of about six, and adiabatic efficiencies in the range 65 to 75%, depending upon the actual operating conditions, have been assumed. The main system pressures for the refrigeration cycles are then in the ratio 6:1, although there may be other auxiliary pressures employed; for instance a third pass through heat exchangers may be added if subatmospheric pressure is to be applied to the liquid tank to reduce the cable inlet temperature. Two major design parameters that have been assumed are a warm end temperature difference for exchanger B of 1 K and a compressor efficiency of 50%.

4. COMPARISON OF POSSIBLE SYSTEMS AND LOAD COMBINATIONS

In table 1 the possible plants are represented by a selection of 'liquid', 'super­critical' and 'gaseous' refrigerators of 1 kW refrigeration power. For each system a single design point is presented out of the many investigated. The conclusions are however more concerned with the basic differences between the systems when the operational limitations introduced in a later section are considered.

Table 1 COMPARISON OF SELECTED 1 kW PLANTS

System

Loading Arrangement

Circulator1

Number of Turbines

Joule-Thomson Valves

Cable Inlet Temp. K

Cable Outlet Temp. K

Cable Operating Pressure MN m"2

Cable Flow kg s_ 1

Total Compressor Flow kg s"1

Total Power Consumption kW

Carnot Efficiency

Supercritical

A

MC

2

— 5.6

6.5

0.37

.09

.11

220

22.0

B

MC

3

— 5.4

6.5

0.37

.07

.12

240

21.0

C

LT.P

2

— 5.8

6.7

*t 0.6-0.3

.09

.11

220

21.0

A

MC

2

2

4.2

5.0

0.3

.27

.33

830

8.0

Liquid

B

MC

2

3

3.8

5.0

0.3 *

.19

.29

700

10.0

C

LT.P

2

1

4.6

5.0

i 0.4-0.3

.40

.16

260

24.0

Gaseous

C

LT.P

2

1

4.6

6.5

1.0

.12

.16

260

20.0

Note 1. MC = main compressor LT.P = low temperature pump

2. No pressure drop in the load has been allowed. 3. The LN2 refrigeration is not included but is represented by 5-10% increase in

compressor power.

4.1 SUPERCRITICAL

For load arrangement A (fig. 3) in which the refrigerator feeds the cable directly, the Carnot efficiency is high and, because of the high specific heat of the fluid in this region, the mass flow is small even with moderate values of AT along the cable.

158

However, while there is no economic penalty in circulating the cable flow with the main compressor there are limitations caused by the inter-relationship between the cable inlet temperature and pressure. For example, with two turbines, inlet tempera­tures below 6.0 K cannot easily be obtained with pressures > 0.4 MNm""2. With higher power consumption and increased complexity an extra stage of turbine cooling can produce a lower inlet temperature (fig. 3 arrangement B). The cable temperatures are raised slightly with arrangement C because of the heat exchanger but there are no constraints from refrigerator considerations on the cable operating pressures. If a pressure drop of < 0.1 M N m - 2 is considered there is nothing to choose between arrangements A and C, and these must be taken as the best forms for a supercritical refrigerator.

4.2 LIQUID

Unlike the supercritical refrigerators the load arrangements do have a considerable bearing on the power consumptions of different liquid refrigerators. Because of the reduced value AT along the cable and the decrease in specific heat the mass flow in the cable is increased by a factor of 2-6. Hence, with arrangement B the extra stage of cooling doubles the AT and the compressor power consumption is reduced despite the increase of 60% in the other refrigerator streams. With a directly fed cable (A or B) the optimum inlet temperature is « 3.8 K. When the refrigerator and cable flows are divorced by use of a low temperature pump the compressor power requirements are significantly reduced and arrangement C is the choice from both economic and operational considerations for a liquid system.

4.3 GASEOUS

Only a brief study has been made of gaseous refrigerators from which it appears that modifications may be made to the arrangement C liquid plant to allow the higher cable exit temperature. Although the AT is large the decrease in specific heat in this temperature range for pressures above ^ 0.6 MN m - 2 means that for 1 kW load the mass flow is likely to be « 0.12 kg.s" 1 and, as such, the gaseous system is expected to fall between the liquid and supercritical systems.

5. OPERATIONAL CONSIDERATIONS

In the previous section it was seen that the low temperature pump is an efficient method of circulating the coolant; the considerations below will show that it is necessary in order to confer the operational flexibility required by a cable.

5.1 OPTIMUM SPACING OF REFRIGERATORS

To capitalize on the benefits of using a small number of large refrigerators the spacing should be as large as possible without excessive pipe friction losses. Opti­mization of this spacing (Swift [6]) leads to values of about 20, 60 and 80 km for the liquid, gaseous and supercritical conditions respectively.

5.2 DUPLICATION AND RELIABILITY

Refrigerator reliability will have a direct bearing on the security of any cryogenic transmission system, and the provision of a suitable degree of redundancy is one of the major problems confronting the designer of any practical refrigeration scheme.

159

It has been shown (Swift [6]) that a cheaper solution than the duplication of refriger­ators at each station is to situate plants at half the optimum distance and sized such that the cable remains operational with one in every three refrigerators out of action. There would still be a requirement for the duplication of crucial components such as compressors and turbines to increase the reliability such that a component failure or maintenance does not affect the operation of the cable. The duplicated plant is also operated so as to increase the refrigeration when a neighbouring refrigerator fails, and after a cable fault. A cable off-load but with no coolant flow will remain in an operable condition for over 30 hours, and so may allow consideration of the provision of mobile cold-boxes to avoid long outages. The requirement now is for more information about the reliability of large scale helium refrigerators, so that the need for alternative stand-by arrangements can be considered.

5.3 STORAGE OF HELIUM

One of the problems that still has to be resolved is the storage of helium during filling or discharging of the cable. Since when off-load the cable will not become dangerous for about a week it is possible to consider the use of mobile stores. However, the size of such a store for a single inter-refrigerator length of 40 km is massive (e.g. « 3 x 102 m3 liquid; « 3 x 104 standard bottles compressed gas; or « 3 x 105 m3

atmospheric store). The simplest solution is probably the use of six liquid helium semi-trailers of the size now becoming available.

5.4 COOL-DOWN OF THE CABLE

The cool-down process will be complicated, not only by the need to provide gas continuously as the temperature falls, but also because of the large pressure drop down the cable to achieve a significant mass flow. The heat to be removed from 10 km of conductor is about 4 x l 0 1 0 J of which 95% is above 77 K. With plant duplication, an extra compression stage and imported nitrogen to achieve a helium flow of 0.2 kg.s" * under a pressure drop of 1 MN m~2, it is estimated that the time of cool-down to 4 K will be about 15 days.

5.5 PULSE HEATING AND ITS EFFECT UPON THE CHOICE OF SYSTEM

If fault currents are carried in resistive backing materials the heating pulse is sufficiently large that the cable outlet temperature must be restricted under normal working condition to below 5 K unless an excessive amount of helium is provided in the cable. This implies that the savings applicable to supercritical systems cannot be realized, and since the ability to boost the flow rate is limited the cable could be inoperative for more than a day after a double fault. On the other hand with the alternative solutions to this problem (e.g. type II superconductor, shunt switching) this heat pulse is small enough for all coolant systems to be possible.

6. PRESSURE DROP AND HELIUM PUMP

Because of the optimum spacing of the refrigerators and the preferred duplication plan there are wide variations in the mass flow and the pressure drop depending upon the operating condition of the cable, (i.e. see table 2). These variations make it difficult to operate the refrigeration system without the use of a low temperature pump. The operating conditions are characterised by a high head and low flow rate (e.g. 0.3 MN m - 2 and 4 x 10" 3 m3 s~ *)—and the need for efficient operation under

160

a number of different conditions. However, cavitation, one of the main difficulties normally associated with cryogen pumps is missing. The efficiency of these pumps is not critical; a 10% variation only affecting the refrigerator costs by 5% but the aim at the moment is to develop a pump with an efficiency of more than 50%. Since these pumps will be inside the cold-box they must be reliable and hence the preference for centrifugal pumps because of their simplicity and the possibility of avoiding seals at cryogenic temperatures.

Table 2 APPROXIMATE VALUES OF THE MASS FLOW AND PRESSURE DROP FOR THE DIFFERENT SYSTEMS

WITH NORMAL AND STANDBY OPERATION

Normal Operation

Standby

Mass Flow Pressure Drop

Mass Flow Pressure Drop

7.

Liquid

.25 .005

.40

.02

Gaseous

.35

.03

.80

.27

CONCLUSIONS

Supercritical

.23

.03

.45 .3

kgs"1

M N i r r 2

k g s - 1

M N i r r 2

It should be possible to design helium refrigerators to match the likely operating conditions of superconducting cables. The best method for circulating the coolant through the cable will be a low temperature pump. Such an item is not yet available but it will probably be of multistage centrifugal design.

Optimized refrigeration plant should realize better than 20% of Carnot Efficiency (i.e. - 350 W input for 1 W brought from 4 to 300°K).

A modest amount of standby plant should achieve reasonable security. On full load, the helium can remain stationary for 30 minutes, and off-load the time for cable to reach the danger limit is a few days. Thus plant repairs, the delivery of mobile units, and the arrangement of helium stores should be possible; although the latter could be a problem with long cable lengths.

Cable cooldown may be troublesome. Using on-line refrigerators with extra compressors and nitrogen, such an operation could take two to three weeks.

ACKNOWLEDGEMENTS

This paper is published by permission of the Central Electricity Generating Board, British Oxygen Cryoproducts Ltd. and British Insulated Calender's Cables Ltd.

REFERENCES

[1] K. EDNEY, (1969). To be published. [2] D.R. EDWARDS and R.J. SLAUGHTER, Electrical Times (3 Aug. 1967). [3] W.T. NORRIS and D.A. SWIFT, Elect. World (24 July 1967). [4] E.C. ROGERS and D.R. EDWARDS, Electrical Review (8 September, 1967). [5] D.A. SWIFT, I.I.R.-XIIth Int. Congress of Refrigeration, Madrid (1967), I, pp. 173-185. [6] D.A. SWIFT, Cryogenics (Aug. 1968) p. 238.

161

DISCUSSION

P.F. CHESTER (U.K.) — After two successive faults how long does it take to bring the cable back into operation and what limits the time ? Is it the circulation time for the helium ?

D.N.H. CAIRNS — The cable outage time after two faults is determined by the minimum circulation time of the helium. Using spare refrigerator capacity to allow increased pumping losses this time is at present estimated to be about 10 hours.

N. KURTI (U.K.) — The frequent mention of duplication, built-in redundancy etc. makes me wonder whether we are applying the same reliability criteria as for conventional transmission lines. Or could it be that since cryocables will probably be used only for transmission in the multi-MW range specifications have been made more stringent ?

P.H. ASHMOLE (U.K.) — Our design philosophy for conventional transmis­sion is to provide redundancy by duplication of circuitry. The degree of duplication provided is related to the reliability of the individual transmission line and the impor­tance of the line in the network as a whole.

Thus if a particular superconducting cable has a lower reliability than a comparable conventional cable, allowance would be made for this by increasing the degree of duplication and there will be an economic balance between the degree of reliability it is economic to build into a particular cable and the degree of circuit duplication. A further consideration is the cable load factor which will decrease with increasing circuit duplication. Thus if there is a 100% duplication of circuit capacity the maximum load factor is unlikely to exceed 50%. On the British Grid system transmission load factors are considerably less than this and operating losses are relatively unimportant in any economic appraisals.

162

ETUDE G£N£RALE DES PROBLEMES LI£S A LA REFRIGERATION D'UNE LIAISON ELECTRIQUE

SUPRACONDUCTRICE A COURANT CONTINU

A. M. SCHWAB (Mile), G. DELILE et Y. JEGOU Direction des Etudes et Recherches, Electricite de France, Clamart {France)

A general study of problems associated with the cooling of a direct current supercond­ucting cable

SUMMARY: The study of general problems is cooling a direct current superconducting cable via the circulation of liquid helium and nitrogen. This cryo-cable is equipped with refrigerating systems regularly spaced out along its entire length; with regard to the cooling of this cable\ one has first to consider two geometric parameters associated with the power carried by the cable and its service voltage and, secondly, its length.

The characteristics of the fluids entering and leaving the cooling systems are a function of these parameters, as well as the number of cooling stations {or cable sections) from which one can calculate the total costs for minimum cooling needs.

The authors have established a certain number of simple and general laws which may be directly used to determine the general characteristics of direct current superconducting cables of any voltage capacity and lengths.

1 — OBJET DE L'ETUDE

Une cryoliaison a courant continu par cables supraconducteurs peut etre caracte-risee du point de vue de la refrigeration par deux parametres geometriques lies a la puissance electrique transitee par la liaison et a sa tension de service; ces parametres sont d'une part, le diametre d0 du canal d'helium (fig. 1) dont les parois constituent le conducteur et qui est done une fonction de l'intensite du courant et d'autre part, le diametre interieur dx du canal annulaire d'azote qui est separe du tube conducteur par une epaisseur d'isolation electrique fonction de la tension de courant.

L'etude generate de la refrigeration d'une liaison supraconductrice peut done etre fractionnee en deux sous-etudes : l'etude de la refrigeration par l'helium liquide liee au parametre d0 et l'etude de la refrigeration par azote liquide liee au parametre dx.

Le cout de la refrigeration d'un cable contribuant pour une part tres importante a son cout total il est primordial de chercher a minimiser ce cout. On determine ici en fonction des parametres geometriques du cryocable et pour chaque etage de refrigeration les caracteristiques des fluides a l'entree et a la sortie des refrigerateurs ainsi que l'espacement de ces refrigerateurs qui conduisent a un cout total de refri­geration minimal.

2 — SYSTEME DE REFRIGERATION

La cryoliaison etudiee de longueur L est formee de deux cables unipolaires de structure coaxiale (fig. 1) assurant Taller et le retour du courant continu. Les stations de refrigeration sont reparties regulierement tout au long de cette cryoliaison (fig. 2) et sont couplees a deux sections adjacentes du cable divisant la cryoliaison en m troncons de longueur /. Les fluides de refrigeration circulent en sens contraires dans deux troncons paralleles ainsi que dans deux sections adjacentes. On a suppose que les troncons des deux extremites de la liaison avaient une longueur 1/2; ainsi tous les refrigerateurs ont le meme flux de chaleur a evacuer et l'etude d'une cryoliaison de longueur L peut se ramener a l'etude d'un troncon de deux cables de longueur /.

163

On n'a pas tenu compte dans le bilan thermique des pertes dues aux amenees de courant.

1 _ Helium liquide

2 _. Couche supraconductrice sur tube d aluminium

3 _ Vide pour isolations thermique et electrique

4 _ Azote liquide 5_ Superisolation 6_ Gaine de protection

Fig. 1 — Structure du cable supraconducteur.

R : Stations de refrigeration

L : Longueur totale de la cryoliaison

t : Longueur d'un troncon de cryoliaison entre deux stations de refrigeration

Fig. 2 — Schema des circuits de refroidissement de la cryoliaison supraconductrice.

A la sortie des refrigerateurs (ou a l'entree d'un troncon de cable) la temperature de 1'helium est prise egale a 4,2°K; a l'entree des refrigerateurs (ou a la sortie d'un troncon de cable) elle est prise egale a 5°K. C'est pour ne pas reduire sensiblement les performances du supraconducteur que Ton a adopte cette faible valeur de l'echauf-fement de rhelium. La pression de l'helium est choisie de facon a le maintenir en phase liquide sur toute la longueur d'un troncon. Ces choix fixent une valeur minimale pour la pression a l'entree de la station de refrigeration egale a 2.105 P. Pour 1'azote, l'ecart de temperature pouvant etre admis est beaucoup plus important; neanmoins comme on desire egalement rester en phase liquide tout au long de l'ecoulement sans etre oblige d'utiliser des pressions de fonctionnement trop elevees, ceci pour des raisons d'ordre technologique, on s'est limite a une elevation de temperature de 77 a 90°K. Dans ces conditions la pression a l'entree des refrigerateurs d'azote devra etre superieure ou egale a 4.105 P.

164

3 — CARACTERISTIQUES DE L'ISOLATION THERMIQUE

Le canal de circulation d'helium liquide (4,2 °K) et la partie supraconductrice du cable sont places dans une enceinte thermique tubulaire sous vide (fig. 1) dont la surface exterieure est maintenue a la temperature de l'azote liquide (77 °K). Les pertes thermiques parietales s{ qui procedent d'un phenomene combine de radiation et de conduction au travers du vide et des cales supportant le conducteur ont ete prises egales a 0,1 W/m2 a 4°K.

L'enveloppe cryogenique exterieure est constitute par une superisolation sous vide de 2 cm d'epaisseur qui isole I'ecran thermique refroidi par circulation d'azote liquide du milieu a la temperature ambiante. II parait raisonnable d'adopter a ce niveau de temperature (77 °K) des pertes thermiques parietales s2 de 2 W/m2 pour tenir compte a la fois des performances des superisolants et des pertes par conduction au travers des differents supports.

4 — MISE EN EQUATIONS

Les equations qui vont etre donnees sont valables pour l'helium liquide ou l'azote liquide.

4.1 —EQUATIONS RELATIVES A UN TRONCON DE CABLE DE LONGUEUR / ENTRE DEUX STATIONS DE REFRIGERATION

— Un fluide cryogenique place dans une enceinte ayant deux parois a des tempera­tures differentes, recoit une quantite de chaleur q en provenance de sa paroi en contact avec le milieu a la temperature la plus elevee et transmet une quantite de chaleur q' au travers de sa paroi en contact avec le milieu le plus froid. Le bilan thermique pour le fluide cryogenique sera :

1i = 4-<l ' Dans le cas du cryocable on a pour l'helium q' = 0 et pour l'azote q' < q. On a done

ql = q = sndl (1)

ou s est le flux de chaleur par unite de surface, d le diametre exterieur du canal et / sa longueur.

— Le fluide de refrigeration en circulation forcee subit une perte de charge et un echauffement qui proviennent des forces de frottement existant entre les filets de fluide voisins animes de vitesse differente.

— La perte de charge peut etre calculee en premiere approximation par la formule generate :

A/7 = p2~Pi = i p M V 2 — \|/ (2) dH

ou /?! et p2 sont les pressions du fluide a l'entree et a la sortie d'un troncon de cable, pM la valeur moyenne de la masse specifique du fluide sur un troncon [2] et [3], V sa vitesse, dH le diametre hydraulique du canal, et i// le coefficient de perte de charge de Darcy. Dans le cas d'un fluide cryogenique incompressible en ecoulement turbulent dans une conduite tubulaire, le coefficient \j/ peut etre exprime par la formule empirique :

i|/ = 4[0,0014 + 0,125 (Re ) " 0 ' 3 2 ] (3)

165

ou Re est le nombre de Reynolds :

Re = ^ ^ (4) V

v etant la viscosite dynamique du fluide.

La chaleur degagee au sein du fluide est

4 2 = A p - V - S (5)

ou S est la section du canal.

— Le debit massique du fluide de refrigeration dans un troncon est done :

D M = p - V - S = i i ± ^ (6) AH

ou AH est la variation d'enthalpie du fluide entre l'entree et la sortie d'un troncon de cable [2] et [3].

— La pompe de mise en circulation doit fournir au fluide une puissance q2 = A/?. V. S; elle consomme une puissance electrique :

ou r\p est le rendement de la pompe Les pertes, soient [(g2lr\P) ~(li\ s o n t dissipees au sein du fluide.

— Les pertes totales a basse temperature qui doivent etre evacuees sont done pour un troncon de cable de :

P, = ?l+flf2 + - - « 2 = < h + - (8)

— La puissance consomme pour la refrigeration d'un troncon de cable est :

T1,P, = ! ! . / « , + - ) (9)

rjr etant le facteur de refrigeration.

4.2 — EQUATIONS RELATIVES A UN TRONCON DE CRYOLIAISON DE LONGUEUR /

Une section de cryoliaison de longueur / comprise entre deux stations de refri­geration (fig. 2) comporte deux troncons de cable de longueur / et un refrigerateur de puissance 2 Pr (a basse temperature).

— La puissance electrique consommee a 300 °K par un refrigerateur est :

2 P r i l r = 2 1 1 r f 9 l + ^ ) (10)

166

— La puissance electrique consommee par les pompes de mise en circulation du fluide pour les 2 troncons de cable est :

2 P , = 2 ^ ( " >

— Le volume de fluide cryogenique necessaire au remplissage des deux troncons de cable est :

0 = 2 Si (12)

— Le cout des pertes electriques dans le refrigerateur et la pompe est :

Ctjai^+^ + 2^] (13)

Cj etant le cout actualise du watt de pertes electriques. — Le cout d'un refrigerateur peut etre exprime en fonction de sa puissance 2 Pr

par la loi de variation suivante [1] :

C ^ Y _ <J«Il±S« f (M) \ A 0 / L *o J

ou C0 est le cout d'un refrigerateur de reference de puissance PQ et n un indice inferieur a 1 qui depend du fluide cryogenique (les deux puissances, 2 Pr et P 0 , doivent etre prises a la meme temperature).

— Le cout du fluide cryogenique est :

C 2 - 2 S - Z (15) C 2 etant le cout de l'unite de volume du fluide cryogenique.

— Le cout total par unite de longueur de cryoliaison de la refrigeration par circulation d'helium ou d'azote liquide est donne par:

5—RESULTATS DES CALCULS ET CONCLUSIONS

On a determine a l'aide des equations precedentes, d'une part pour la refrigeration par Thelium liquide en fonction du diametre d0 du cylindre conducteur et d'autre part pour la refrigeration par I'azote liquide en fonction du diametre dt du canal annulaire, les valeurs de la distance entre stations de refrigeration et des pressions d'entree et de sortie du fluide cryogenique qui conduisent a un cout total de refri­geration minimal. II faut noter que les couts et les facteurs de refrigeration sont connus avec une assez grande incertitude. Avec les valeurs generalement admises par differents auteurs on est conduit aux resultats suivants :

5.1 — REFRIGERATION PAR HELIUM LIQUIDE

Les courbes representees sur les figures 3 a 5 montrent qu'a 1'optimum economique: — Les pressions de service pY et p2 sont faibles et pratiquement constantes •

p2 optimale ~ 2.105 P. 2,3.105 P ^ px optimale < 2,5.105 P. La perte de charge Ap = px — p2 reste faible et varie peu.

167

4 Distance optimale entre r6frigerateur d'heiium (km)

FIGURE 3 , Variation de la distance entre refrigerateur d'helium,correspon­dent au cout de refrigeration minimal en fonction du djam&tre du canal d'heiium.

4 6 Diametre du canal d'heiium (cm)

^ Pression optimale d'entr6e (P )

2,5.10 FIGURE 4 - Variation de la pression d'entree de fheiium, correspondent au cout de refrigeration minimal, en fonction du diametre du canal d'heiium.

4 6 Diametre du canal d'heMium (cm)

FIGURE 5 _ Variation du cout minimal de la refrigeration par helium,par unite de longueur de cryoliaison,en fonction du diametre du canal d'heiium.

4 6 Diametre du canal d'heiium (cm)

168

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— Le quotient ljd0 est approximativement constant et egal a 4,6.105. — Les pertes de chaleur par frottement visqueux restent faibles devant les pertes

parietales. — Le cout total par unite de longueur de cryoliaison de la refrigeration par helium

liquide peut se mettre sous la forme :

Cr ( H e ) # 1,2.103 d0 + 2,5.102 C 2 2 + 3 , 1 4 . 1 0 4 ^

ou Cr(He) est exprime en F/m et d0 en m.

5.2 — REFRIGERATION PAR AZOTE LIQUIDE

Les courbes tracees sur les figures 6 a 8 donnent en fonction du diametre dx du canal, son diametre hydraulique, le pas des refrigerateurs, la perte de charge sur une section de cryocable et le cout specifique minimal de refrigeration correspondant. On remarque qu'a l'optimum economique :

— Les pressions de service pY et p2 sont assez elevees

p 2 optimale ~ 4 .105P 12.10 5 P < px optimale ^ 8 .10 5 P ;

— Le quotient //2.a, du pas des refrigerateurs sur le diametre hydraulique du canal, est sensiblement constant et egal a 1,15.106;

— Les pertes de chaleur par frottement visqueux restent faibles devant les pertes parietales;

— Le cout du remplissage de la cryoliaison par l'azote liquide est negligeable devant les autres couts.

REFERENCES

[1] K.J.R. WILKINSON, Proc. IEE, 113, 9 (1966), pp. 1509-1521 et Proc. IEE, 114, 12, (1967), pp. 1892-1898.

[2] D.B. MANN, NBS Technical Note, 154 (Jan. 1962). [3] T. R. STROBRIDGE, NBS Technical Note, 129 (Jan. 1962).

DISCUSSION

D. A. SWIFT (U.K.) — Since coaxial conductive arrangements are probably cheaper than two separate cores, have the authors analysed optimum coolant condi­tions for this type of arrangement?

Ml l e A.M. SCHWAB — Nous n'avons pas etudie les conditions optimales de refrigeration pour une disposition coaxiale des conducteurs dans le cable. Cepen-dant, les resultats obtenus etant tres generaux, ils pourraient facilement etre appliques a ce type de structure.

170

COOLING OF SUPERCONDUCTING CABLES

A. SELLMAIER Linde A G, Hollriegelskreuth (West Germany)

Refroidissement des cables supraconducteurs

RESUME : Le cout des cables supraconducteurs depend du cout du materiel frigorifique et du cout du cable.

Theoriquement on peut obtenir de tres faibles pertes desolation thermique des cables, mais, dans ce cas, le cout des cables sera relativement eleve. Vun des principaux problemes est de parvenir a une solution optimale de la conception des cables, compte tenu des groupes frigori-fiques necessaires.

En vue de cette optimisation, on donne quelques points de vue sur les possibilites de refroidisse­ment des cables supraconducteurs.

Dans ce rapport, on decrit quelques procedes de refroidissement des cables. On examine les limites thermodynamiques theoriques et le cout d'installation et d'exploitation des groupes frigorifiques realisables dans la pratique.

The development of superconducting cables may be divided in two parts. One part includes all electrotechnical questions and the special problems of the super­conducting material. The other part encloses all questions of the cryogenic envelope and refrigeration. Connections between the two parts are given by the necessary volume and weight of the conductors with their electrical insulation. The heat transfer from the cooling flow to the superconductor and its friction coefficient has an influence on the design of the cryogenic envelope.

The problems are characterised by a great number of thermodynamical and electrical parameters. For further development work it seems necessary to get more knowledge of the influence of these parameters on the cable design and finally on the total economy of superconducting cables. About this matter some contributions with quantitative considerations have been published. Wilkinson [1], Klaudy [2], Swift [3] and Rogers and Edwards [4] have described a.c. cables, Garwin and Matisoo [5] d.c. cables. A special optimization of the flow conditions for a.c. cables has been published by Swift [6]. In these papers questions of optimization especially of a.c. cables have been discussed ant the results give an idea of the possibilities and diffi­culties in the cooling of superconducting cables. The following contribution deals with some special cooling problems.

In the manufacturing of superinsulated pipelines for liquid or gaseous hydrogen or helium, relatively good insulation values have been obtained. These values would be sufficient in superconducting cable application. But the costs of these arrangements are too high. The question is to find a solution with an optimal relation between the costs for manufacturing and installation of the cable on the one side and the costs for the refrigeration equipment on the other side.

Because of this it seems useful to give a survey about the influence of some cryo­genic parameters. We begin with the necessary power consumption and the costs of the refrigeration plants under special conditions.

It seems that in any case the cable has to be thermally insulated with superin-sulation. Furthermore it also seems necessary that the cooling has to be done in more than one stage to keep the costs for manufacturing and installation of the cable in economical limits.

In the following a two stage cooling will be taken into consideration, schematically shown in figure 1. The first stage is operated at about 80 °K, the second stage at about 4.5 °K minimum temperature. The heat fluxes which have to be removed are

171

then Q t and Q2 and related to the unit of the length L of the cable qt = Q t /L and q2 = Q2/L-

Helium flow

Helium return flow

Nitrogen flow

Nitrogen return flow

Fig. 1. — Cryogenic envelope schematically.

Most important for the cable design is the quality of the insulation between the different stages. If we had to calculate only the heat transfer through the superinsulation we could give some values for the optimal insulation. But in industrial manufactured cables the heat transfer will be much more influenced by the supports between the different temperature levels than by superinsulation. Therefore we like to show the influence of the insulation qualities of the cooling stages on the power consumption and on the costs of the refrigerators. For this purpose we define the heat flux ratio v = #2Ah- F ° r the further considerations q2 may also include alternating current losses. To the values shown, the expenses for transportation of the cooling media must be added, i.e. the power consumption and the costs of pumps and of the addi­tional capacities of the refrigerators. But for first information it seems not useful to include data which largely depend on a great number of different parameters. In some cases the influence of this expense is small in relation to the total economy of the cable. These questions will be discussed later on.

Figure 2 shows the power consumption of the refrigeration plants against the heat flux per unit of length at 80 °K, qx with v as parameter. From the equation

Nel/L = Aqi+Bq2 = ^ ( A + vB) (1)

wit A = 8 and B = 400, we obtain the actual specific electrical power for cooling, with the relatively large plants needed. One can see the very important influence of the insulation between the two cooling stages. For comparison also the power con­sumption for a one stage cooling at 4.5 °K, with q2 instead of q{ is shown. A very high quality of insulation is necessary to obtain a comparable power consumption.

To get an idea of the costs for the refrigerators we also use the heat flux ratio v as parameter. The costs depend on the distance between the refrigerators. So we have to

172

1 S * 400

•c * -? 300 M 2 § 200

Si 700

I Cooking at V = 0./

= 0.05

y = 0.01

10 12 qj w/m

Fig. 2

take as second parameter the distance between the refrigerator stations. With the following calculations, we get a survey of the costs for the refrigeration plants under different conditions.

The plant costs for two stage cooling are given by the equation

K = C 1 Qr + C2Q"22;

with Q = q L and q2 = vq± we get

K/L = [C1(«1L)"' + C2(«1vLn/L

(2)

(3)

for the costs of the refrigeration plants per unit of length with the distance L of the plants.

For cooling at 80°K we take Cx = 102; nx = 0.8 and for 4.5°K C 2 = 5.104; n2 = 0.5

With these values we get the diagram figure 3. We see again the very important influence of the insulation of the second stage

and also of the plant distance L. It seems that the insulation at higher temperature is not a serious problem and that greater heat fluxes have not much influence on costs and power. But one has to consider the coolant quantities.

Because it seems not practicable to use the latent heat of the media we always have a single phase flow, which may be subcooled liquid or gas with supercritical pressures.

One can say that maximal heat capacities can be obtained if the cooling flow under supercritical pressure is cooled in the refrigerator to the lowest possible tem­perature and is warmed in the cable to temperatures above the critical temperature. To do so one takes advantage of the very high specific heat of the compressed gases above the critical point. The disadvantage to work near the critical point is that higher cooling temperatures must be allowed. For cooling at the low temperature level this is not as important for d.c. cables as it is for a.c. cables with superconductors which have relatively low transition temperatures. For the first cooling stage higher temperatures of up to about 140°K do not present serious insulation problems. But there is another disadvantage in the relatively high pressure.

For nitrogen the critical pressure is 34.6 atm. The application of pressures up to about 40 atm involves mechanical problems in cable design and new questions of optimization.

173

$600

I "> o500

o $300

200

WO

ysSi q2forT2=4 5°K 91 q1 for Tf = 80°K

L=10 km

05 L=10 km L = 20 km

L=30km 05 L = 20 km

L =40 km

L=30km QJ L=Wkm [5 L =40 km

L=20km

L=30km 01 L=40km

10 12 14 Heat flux q1 w/m

Fig. 3

Because of the great change of the specific heat with changing of pressure and temperature, an exact description of working in the supercritical region is not possible in this short paper. But one can give an approximation of the possible heat capacities for some temperatures and pressures. Table 1 shows some values for the helium and nitrogen cooling streams.

Table 1

Gas

He

N2

T A min °K

4.5 4.5 4.5 3.0

80 80 80 80 80 80 65 65

Tm a x °K

5 6 6.5 5

90 100 110 120 130 140 90

110

Pmin atm

2.6 2.6 2.6 2.6

3.6 8.8

16.5 27.5 36 36

3.6 16.5

Pmax atm

3 5 6 3

10 15 23 34 42 42 10 23

Ah Wh/kg

0.8 6.3 7.2 2.1

5.1 1 10.8 [ 16.9 J 26.0 50.6 59.0 11.6 22.7

Remarks

Tm a x <C Tc = 5.2 K. pc = 2.4 atm

* max "^ *c

T m a x T c = 126.1 °K Pmin Pc = 34.6 atm

Pmin ^ J-l /'saturation

174

One can see that for cooling with maximum temperatures of about 5 °K relatively unfavourable conditions are given. In this case, one should take into consideration the possibility to reduce the minimum cooling temperature to values below 4.2 °K. The obtainable difference of enthalpy for instance with 3 °K is about 260% of the value that will be obtained with minimum temperature 4.5 °K, with the same quantity and pressure drop of the cooling flow. On the other hand the price of a plant for temperatures down to 3°K is higher and requires further expenses to guarantee uninterrupted operation. The helium refrigerator has to work with pressures below atmospheric pressure.

We have the same situation for the first cooling stage. We see in table 1 that for minimum temperature 65 °K, maximum temperature 90 °K, and a pressure drop of 6.4 atm we obtain an enthalpy difference of about 11.6 Wh/kg instead of 5.1 Wh/kg with the same quantity of coolant and pressures. This is a real advantage because there is not much difficulty and price difference if refrigeration at the liquid nitrogen temperature level is done by helium gas cycles or Stirling cycle machines.

With the values of table 1 and equation

M = qL/Ah (4)

we get the necessary quantities of coolant. This and the following equations give us now an approximative information about the size of the channels for the coolant. With

V = M/y ; (5) Ap = v|/yW2L/20d (6)

and d = 4 F / U (7)

we get F3/U = 9.83.10" 10v|/LV/ApyA/!2 [m5] (8)

Where U is the circumference (m), d the equivalent diameter (m), \|/ the friction factor, F the cross section (m2), w the velocity (m/s), g the gravity constant (m/s2)* and q in W/m; Ap in kg/m2; y in kg/m3 and Ah in Wh/kg.

For exact calculation, the change of y with pressure and temperature and the corresponding change of the specific heat has to be taken into consideration.

The application of equation (8) to the values of table 1 and the diagram figure 3 shows that for L = 30 km, circular cross section and \|/ = 0.05, the diameters for nitrogen are approximatively in the range between 40 mm and 250 mm.

For helium the values are between 30 mm and 400 mm. The prohibitively great value of 400 mm corresponds to a heat flux of q2 = 1 W/m and an enthalpy difference of Ah = 0.8 Wh/kg.

The next cryogenic question which has already been mentioned above is the expenditure for circulation of the coolants. For nitrogen this parameter may be neglected for the first stage of development of superconducting cables. For instance, with L = 30 km; A/* = 11.6 Wh/kg, with the mean value y = 800 kg/m3

and the heat flux qx = 10 W/m, we get a coolant flow of about 32 m3/h. With a pressure drop of 6.4 atm the necessary power consumption on the shaft of the pump is in the order of 9 kW. The energy which has to be removed at LN2 temperature level is 300 kW and the additional load of the nitrogen refrigeration only about 3%. The costs of the pump with accessories, together with the additional costs on account of the enlarging of the nitrogen refrigeration will be only about 4% of the costs for the first cooling stage.

Pumping at LHe-temperatures has much more influence on the plant costs and power consumption. Energies in the order of about 0.5 kW to 1.5 kW have to be

175

/—-*■ '—1

H2-*—( IN ILZM

feJ n Fig. 4.

1. Gas compression, purification and storage. 2. First cooling stage 3. N 2 pump 4. Second cooling stage 5. He subcooler 6. He pump

176

absorbed by the second refrigeration stage and the additional expenses in costs and power may be 20% and more. Fortunately in some cases if the needed quantity of coolant is smaller than the helium flow of the refrigerator which has to be expanded in the Joule-Thomson valve, a part of this stream may be subcooled in a LHe-bath and fed into the cable channel. The return flow is expanded into this bath. This is shown in a flowsheet for the plant arrangement figure 4.

The plants contain gas purification, compression and storage elements for the circulating helium and nitrogen.

In the first stage with refrigeration by means of turbines or Stirling machines, the compressed He-gas is precooled and the N 2 gas condensed and subcooled. In the second stage, with the aid of turbine expansion, the helium gas is cooled far below the inversion temperature. One part is expanded directly, the other part is subcooled in a helium evaporator 5. The evaporated return flow is warmed up in the heat exchangers of both stages.

In case of the coolant quantities being similar or greater than the quantities of expanded gas, a helium pump 6 has to be installed. In any case subcooling of the circulating gas before and after the pumps is necessary.

One can see, if the cable has a constant configuration, the plants on the ends would get half capacity. But on the ends also the current leads have to be cooled and the plants must have additional refrigeration capacity for this purpose. If relati­vely small power in the order of some hundreds MW has to be transferred over long distances, one may consider to install on the ends only the refrigerators which are necessary for cooling the current leads.

With one refrigerator operating in both directions plant distances in the order of 40 km and more seems possible if a cooling temperature up to about 6 °K is allowed (d.c. cables). For a.c. cables distances up to 20 km seem to be practicable.

But final considerations are only possible if one has an approximate knowledge of the costs for the industrial fabrication and installation of cables in lengths of 100 km and more. This knowledge can only be obtained by technical development work with the possibility of erecting and testing cables in lengths which are comparable with cable elements for future industrial applications.

REFERENCES

[1] K.R. WILKINSON, Proc. IEE, 113, 9 (1966), p. 1509. [2] P. A. KLAUDY, Adv. Cryog. Engg, 1 (1966), p. 864. [3] D.A. SWIFT, I.I.F.-12e Congr. Int. du Froid, Madrid (1967), I, pp. 173-185. [4] E.C. ROGERS and D.R. EDWARDS, Electr. Rev., 8 (Sept. 1967), p. 348. [5] R.L. GARWIN and MATISOO, Proc. IEEE, 55, No. 4 (1967), p. 338. [6] D.A. SWIFT, Cryogenics, 8 (1968) p. 238.

DISCUSSION

W.T. NORRIS (U.K.) — Your paper emphasized the predominance in refriger­ation costs of the heat inleak to helium temperatures. What would you feel would be the heat inleak expressed per unit area between a LN2 shield and the LHe inner vessel ?

A. SELLMAIER — The heat inleak to helium temperature is given by heat transfer in superinsulation and in the mechanical supports. We feel that the mechanical supports will influence the total heat inleak considerably. The heat transfer through

177

superinsulation per unit area obtained in laboratory measurements is very small, but it is not known what values can be obtained in industrial manufactured cables.

W. T. NORRIS — It is certain that heat inleak along mechanical support will be important, but what is the range of values one might expect ?

A. SELLMAIER — The best values for superinsulation between LN2- and LHe-temperatures obtained in the laboratory are of the order of 0.01 W/m2. A cable would certainly have worse heat leak, mainly because of the supports.

W.T. NORRIS — It is unlikely to be better.

178

SPECIAL CRYOGENIC PROBLEMS ON THE PERFORMANCE OF SUPERCONDUCTING CABLES

U. HILDEBRANDT Linde AG, Hollriegelskreuth {West Germany)

Problemes cryogeniques particuliers poses par le comportement des cables supra-conducteurs.

RESUME : Dans ce rapport on examine quelques problemes cryogeniques essentiels des cables supraconducteurs.

La contraction thermique provoquee par le refroidissement d'un cable et la contrainte des parties elastiques provoquee par la pression du gaz ou du liquide froids entrainent une elevation des pressions du support entre les parties froides et les parties a temperature ambiante. Par la suite, les apports de chaleur augmentent. On montre comment il est possible de reduire les pressions du support.

On decrit Vinfluence des pressions du support sur Visolation thermique. On en tire Vamena-gement optimal des conducteurs et des supports. On etudie les cas dans lesquels une protection thermique est utile. On indique les isolants appropries. On examine les problemes a"evacuation des canaux longs et etroits. On donne des exemples pour la conception de risolation thermique.

On examine quelques schemas de construction des cables supraconducteurs.

In recent years a number of papers on cryogenic cables has been published. The special cryogenic problems however were not mentioned in detail. As efforts in the realization of prototypes of such cables are made at present, it seems necessary to emphasize some of these problems in order that the relevance of cryogenics is adequately taken into account.

The price A of a refrigerator for a cold production at 4.5 °K is roughly proportional to the square root of the refrigeration capacity QK.

A ~ Q* (1)

Therefore it is said, that the distances between the refrigerators along a cable should be made as great as possible. This is not necessarily economic, especially for cables with only one conductor, as will be explained. If an enthalpy difference of the helium flow between two neighbouring refrigerators of the cable of at least 12.5 J/g can be allowed, liquid helium pumps are not needed when a suitable refrigerator circuit is used. At a given diameter of the cold parts of the cable optimum flow conditions for the refrigerant are reached with a circular cross section of the same diameter, the inside mountings having a small cross section compared with the total cold cross section and having a hydraulic circumference in the same order of magnitude as the circumference of the cold cross section. Under these conditions, the following simpli­fied calculation can be made.

If the cable is considered as a pipeline with circular cross section for the transport of helium by turbulent flow at very low temperatures, equation (2) can be derived from the known equation for the pressure drop, from the continuity law and from the heat balance:

L - K . - r f ^ ^ (**)*]* (2)

where L is the length of the line, K a constant, d the inner diameter, Ap the allowed pressure drop, pm the mean density of the helium, % the friction factor, Ah theenthalpy

179

difference of the helium between outlet and inlet and q the heat flux into the pipeline per unit of the inner surface. The refrigeration capacity QK is the product of the heat flux density q and the inner surface of the line between two refrigerators, mounted at a distance L from each other:

QK = n- d-L- q (3)

If the price A for the refrigerators in equation (1) is related to the unit length of the line and equations (3) and (2) are inserted, we get

- = K2 • q* • A/r" (—^—V (4) L \Ap-pm/

K2 is a dimensional constant. According to equation (4) the specific capital costs for the refrigerators do not

depend on the diameter d. As the inner surface of the pipeline is proportional to d, the required refrigerator capacity and the power consumption per unit length of the line rise at a constant q proportional to d. Because of this and because of the lower helium hold-up, the lower weight, the lower costs of the cable and other advantages the diameter d should be made as small as possible, although the distance between the refrigerators could be increased using a higher diameter d (see equation 2).

Equation (4) shows clearly, what has to be done in technology under the conditions given above. The minimization of the heat input represents the main problem. It is mentioned in another paragraph. Firstly the thermodynamic values are considered.

By suitable assumptions of the gas states at the inlet and the outlet of the cable the product

can be minimized. Supercritical states must also be taken into account. The enthalpy difference depends on the state at the inlet of the cable and on the allowed maximum temperature in the cable. Because the change of state of the helium is taking place near the critical state, the maximum temperature in the cable can arise at a point different from the helium outlet. For superconductivity the maximum temperature should be low, for refrigeration it should be high. Therefore, the maximum temperature has to be optimized carefully. It seems, that the favourable maximum temperature at least for d.c. cables is higher than the critical temperature of helium, i.e. 5.2°K.

However the most important parameter in equation (4) is the heat flux density q, as it has the highest exponent. Moreover, the power consumption of the refrigerators is proportional to q. Therefore the minimization of the heat input is, as known, the most important cryogenic problem on the performance of superconducting cables.

Heat input occurs by insulation losses, i.e. by radiation and heat conduction. Other heat sources are the electrical losses and the heat of friction of moving cold parts.

The first prerequisite for good thermal insulation is an evacuated space between the cold and warm parts. In this way heat transfer by convection is prevented and at sufficiently low pressures heat conduction of the residual gas is reduced significantly. As is known a high quality insulation requires a pressure < 10" 4 Torr. The maximum distance between the vacuum pumps along the cable depends on the flow resistance per unit length of the vacuum channel, on the leakage per unit length of the vacuum channel, on the absolute temperature, on the molecular weight of the flowing gas and on the allowed pressure drop/?m a x-/?m i n , where pmaK is the allowed maximum pressure

180

and pmin the suction pressure of the pumps. At a pressure of 10 4 Torr molecular flow exists. The pressure drop is calculated analogously to Ohm's law (see fig. 1).

dp = — m R dx

where m is the gas flow, R the flow resistance per unit length.

(5)

vacuum pump Q-

mcaJ>le Pmax

2

m Pmin — ? vacuum

\pump

2

Fig. 1

At a constant leakage s per unit length we get

m = s • x

Equation (6) is inserted in equation (5):

I Pmin f i dp = — s - R x dx

Pmax J 0

(6)

(7)

The solution is

/ = " Pmin

s -R (8)

/ is obtained in cm when Ap is taken in Torr, s in Torr litre/sec.cm at 293 °K and R in sec/litre, cm. As is known R is given by the expression

R 6 1 . 8 - p - F '

M_ 293 r sec 1 28.7 T L litre.cm J

The circumference u of the cross section F of the channel has to be taken in cm, Fin cm2 and T in °K. M is the molecular weight, {3 a factor which equals unity for circular cross sections. For rectangular and for annular cross sections P is given by figure 2.

Using equation (8) the maximum distances between the vacuum pumps at steady state conditions can be calculated. The actual minimum leakage per unit area attain­able with modern methods of manufacturing is roughly 1 0 " 1 0 Torr litre/seam2. For the calculation of the distances between the pumps starting conditions must also be considered. The outgassing of metals and insulation materials is initially some orders of magnitude higher than at steady state conditions, for stainless steel for instance 10"3-10"4Torr litre/sec. m2. It is difficult, to reduce the time of outgassing by heating, as the heating temperature is limited by the heat-sensitive insulation materials and by the thermal expansion of the heated materials. It is favourable, to calculate the

181

distances according to the outgassing and to stop the unnecessary vacuum pumps when outgassing has ceased. Using equation (8) in this case, a maximum pressure of 10" 2 Torr can be allowed. At this pressure refrigeration can begin and outgassing is reduced with falling temperature. When working temperature is approached, the cold

0 0.2 0.4 0.6 0.8 1.0

Factor pr for rectangular and pa for annular cross sections

b

a ' - 6 Z =

d_2

Fig. 2

parts of the cable act as cryopumps and the vacuum pumps have to remove only the gas flow coming from the helium leaks. Other leakages would affect the thermal insulation as the gas flow freezes out on the cold parts of the insulation, thus worsening the emissivities of the radiation shields. When a great air-leakage occurs during operation and air freezes out in presence of combustible materials danger of explosion exists because of the high density of oxygen. For these reasons an excellent tightness of the vacuum channel is required. With a modern leak detection equipment single leaks of 10~1 0 Torr litre/sec can be localized.

The problem of the thermal insulation must be seen in connection with the support forces which have to be transmitted between the cold and the warm parts of a cable. The forces are caused by the weight of the cold parts, by the pressure of helium, by thermal contractions and by the magnetic field of the conductors if the electrical insulation has ambient temperature.

To avoid the transmission of the magnetic forces through the thermal insulation the conductors can be arranged concentrically, or parallel conductors can be connected by a bandage, the electrical insulation being on low temperature. This is advantageous if the forces caused by the weight of the electrical insulation and the bandage are low compared with the magnetic forces and if the additional heat input caused by the high circumference of the cold electrical insulation is lower than the saving of heat input caused by the compact arrangement of all conductors inside a single thermal insulation. The cryogenic decision, whether a parallel or a concentric arrangement of the con­ductors with cold electrical insulations respectively should be chosen, only depends on the minimum weight and the minimum circumference of the cold parts.

The changes in the length of the cold parts of a cable represent a difficult problem. The linear contraction of most metals and alloys during cool down to very low temperatures is about 0.2-0.5%, the contraction of plastics however is about ten times this value. As an axial change of length of the cold parts must be prevented, possi­bilities for elastic deformations must be provided. That means, that the pressure of helium counteracts the thermal contraction. As this pressure is not constant along the cable, a fairly good compensation of these forces can be achieved by the close arrangement of conductors with cold electrical insulations and cold mechanical couplings between those parts in which the helium is flowing in opposite direction.

182

The cold parts should be designed in such a manner, that a low tensile force exists during operation to prevent buckling and that they have a constant elasticity over the length of the cable, rather than an elasticity at single points. Thus axial movements during cool down are prevented. Furthermore a continuous manufacturing and even­tually reeling are made possible.

For the transfer of forces between the cold and warm parts special supports or the thermal insulation itself can be used. Without pressure load currently only multi­layer insulations have sufficient insulation properties. The increase in their heat flux is roughly proportional to the two-thirds power of the compressive load as was measured by Glaser et al. [1]. That means, that the pressureloaded multilayer insulation is most effective, when it is not loaded homogeneously with low pressure but when it is loaded at few points with high pressure.

Glaser has described the properties of different multilayer insulations in detail. In principle they consist of many alternating layers of metallic radiation shields and of spacers with low thermal conductivity, e.g. of aluminium foils with glass paper spacers or of polyester films coated on one side with vacuum-deposited aluminium and crumpled for reducing the heat conducting contact areas. The radiation heat transfer is inversely proportional to 1 + n, if all layers have the same emissivity, n being the number of layers. Solid heat conduction is very low because of the loose packing of layers, giving low pressure loads.

Without pressure load, a multilayer insulation consisting of 0.05 mm aluminium foil and 0.5 mm fiberglass-netting with 30 layers at a total thickness of 25 mm has a heat flow of about 0.24 W/m2 between 293 °K and 4°K. This is one of the best values obtained at present.

warm support rings 8 space for

evacuation

Fig. 3

When the total thickness of an insulation is limited, the use of cooled interfaces is advantageous. In this case, the insulation can be designed in such a way, that at high temperatures low insulation thickness and at low temperature high thickness is used. This is especially important when forces have to be transmitted through the insulation. Superconducting cables should therefore be equipped with a cooled inter­face.

Of course also the heat producing electrical losses and the heat production by friction of moving cold parts must be prevented. Especially the oscillations of cold parts of three phase current cables with mains frequency can lead to serious losses. With d.c. cables such losses can occur by an insufficient noise elimination or by

183

frequent changes of current loading. Therefore radial movements must be made impossible by stable joints, on the other hand axial elastic deformation must be possible.

Figure 3 shows the first example of a thermal insulation. All cold parts, the electrical cold insulation included, are arranged in a flexible tube 2 as proposed by Klaudy [2] The cold support rings 4 are fastened to the flexible tube, the warm support rings 3 can slide in the outer tube, their distances remaining constant because of the spacer 7. The

\ \ N \ N \ \ i k v u v v i vV r\ \ v v v v v v v v v i v i v i

F i g . 4

/ warm surface 2 multilayer or plastic

support spiral 3 space for

evacuation 4 multilayer

insulation 5 cold surface

1 warm surface 2 vacuum channel 3 single spacers 4 perforated tube 5 multilayer

insulation 6 multilayer supports 7 cold surface

Fig. 5

1 2 3 4 5

1 warm surface 2 space for

insulation and evacuation

3 spiral tubes 4 space for

insulation 5 cold surface

Fig. 6

184

wires 5 transmit the axial and the radial forces as well. An axial vacuum gas flow exists. The same principle of insulation can be used with a cooled interface. In this case the warm surface 1 would be the interface.

In figure 4 another possibility of transmission of radial forces is shown. Axial forces can only be transmitted by special arrangements provided in certain distances along the cable. Manufacturing of the insulation is relatively simple as all layers can be wound continuously. If a plastic spiral is used it is designed in such a manner that it has only few contacts per unit length with the cold and the warm parts. For evacu­ation, this construction is unfavourable, as the vacuum channel has a spiral shape with a great length. If 1 is the cooled interface and the vacuum channel lies outside the interface the construction becomes more favourable.

The multilayer supports in figure 5 are advantageous as they can be spaced arbi­trarily. The total cross section of all supports can be minimized in this way. The disadvantage is, that a separate vacuum channel is used and that axial forces cannot be transmitted.

A cooled interface could be realized by spiral tubes with rectangular cross section as to be seen in figure 6. Also fin tubes are suited. Continuous manufacturing would be possible. Axial contractions can be compensated by elastic deformations. In the spaces 2 and 4 insulations according to figures 3-6 can be arranged.

Attention is called to the fact, that the examples shown here do not represent solutions of general validity as development on this subject has just begun in recent years. For the same reason a series df the other problems mentioned can only be solved in future.

REFERENCES

[1] P.E. GLASER, J.A. BLACK, R.S. LINDSTROM, F.E. RUCCIA, A.E. WEXLER, Thermal Insulation Systems, a Survey, NASA SP-5027 Technology Utilization Division, Washington, D.C. (1967).

[2] P. KLAUDY, Elektrotechnik und Maschinenbau, 82 (1965) 6, pp. 275-281.

185

PROTECTION DES DISPOSITIFS SUPRACONDUCTEURS CONTRE LES SURINTENSITES

CONSEQUENCES DIMENSIONNELLES

Ph. BARRET Direction des Etudes et Recherches, Electricite de France, Clamart {France)

Protection of superconducting devices against overload. Dimensional characteristics

SUMMARY: In order to use superconductors in A. C. cables or transformers very exacting conditions must be satisfied because of the inevitable overloads particularly in case of a short-circuit. The studies published hitherto accepted either a heavy overdimensioning of the apparatus or a non-availability of long duration in case of a short-circuit.

A protective device is proposed here which allows to avoid almost entirely both these draw­backs. The paper examines the consequences of the use of this device from the point of view of both dimensioning and the exploitation of the grid.

1 — INTRODUCTION ET HYPOTHESES

Le probleme des surintensites a une importance tres grande pour le dimension-nement des dispositifs utilisant les supraconducteurs pour l'Electrotechnique de puissance, a cause de la dissipation calorifique due au passage d'un courant important dans un supraconducteur redevenu normal. Si nous considerons les etudes publiees sur les cables supraconducteurs triphases, nous voyons que ce facteur necessite soit un surdimensionnement tres onereux du cable afin qu'il reste supraconducteur meme en cas de court-circuit [1], soit une indisponibilite de longue duree apres un court-circuit si le cable est dimensionne pour le courant nominal [2]. II est possible d'eviter simultanement ces deux inconvenients, en cherchant a realiser une detection et une elimination ultra-rapide du court-circuit. Nous nous proposons de montrer les possibilites qui existent a Theure actuelleen ce domaine, en etudiant tout d'abord un cas particulier puis en cherchant a le generaliser.

Nous considerons done un cable triphase du type de celui etudie par Swift [2], de longueur moyenne (environ 20 km) et destine a alimenter une clientele urbaine; e'est en effet ce type d'utilisation qui a le plus de chances de realisation prochaine, a cause de la faible longueur et du besoin des grandes villes en canalisations electriques de grande puissance. Dans ce cas on n'a pas a considerer les surintensites dues aux oscillations pendulaires entre generateurs et on examinera seulement le cas du court-circuit, suppose triphase, exterieur au cable (un court-circuit interieur au cable necessitant de toutes f aeons une longue duree de mise hors service).

2 — PRESENTATION D U PROCESSUS ETUDIE

Le schema est represents figure 1. Le cable etant parcouru par son courant nominal I„, un court-circuit triphase apparait a Tinstant / = 0 a l'exterieur du cable de sorte que la valeur efficace devient ki I„; kx caracterise la puissance de court-circuit du reseau.

Le systeme de detection se declenche lorsque le courant instantane dans une phase quelconque atteint la valeur k2 I„ y/2; k2 est un nombre superieur a 1 que Ton peut choisir arbitrairement. Cette valeur est atteinte au temps t0 ; nous appelons tx la duree de la detection, telle qu'au temps t0 + tx la puissance necessaire pour actionner le disjoncteur est disponible.

189

L'elimination du defaut se fait en enclenchant les deux disjoncteurs triphases Dj (fig. 1) qui court-circuitent le cable a ses deux extremites. Nous appelons t2 le temps d'action des disjoncteurs, de sorte qu'au temps t0 + tt +1 2 les deux disjoncteurs sont fermes. Nous appelons T = tx +t2 le temps d'elimination.

Le court-circuit est ensuite definitivement elimine par l'ouverture des deux disjoncteurs classiques D 2 (fig. 1).

Reseau generateur

Cable supraconducteur

Reseau D2 recepteur

Defaut

Fig. 1

Transmission lumineuse

Transformateur d' intensite

= *■

Fig. 2

Les phenomenes dissipatifs dans le cable sont alors les suivants : pertes par effet Joule dans le support du supraconducteur pendant la duree du court-circuit; dissi­pation de I'energie magnetique emmagasinee dans le cable; dissipation de I'energie electrostatique.

Nous allons examiner tous les points qui viennent d'etre evoques afin de montrer comment on peut obtenir une disponibilite immediate du cable apres le court-circuit, meme si le cable n'est pas surdimensionne.

DETECTION RAPIDE

Le schema de principe de la detection est represents figure 2. Le transformateur d'intensite est connecte au detecteur A. Une diode a effet tunnel fixe le seuil de detection et, a partir de ce seuil, actionne une diode electro-luminescente par l'inter-mediaire d'un amplificateur. La transmission lumineuse L assure l'isolation electrique entre A et B. Elle permet de remplacer le transformateur d'intensite par un shunt dans le but d'accroitre la rapidite et la surete d'intervention. Le recepteur B comporte une photodiode qui actionne un thyristor de puissance par Fintermediaire d'un amplificateur. A son tour, le thyristor peut actionner la commande du disjoncteur. Le dispositif a ete essaye a la Station d'Essais de Disjoncteurs de l 'E.D.F. Le temps d'action de l'ensemble A + B, c'est-a-dire le temps qui separe le signal de la diode tunnel de la commande du thyristor de puissance, a ete trouve egal a 20 (is. Nous pouvons done considerer que le temps de detection tx du paragraphe 2 sera de l'ordre de 20 us.

190

4 — DlSJONCTEUR RAPIDE

On peut envisager d'utiliser pour les disjoncteurs Dx de la figure 1 des enclencheurs rapides dont un prototype est en cours d'etude a la Direction des Etudes et Recherches de l'Electricite de France. Cet enclencheur est constitue par un piston mis en mouve-ment par explosif. La longueur totale de la course du piston est de 45 mm, et la duree de pafcours, compte tenu du temps de mise en vitesse, est de 800 [is. Le temps qui separe la commande du thyristor du paragraphe 3 et l'explosion est negligeable. On peut done considerer que le temps d'action du disjoncteur t2 (paragraphe 2) sera de l'ordre de 800 us. Une duree d'enclenchement encore plus courte (de l'ordre de 50 jis) pourrait etre obtenue par amorcage d'un arc sous vide.

En consequence, le temps d'elimination T = tl + t2 peut etre pris egal a 10" 3 s dans le cas de l'enclencheur mecanique a explosif, et on peut esperer atteindre 10 ~4 s en utilisant un enclencheur sous vide.

Nous examinerons en premier lieu les consequences d'une elimination en 10" 3 s.

5 — DISSIPATION CALORIFIQUE PENDANT LE TEMPS D'ELIMINATION T

La loi de variation du courant i en fonction du temps est representee figure 3: court-circuit au temps 0, detection au temps t09 disjoncteur enclenche au temps t0 + t1 + t2 = t0 + T. Nous supposons que le cable supraconducteur est dimensionne

Fig. 3

pour le courant nominal; il en resulte que le courant critique est superieur au courant nominal ([2] et [3]), et nous pouvons admettre que I'etat supraconducteur subsiste jusqu'a l'instant t0. Entre l'instant t0 et l'instant t0 + T, le courant passe dans le support du supraconducteur constitue par un metal normal de tres faible resistivite. La dissipation calorifique correspondante est obtenue en tenant compte de l'effet de peau en regime transitoire [4]. Soit Wx la quantite de chaleur ainsi dissipee par metre de longueur du cable. En prenant comme exemple le cable donne en reference [2], avec kY = l,k2 = 1,5 et T = 10~3 s, nous trouvons : Wt = 6 J/m.

191

6 — DISSIPATION CALORIFIQUE DUE A LA DISPARITION DE L'ENERGIE MAGNETIQUE

Au bout du temps t0 + T les disjoncteurs D x sont enclenches et les courants dans les 4 conducteurs du cable ont atteint les valeurs + I l 5 — I t , + I 2 , — I 2 ; l'energie magnetique correspondante par metre de longueur W2 = Vi(LilJ + L2I2) doit alors Stre dissipee dans la resistance du circuit. Les courants suivent une loi de decroissance exponentielle repr6sentee figure 3 apres le temps t0 + T. En prenant les m£mes hypotheses qu'au paragraphe 5, on trouve W2 = 137 J/m.

7 — DISSIPATION CALORIFIQUE DUE A L'ENERGIE ELECTROSTATIQUE

Ce terme peut etre neglige, car dans l'exemple choisi il est inferieur a 1 J/m.

8 — CONSEQUENCES SUR L'ECHAUFFEMENT DU CABLE ET SUR LA DUREE DE MISE HORS SERVICE

Le terme preponderant dans l'energie totale W t + W2 a dissiper est done celui qui correspond a l'energie magnetique, soit W 2 . Si cette energie devait etre dissipee dans le cable, l'augmentation de temperature du support serait de 12°K, et il faudrait 20 minutes pour ramener le cable a la temperature de 4°K. II est done necessaire de dissiper l'energie magnetique en dehors du cable, et pour cela on place des resis­tances en serie avec les disjoncteurs D j aux points A et B (fig. 1). Si chacune de ces resistances a pour valeur 5.10"3 ohm, le dixieme de W2, soit 13,7 J/m seulement, sera dissipe dans le cable.lDans ces conditions 1'eleVation de temperature du support est seulement de 5°K. La temperature est d'ailleurs rapidement abaissee par le contact avec l'helium; celui-ci subira une vaporisation partielle correspondant a l'energie W 1 + W 2 / 1 0 = 19,7 J/m.

La temperature necessaire a la supraconduction est done retablie instantanement et les conditions de fonctionnement du cable sont restaurees au bout de 8 minutes du moins sans faire appel a un supplement de puissance frigorifique.

La presence des resistances en serie avec les disjoncteurs entraine l'existence d'une tension residuelle rapidement amortie. L'amplitude crete de cette tension ne depasse pas 1 000 V. De toutes manieres l'enclenchement des disjoncteurs Dx est rapidement suivi par le declenchement des disjoncteurs D 2 (fig. 1) qui isolent le cable du reste du reseau, et assureront sa remise en service sitot disparue la cause du court-circuit.

Du fait de la presence des resistances en A et B (fig. 1), la constante de temps de la loi exponentielle de la figure 3, apres le temps t0 + Tx est d'environ 0,03 s.

[9 — AMELIORATION DU PROCED£ ET GENERALISATION

Nous avons choisi pour le calcul de l'energie a dissiper un temps d'elimination T = 10~3 s mais nous avons vu au paragraphe 4 qu'il est envisageable dans l'avenir d'atteindre T = 10~4 s. II en r6sulte une diminution de W t et W 2 . En particulier, avec nos hypotheses, W t serait alors egal a 0,7 J/m et W2 a 30 J/m. L'utilisation de resistances en serie avec les disjoncteurs permettrait de reduire a 3,7 J/m la dissipation calorifique a evacuer, ce qui correspond a un echauffement de 2°K du support. La supraconduction est done restauree sitdt le court-circuit eiimine et les conditions normales de fonctionnement du cable sont retablies au bout de 30 s, du moins sans faire appel a un supplement de puissance frigorifique.

Le procede etudie peut etre utilise pour la protection de tout autre materiel utilisant des supraconducteurs en courant alternatif, par exemple un transformateur

192

triphase. La possibility de calculer un appareil de ce type sans avoir a tenir compte des surintensites de court-circuit, grace a une protection ultra-rapide, permettrait de reduire condiderablement les contraintes qui pesent sur son dimensionnement et de rendre un peu moins sSveres les performances requises de la part des supra-conducteurs pour que son emploi economique soit envisageable.

10 — CONCLUSION

L'elimination ultra-rapide des courts-circuits permet simultanSment d'eviter le surdimensionnement des appareils de puissance a courant alternatif triphase tels que cables ou transformateurs utilisant des supraconducteurs, et de pouvoir les remettre en service dans le plus bref delai. Des dispositifs permettant d'obtenir ce resultat (detection et disjoncteur ultra-rapides) sont decrits, permettant d'atteindre un temps d'elimination de 10 ~3s et d'envisager 10~4 s a l'avenir. Le probleme le plus delicat reste l'evacuation de l'energie magnetique que Ton peut cependant dissiper en partie a l'exterieur du cable. Moyennant ces dispositifs la remise en service de l'appareil, suppose non surdimensionne en courant, peut etre effectuee sans delai. Le temps necessaire au retour de l'helium a ses conditions thermodynamiques initiales est inferieur a 10 minutes si le temps d'elimination est de 10" 3 s, inferieur a 1 minute si le temps d'elimination est de 10~4 s.

REMERCIEMENTS

L'auteur tient a remercier tout particulierement MM. P. Heroin et J. P. Barret de la Direction des Etudes et Recherches de l'Electricite de France pour leur contribu­tion a la preparation de ce travail dans le domaine des disjoncteurs ultra-rapides et dans celui de la detection utra-rapide.

BIBLIOGRAPHIE

[1] WILKINSON, Prospect of employing conductors at low temperature in power cables and power transformers. Proceedings IEE, 133, 9, (Sept. 1966), p. 1 509.

[2] SWIFT, Prospects for the superconducting A.C. power cable, I.I.F.-126 Congres Inter­national du Froid, Madrid (1967), I, pp. 173-185,

[3] ROGERS et EDWARDS, Design for a 750 MVA superconducting power cable, Electrical Review (8 sept. 1967), p. 348.

[4] Ph. BARRET, Dissipation caloriflque dans un cable supraconducteur triphasS en cas d'eli­mination ultra-rapide d'un court-circuit-Rapport interne Etudes et Recherches de l'£lec-tricit<§ de France, n° HM 041-50 (d6c. 1968).

DISCUSSION

D.A. SWIFT (U.K.) — I should like to make a point concerning how rapidly a shunt switch will have to be operated. With present designs of superconducting cables this time need only be about 100 ms. I don't think this should cause any serious problems. Would the author please like to comment?

Ph. BARRET — The purpose of the communication was precisely to show that there are under development means of clearing very quickly faulty currents, so that no over dimensioning is needed.

193

H.J. SHEPPARD (U.K.) —The device described by the author represents a great advance in the speed of fault clearance by protection which involves the use of a fault detecting relay and the operation of circuit-breakers. The total clearance time of about 1 ms compares with 50 to 100 ms for conventional equipment.

I would ask the author how long the experimental fault detector has been in service and how many times it has operated on fault current, correctly and incorrectly. Can he give similar information for the fast-operating shunt switch ? Is its operating characteristic based on a theoretical study, on design data, or on an oscillogram of an actual operation ?

Ph. BARRET — Figure 3 is the result of calculation. The detector has been tested on full scale, with a current of 15,000 amp. The switch itself still requires some final touches.

C.S. FURTADO (U.K.) — Would it be possible to devise a superconducting "switch", in series with the cable, using for example a lower critical current super­conductor alone or in combination with an applied magnetic field. With such a device the current could either be interrupted, or reduced to a safe value.

Ph. BARRET — The resistance of the superconductor in the cable, returned to normal state after a fault, would not by itself be sufficient to lower significantly the fault current (assuming there is no aluminium coating).

Consequently the superconducting switch in series with the cable would require much more superconducting material than the cable itself, since it must remain superconducting under normal operation of the cable.

194

THE COOL-DOWN OF LONG DUCTS

K. KELLNER, I.P. MORTON S. SALDANHA and R.G. SCURLOCK

Department of Physics, University of Southampton {United Kingdom)

Le refroidissement des longs conduits

RESUME : Le temps de passage des fluides frigorigenes dans un cryocable sera particulierement long pendant les premieres etapes du refroidissement a partir de la temperature ambiante. La resistance au flux du fluide frigorigene sera importante jusqu'd ce que le cryocdble entier soit refroidi a une temperature voisine de celle du regime de fonctionnement. On s'attend par con­sequent a ce que le refroidissement le plus simple et le plus rapide soit obtenu en appliquant la capacite refrigerante totale aux entrees de fluide frigorigene dans le cryocdble.

D'apres V experience sur les lignes de transfer t, on s'at tend a ce que le cryocdble soit alors refroidi par le progres d' un front froid a Vavant duquel la temperature demeurera celle de Vambi­ance et a Varriere duquel elle sera egale a la temperature de fonctionnement. La vitesse de ce front peut etre evalue'e, mais les gradients de temperature a Vinterieur du front ne sontpas connus.

On ddcrit des essais preliminaires sur les profils de temperature pendant le refroidissement a"un long conduit (L/D = 730), thermiquement isole par le vide, par azote liquide introduit par une extremite en utilisant des pressions basses. Les resultats fourniront une base pour la discussion des procedes de refroidissement des cryocdbles.

INTRODUCTION

This paper contains some observations on the problem of cooling-down a cryo-cable to its operating temperature based on previous work by other workers and on some preliminary studies on the cool-down of a vacuum insulated duct with liquid nitrogen at low pressure drops.

In general, there will be two temperatures at which refrigeration will be required in a superconducting cryo-cable; 70-80°K for the radiation shield using nitrogen and 4-7°K for the superconducting elements using helium. At the operating temperatures, two phase flow of the refrigerant will be avoided by the use of sub-cooled liquid (for nitrogen) and supercritical vapour (for helium) at pressures above 1 atmosphere. Calculations show that, with correct design, high mass flows can be conveniently achieved with small pressure drops over lengths of the order of 10 km.

During cool-down of the cryo-cable, however, small mass flows and high pressure drops will arise from the high flow resistance of the refrigerant at ambient temperature, and the cool-down time will tend to be prolonged. It is therefore necessary to examine the cool-down mechanism to see what allowances should be made in the design and operation of the cable and its associated refrigerators to make certain that cool-down is not a major problem.

This paper will be confined to the cool-down of long uniform ducts with liquid nitrogen. The helium cooled core of a cryo-cable will have to be cooled entirely by high pressure helium vapour and radiant heat transfer to the nitrogen cooled radiation shield, and the differences between cool-down and steady operating condi­tions will not be so severe.

Previous workers (see for example Steward et al. (1968) [1]) have examined the cool-down mechanism of ducts with L/D ratios up to 103 using liquid nitrogen (L = duct length, D = hydraulic diameter). Most of this work has been concerned with large pressure drops of the order of 10 atmospheres, with relatively unlimited refrigeration available, and with reducing cool-down times to a minimum. For cryo-cable operation, much smaller pressure drops will have to be tolerated in ducts with L/D ratios in excess of 105.

195

The likely method of refrigerating at a particular temperature level within a cryo-cable is outlined in figure 1. The flow and return lines must be thermally isolated from one another and it is therefore envisaged that they will be situated in separate cables of a pair. The cable refrigerant circuit is a closed one and heat is exchanged with the refrigerator through heat exchangers. Further cooling of the refrigerant may be required after it passes through the circulating pump. The refrigerator and heat exchanger will have to be designed primarily to meet efficient operating require­ments at the working temperature of the cable. The refrigerator will of necessity

Fig. 1

be simple and reliable in design for providing refrigeration efficiently at one tempe erature only. The heat exchanger may require some additional refinements to enable the same refrigerator to be used for cooling-down the cable from ambient temperature. For example, during any cool-down procedure, the return gas will be at ambient temperature for a considerable period of the cool-down time and the heat exchanger must therefore be capable of cooling a small mass flow of return gas over a large temperature interval.

The circulating pump will be designed for the operating condition of high mass flow and low pressure head. However, for simplicity of operation, it will have to be designed to operate during cool-down to provide high pressure head and low mass flow. Under these cool-down conditions, it is envisaged that the most efficient use of the available refrigeration capacity will be achieved by ensuring that the refrigerant which enters the cable is at a temperature close to the operating temperature.

From experience with cryogenic liquid transfer lines, it is expected that under this condition, the cable will cool by the progress of a cold front, ahead of which the temperature will remain at ambient and behind which the temperature will be in the region of the operating temperature. The behaviour of these cold fronts has been studied for conditions of high mass flow and short cool-down times, but they have not been studied under conditions of low mass flow, and there is little information on temperature gradients within the front.

196

An attractive method for refrigerating the radiation shield at 70-80°K is by the use of sub-cooled liquid nitrogen operating at pressures between 5 and 20 atmospheres, It is therefore realistic to consider cool-down with sub-cooled liquid nitrogen and study the temperature profiles of the cold front during liquid flow, two-phase flow and vapour flow within the front. Some preliminary studies have therefore been carried out on the cool-down with liquid nitrogen of a vacuum-insulated copper tube with an L/D ratio of 730, using pressure heads of a few cm of liquid nitrogen.

COOL-DOWN OF A LONG DUCT WITH LIQUID NITROGEN

The experiments were carried out on a 460 cm length of 1/4" copper Yorkshire tubing located in a vacuum-insulated enclosure. The duct was in the form of a U with inlet and outlet points located at the same end of the vacuum enclo­sure. Temperatures at the wall and approximately at the axis of the tube were measured at a number of points along the duct with copper-constantan thermocouples. The outputs from the thermocouples were recorded simultaneously with a multi-channel U.V. recorder operating with chart speed of 1 mm per second. Temperature-time plots for the different thermocouple positions were recorded during cool-down runs with different liquid nitrogen heads. A constant pressure drop across the duct was then maintained approximately by ensuring a constant head of liquid nitrogen at the inlet during each run.

A typical example of the temperature behaviour within the fluid and along the wall of the duct during a cool-down with high mass flow is illustrated in figure 2.

O to 40 6o So loo ao no I bo TlMC IN 5CCONM

Fig. 2

Large temperature oscillations within the duct are evident during the first part of the cool-down when slug-flow occurs. Figure 3 shows the wall temperature behaviour at a lower flow rate. The temperature oscillations observed inside the duct were not so large and it is possible to estimate the heat transfer rates during the whole cool-down. Furthermore, by converting temperature-time plots into temperature-distance plots at different times, a picture of the temperature profiles along the duct can be obtained. The temperature-distance profiles obtained from figure 3 are shown in figure 4. It is clear that only a section of the cold front is being observed at any one time, and that the total linear extent is in the region of 600 cm. While the high temperature edge of the front is diffuse, it should be noted that the cold (77°K) edge is sharply defined and hence has a measurable velocity.

Figure 5 shows the characteristics of one point in the duct during this cool-down experiment. It shows the variation with time of the wall temperature, temperature

197

FLOW RATS = Sg/sec

PRESSURE HEAD- io-Zcm LIQUID NZ

MEW ^ELCOTf OF COLO EDQ£=0. / 7cmhec

TIME IN MINUTES

Fig. 3

26o

198

Fig. 4

difference between wall and refrigerant, and the heat transfer rate. Below 100°K, the heat transfer rate increases sharply with the onset of nucleate boiling heat transfer, giving rise to the sharp definition of the cold edge to the front.

-TIME MHVNUTES

Fig. 5

DISCUSSION

The conclusions from these preliminary studies can best be made by summarising some observations relevant to the cool-down of cryo-cables. 1. At the start of a cool-down, the inlet to the duct should be pre-cooled slowly, preferably with cold gas, and definitely not with sub-cooled liquid, to avoid violent pressure oscillations and flow reversals with consequent damage to the circulating pump and flowmeter. Once the inlet section is cold, the inlet pressure and flow can be slowly increased to the optimum cool-down conditions [1]. 2. The linear extent of the cold front is quite large, in the region of 103 times the hydraulic diameter of the duct, but small in comparison with the length of a cable. Temperature gradients within the front are small at the higher temperatures in the region of vapour heat transfer, become greater during two-phase flow and are greatest during nucleate boiling heat transfer. 3. The velocity of the cold front can be well-defined in terms of the velocity of the rear cold edge of the front. This velocity is related to the liquid mass flow less the mass flow passing through the front. 4. The cool-down mechanism can be illustrated by considering the history of an element of refrigerant. The element passes along the duct from the inlet as sub-cooled liquid with a velocity greater than that of the front. When it reaches the cold edge of the front, the element is heated to its saturation temperature, evaporation com­mences and heat is transferred from the wall to the element as it passes through the front, initially in a two-phase condition and subsequently as a cold gas. Ahead of the

199

front, the element moves along the duct with greatly increased velocity as gas at ambient temperature. 5. Once the complete cold front is established beyond the inlet section, it is expected that the temperature profile of the front will become invariant with time. There will then be three well-defined regions of flow, the sub-cooled liquid stream, the front and the warm gas stream, which can be considered separately. The associated pressure drops across the low velocity liquid stream and cold front will be small compared with the pressure drop across the low density, high velocity, warm gas stream. To a good approximation all the available pressure drops can be assumed to occur across the warm gas stream. 6. Provided sufficient refrigeration is available, the warm gas flow will be limited by the condition that the discharge velocity cannot exceed the velocity of sound. In this type of flow, called "Fanno "-flow [2], the upstream gas velocity (just in advance of the cold front) decreases with increasing L^s/D for constant ratios of inlet to discharge pressures. For example, for L^as/D in the region of 105, the upstream gas velocity is approximately 2% and 3% of the velocity of sound for pressure ratios of 1.5 and 10 respectively. Thus the mass flow of refrigerant through the cold front, and hence the maximum velocity of the front, will be determined by the conditions of Fanno-flow in the warm gas stream. On the other hand, by applying heat balance conditions a second limit to the velocity of the cold front will be determined by the available refrigeration at the inlet. If this second limit is less than that set by the Fanno-flow, then it will be worthwhile providing extra refrigeration from an external reservoir for cool-down; but not, otherwise. 7. The limiting velocity of the cold front set by Fanno-flow of the warm gas can be raised by increasing the discharge pressure, and hence the inlet pressure, keeping the pressure ratio in the region of 1.5. The use of high pressures will be limited by the mechanical design of the cryo-cable, but would be advantageous in reducing the problem of external storage of refrigerant gas. 8. It is almost certain that the pressure and flow conditions for optimum cool-down procedure will not be the same as those for operating at the working temperature. The circulating pump and refrigerator heat exchangers should therefore be designed to cope with both sets of conditions. 9. Flow and return ducts cannot be tolerated within the same cable envelope-Thus cables will have to be operated in pairs. 10. Most of these observations are relevant to the cool-down of the superconducting core of a cryo-cable with helium gas.

Further experimental and theoretical studies are being carried out.

ACKNOWLEDGEMENTS

The experimental assistance of Mr. P. Ward is gratefully acknowledged. The multi-channel U.V. recorder was kindly loaned by A.S. M. (Mullard) Ltd., Southampton.

REFERENCES

[1] W.G. STEWARD, J.A. BRENNAN and R.V. SMITH, "Transient flow of Cryogenic Fluids," NBS report 9709 (1968).

[2] R.B. BIRD, W.E. STEWARD and E.N. LIGHTFOOT, "Transport Phenomena" John Wiley and Sons H964).

200

DISCUSSION

A. SELLMAIER (W. Germany) — What can you say about the pressure fluctuations during the cooling down of the pipe. These fluctuations seem very impor­tant because the cooling channels of the cables have to be flexible in any case.

R. G. SCURLOCK — We did not make any pressure measurements in our experiments. There were definitely pressure fluctuations greater than the liquid nitrogen head at the beginning of every cool-down, causing blow back of the liquid nitrogen. However, we observed that they were considerably reduced once the inlet section was to 77 °K. These inlet fluctuations have been studied theoretically and experimentally at high mass flows in Ref. [1], and may be reduced by avoiding the use of sub-cooled liquids to start the cool-down. The possibility of fluctuations arising during the progress of an established cold front down a flexible cooling channel could be a much more serious problem. The conditions exist but are slightly different from those for inlet fluctuations owing to the inertia of the liquid behind the cold front. It is clear that the relationship between such pressure fluctuations and the mechanical modes of resonance of a long flexible line during cool-down with low mass flow of vapour or liquid requires some experimental and theoretical studies.

N. KURTI (U. K.) — Since cryo-cables are unlikely to be cooled down frequently is there much point in living to find the optimum conditions for this operation?

R.G. SCURLOCK — Yes. (1) for prototype test sections which will be fre­quently cycled between ambient and working temperatures. (2) For rapid recovery of working conditions in an operational cryo-cable following a fault.

Cooling cryo-cable and refrigerator together from ambient temperature will take a much longer time than the method outlined in this paper. Low mass flow of refrig­erant will be obtained during most of the cool-down, particularly at the higher tempe­ratures where the thermal capacity of the cable is large. Taking heat fluxes along the cryo-cable into account, it is obvious that the likelihood of zero delivery of "cold" must be avoided during any cool-down procedure when the mass flow is low.

It is therefore important to establish minimum (not optimum) requirements for cool-down, however infrequent, which will have to be incorporated in the refrigeration circuit.

201

LA REFRIGERATION DES APPAREILS CRYOfiLECTROTECHNIQUES

PROBLEME DES CONNEXIONS fiLECTRIQUES

J.M. LEROUX, S. LEHONGRE et E. CARBONELL Centre a"Etudes Cryogeniques de VAir Liquide, Sassenage {France)

Refrigeration of cryoelectrotechnical apparatus — Problem of electrical connections

SUMMARY: The heat that must be removed by a refrigeration cycle in a cold cryoelectro­technical machine comes from specific losses of the conductor, poor thermal insulation and the electrical connections between room temperature and the cold temperature.

The optimisation of the energy consumed by the refrigeration cycle as a function of amount of heat produced by the cryoelectrotechnical apparatus and the current passing through the connections shows the utility of producing the refrigeration at several temperature levels.

A certain number of configurations will be examined.

Dans un appareil cryoelectrotechnique, la chaleur a evacuer par le cycle de refrigeration provient, d'une part des pertes propres de l'appareil, d'autre part des defauts d'isolation thermique de l'enceinte cryogenique contenant l'appareil, enfin des connexions electriques entre la temperature ambiante et la temperature froide. L'importance relative de ces pertes depend de la nature de l'appareil et de sa taille.

Ainsi, par exemple, lors de la mise en oeuvre d'un bobinage supraconducteur de moyenne ou de grande dimension, les entrees de chaleur dues aux amenees de courant peuvent etre preponderantes, tandis que pour une cryoliaison de plusieurs kilometres les pertes par effets Joule dans le cable et les pertes par defauts d'isolation risquent d'etre les plus importantes.

Cependant, quel que soit le cas, il y a toujours interet a reduire le cout energe-tique des pertes par les connexions. Ce cout energetique est, rappelons-le, l'energie qu'il faut fournir a un refrigerateur reel pour evacuer la chaleur a une^temperature T[ l ] .

TYPES DE CONNEXION UTILISABLE

Considerons un appareil electrotechnique froid pour lequel la chaleur a prendre en charge par le cycle de refrigeration est due essentiellement aux connexions elec­triques : par exemple bobinage supraconducteur, parcouru par un courant de 10 000 A et fonctionnant a une temperature d'environ 4,5 °K.

Plusieurs types de connexions sont possibles; — Connexions sans recuperation (du type de celles etudiees en particulier par Mc-

Fee)[l]; — Connexions a recuperation totale optimale [1]; — Connexions a recuperation totale autonome; — Connexions optimum optimorum [2]; — Connexions a plusieurs troncons sans recuperation [2].1!

C O U T ENERGETIQUE THEORIQUE

Dans le cas particulier de connexions 300 °K — 4,2 °K realisees avec un metal obeissant a la loi de Wiedemann-Franz, et en considerant des refrigerateurs fonc-

203

tionnant suivant un cycle de Carnot th6orique, MM. Carbonell, Renard et Berard ont calculi [2] le cout erierg&ique de ces divers types de connexions. Les resultats font apparaitre qu'une connexion sans recuperation a un cout energetique de 3,3 W/A, tandis qu'une connexion en deux troncons optimaux sans recuperation a un cout energetique de 0,78 W/A, et une connexion optimum optimorum de 0,344 W/A.

COUT ENERGETIQUE REEL

Les calculs peuvent etre repris avec un metal reel et des rerrigerateurs r6els. Le metal est de 1'aluminium super-raffin6 presentant un rapport de resistivity de

1 400 entre 300 et 4,2 °K, dont on a mesure la resistivite et la conductivity thermique entre ces temp6ratures.

Le rendement des cycles de refrigeration utilis6 pour le calcul des couts energe-tiques est celui donne par le Professeur Kurti dans sa communication «Low tempera­tures in the generation and transmission of electric power»[3]. Nous allons calculer le cout 6nergetique d'une connexion en aluminium, optimisee pour faire passer (entree et sortie) 10 000 Amperes entre 300 et 4,5 Kelvin.

Dans la suite des calculs, nous designerons par Q^ la puissance evacuee a l'extr£-mite froide de la connexion et par Wth le cout energetique correspondant. — Connexion sans recuperation entre 300 et 4,5 °K :

Dans ce cas, Qf = 830 Watts (41,5 mW/A). En prenant, sur les courbes publiees par le Professeur Kurti, le rendement d'un

cycle entre 300 et 4,5 °K pour une puissance de 1 kW environ (rapport = 500), on a:

Wth = 415 kW

— Connexion a recuperation totale — autonome — Dans ce cas, Qf = 16,2 Watts. Au cout energetique correspondant a cette puissance evacuee a 4,5 °K, il

faut ajouter celui necessaire a la puissance evacuee par le gaz recuperant a tous les niveaux de temperature entre 4,5 °K et l'ambiante. Cela revient a considerer que, dans le cas de la connexion autonome, nous n'utilisons plus un refrigerateur pour fournir de la puissance a 4,5 °K, mais un liquefacteur, puisque le gaz est repris par le cycle frigorifique apres avoir 6te rechauffe jusqu'a temperature ambiante lors de son passage dans la connexion.

Les courbes donn£es par le Professeur Kurti ne peuvent plus alors etre utilises seules. II convient de comparer les puissances fournies a froid, pour une meme puissance consommee par le compresseur du cycle, selon qu'il s'agit d'un refrigera­teur ou d'un liquefacteur. Cette comparaison, basee sur des projets et des realisations industrielles, nous permet de donner un rapport 5 environ entre ces valeurs, c'est-a-dire que, pour une meme puissance P consomrree par le compresseur d'un cycle, il est possible de fournir a 4,5 °K Wf avec un liquefacteur, et Wr = 5 W, avec un refrigerateur, le rapport P/Wr etant donne par le Professeur Kurti.

Ces elements nous permettent desormais de calculer Wth dans le cas de la connexion autonome. S'il s'agit d'appareils cryoelectrotechniques, dont les connexions consti­tuent la principale source de chaleur a basse temperature (bobine supraconductrice par exemple), le rendement du cycle doit etre celui correspondant sensibleirent a 100 Watts en refrigerateur. Nous obtenons alors un rapport de 850 et Wffc = 69 kW environ.

Remarquons que la connexion autonome n'est pas a recuperation optimale. Cependant, les debits de gaz recuperants sont tres peu differents et les resultats restent valables.

204

— Connexion en trois troncons sans recuperation Cette connexion est pr£vue en trois troncons sans recuperation (type Mc Fee),

la refrigeration 6tant fournie a trois temperatures. A titre d'exemple, nous prendrons les temperatures usuelles 77 °K, 20 °K et 4,5 °K.

Les flux de chaleur qu'il faudra eVacuer, toujours en utilisant un conducteur en aluminium super-raffin6 dimensionne pour 2 fois 10 000 A, sont: r

.A 77 °K Q/i = 830 Watts

.A20°K Q / 2 = 118 Watts

.A 4,5 ° K | Q / 3 = 52 Watts

A Qfl9 Qf2, Q/3 correspondent des couts energ&iques Wffcl, Wfh2 et Wf/f3. Le cout energetique total est: Wth = Wrtl + Wffc2 + Wrfl3.

En reprenant, comme pour la connexion autonome, un rapport correspondant a un reTrigSrateur dimension^ pour 100 Watts a 4,5 °K, on obtient Wth = 830 x 7 + 118 x 70 + 52 x 850 en Watts, Wth = 58 kW environ.

COMPARAISON DES COUTS S N E R G E T I Q U E S

Nous r6sumons dans le tableau ci-dessous les resultats precedents et les comparons aux valeurs obtenues en utilisant un cycle de Carnot avec un m£tal obeissant a la loi de Wiedeman - Franz [2].

Connexion sans recuperation

Connexion a recuperation

Connexion en 3 troncons 300-77-20-4,5 °K

Aluminium super-raffing, cycle rdel {pour 20 000 A)

20,7 W/A

3,5 W/A (connexion autonome)

2,9 W/A

Cycle Carnot, mital suivant la loi de Wiedeman-Franz[2\

Wth = 3,30 W/A

Wth = 0,42 W/A (connexion a debit optimal)

Wth = 0,58 W/A

Nous remarquons que le type de connexion en trois troncons est le plus interessant du point de vue cout energetique. Comme du point de vue realisation une telle con­nexion est comparable a celle a recuperation totale, il est avantageux de l'utiliser, ceci d'autant plus que le niveau de77 °Kest aisement fixe par l'azote liquide bouillant sous la pression atmospherique et qu'un echangeur a azote liquide est tres facilement realisable. L'avantage des connexions en trois troncons sans recuperation s'amenuise lorsque la cryomachine electrique necessite une puissance froide nettement superieure aux entrees de chaleur par les connexions (casd'une cryoliaison supraconductrice par exemple).

REALISATION PRATIQUE

A titre d'exemple nous donnons un modele de realisation pratique qui com-prendrait : — Entre 300 et 77 °K, un conducteur optimise, sans recuperation, fait avec 1'aluminium defini precedemment. LI/S = 24 980 A/cm(*).

(*) L = Longueur en cm. S = Section en cm2. I = Intensite en Amperes.

205

— A 77 °K, un echangeur a azote liquide constitue par un tube enroule en helice autour du conducteur a l'interieur duquel l'azote liquide se vaporise (echange de chaleur en double phase).

— Entre 77 et 20 °K, un conducteur optimise sans recuperation. LI/S = 220 000 A/cm. — A 20 °K, un echangeur de chaleur dont le conducteur est bobine suivant la technique

courante des echangeurs bobines, bien connus dans les refrigerateurs et lique-facteurs. La section totale du metal est importante a ce niveau pour deux raisons essentielles : ne pas etre le siege de pertes Joule importantes, et avoir une grande surface d'echange avec le gaz de telle facon que l'ecart de temperature gaz-metal soit faible.

— Un troncon sans recuperation entre 20 et 4,5 °K; LI/S = 700 000 A/cm.

CONCLUSION

La comparaison des couts energetiques des differentes solutions envisagees montre que la connexion en trois troncons est la plus economique, ceci d'autant plus que la puissance installee (compresseur du cycle) est plus faible.

L'extraction de puissance a plusieurs niveaux de temperature est facile lorsqu'il s'agit d'un cycle de refrigeration du type Claude a deux turbines ou de Stirling a deux etages.

Enfin, on constate, d'apres les courbes publiees par le Professeur Kurti, que pour une puissance donnee, installee surle compresseur, 1'investissement est d'autant plus faible que la refrigeration est delivree a une temperature moins basse. La con­nexion en trois troncons sans recuperation permet done de gagner a la fois sur le cout d'investissement et sur le cout de fonctionnement.

REFERENCES

[1] P. BERARD : «Connexions 61ectriques cryog6niques. Aspects theoriques». I.I.F.-126 Cong. Int. Froid, Madrid (1967), I, pp. 157-171.

[2] E. CARBONELL, M. RENARD et P. BERARD: Thermodynamic optimum for electrical connections at cryogenic temperatures. Cryogenics (Octobre 1968).

r3] N. KURTI. «Low temperatures in the generation and transmission of electric power». I.I.F.-12e Cong. Int. Froid, Madrid (1967), I, pp. 1-13.

DISCUSSION

C. DAMMANN (France) — 1. Dans le cas d'une connexion a trois troncons que vous preconisez, quel est le flux thermique a la temperature de 4.2 °K?

2. Je pense que dans 1'etude economique des traversees de courant pour enceinte cryogenique, il faut prendre en consideration les conditions d'exploitation du systeme.

M. CARBONELL—1. Les flux thermiques aux trois temperatures 77, 20 et 4.5 °K sont donnes dans le texte de la conference. Ces valeurs sont:

— a 77 Kelvin 41,5 mW/A — a 20 Kelvin 5,9 mW/A — a 4,5 Kelvin 2,6 mW/A

2. II est bien evident que les connexions electriques en trois troncons doivent s'integrer a la technique generale du systeme. Cette solution semble bien adaptee lorsque le niveau de temperature de 77 °K est fixe, par exemple, dans le cas d'un cable supraconducteur, par l'interet economique de placer un ecran refroidi par circulation d'azote liquide.

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SURVEY PAPER

MOTORS, GENERATORS AND FLUX PUMPS

A.D. APPLETON Electrical Engineering Department, International Research and Development Co. Ltd.

Newcastle-upon-Tyne {United Kingdom)

Moteurs, generateurs et pompes a flux

RESUME : Ce rapport etudie Vetat technique et commercial des moteurs, generateurs et pompes a flux supraconducteurs. On fait Vhistorique de la mise au point des machines homo-polaires supraconductrices a VIRD en se referant particulierement a la conception et a la fabri­cation du moteur de 3250 ch qui doit etre essaye a la station electrique du CEGB a Fawley. Cette machine se trouve a un stade avance de fabrication et on presente un certain nombre de figures pour montrer ses elements. On discute des applications probables pour les moteurs supraconducteurs. On presente la conception generale des grands generateurs homopolaires supraconducteurs jusqu' a 200 MW, ainsi qu'un expose des besoins industriels en courant continu.

La revue de synthese rend compte d'autres progres connus des moteurs et generateurs supraconducteurs. On rend compte brievement de la conception, de Vapplication et des mises au point actuelles des pompes a flux.

Finalement on donne une vue des tendances futures des machines rotatives employant des supraconducteurs dans le domaine des progres technologiques et de Vinfluence du cout des materiaux supraconducteurs et des refrigerateurs a helium sur Vutilisation de ces machines dans VIndustrie electrique.

1 — GENERAL REMARKS

Over the past ten years or so, many studies have been made on the possible appli­cations of superconductors to electrical power equipment. It is probable that up to about a year ago most effort has been devoted to transformers with decreasing attention given to transmission, a.c. machines and d.c. machines respectively. In terms of the probability of success, both technically and commercially, this is precisely the reverse order in which resources should be allocated. It is often the case that new technological developments experience an initial grey period of commercial uncertainty because the cost of materials are not stabilized to a production envir­onment. We all wish to see superconducting transformers in production and large scale superconducting transmission systems under construction, but if investment in the development of superconductors and helium refrigerators is geared to the success of these applications progress will be slow. It is much more important to develop superconducting equipment which can be taken quickly and successfully to the large scale hardware stage; this will stimulate investment to reduce production costs of superconductors and refrigerators. Thus, although superconducting a.c. generators will eventually be more attractive commercially than superconducting d.c. machines, the latter must be developed at the present time.

The superconducting d.c. machines with which my paper is largely concerned, are now emerging from this grey area and their production may soon take place for a number of specific applications. This is most encouraging at such an early stage in their development, and indications are that future costs will improve to increase still further their commercial status. It is the opinion of many electrical engineers that d.c. machines should be avoided if at all possible and, as far as conventional machines are concerned, there are good reasons for this view. However, the advantages of superconducting d.c. machines are such that there will be a change in this attitude.

207

There appears to be no immediate prospect of obtaining a high field superconductor which will operate at power frequencies with acceptable losses. It is unlikely therefore that an a.c. machine will be developed in the near future with a superconducting armature; however, it may be quite feasible to produce a large a.c. generator with a superconducting excitation winding.

Most of the work on superconducting machines described in this paper has taken place at International Research and Development Co. Ltd., and this is the first occasion of its publication in any detail. Other known work on superconducting machines is described except where this has a security classification. A brief account is given of the status of flux pumps, but since much data are published already I feel justified in devoting most of the paper to machines.

2 — THE CONCEPT OF SUPERCONDUCTING d.c. MACHINES

2.1 CHOICE OF THE TYPE OF MACHINE

We may list a number of points relating to the use of superconductors in electrical machines and deduce the most appropriate type of machine. (a) The most important advantage of a Type 2 superconductor is its ability to carry

a high current density in a high magnetic field, and the first major consequence of this upon machine design is that we may break free from the magnetic limitations of iron. In fact, we have no choice since it may be shown that, in gereral, an iron magnetic circuit and a superconducting winding are mutually exclusive economi­cally. This is not to say, of course, that iron cannot be used for the magnetic screening of a superconducting winding.

(b) The most significant disadvantage of superconductors is that they will operate only at extremely low temperatures and this introduces a number of problems which are alien to machine designers. It must be appreciated, for example, that a logarithmic law applies to refrigeration power as a function of temperature. To remove heat at a rate of 1 W from a body at 4.2°K requires a refrigerator power of about 1400 W, while the corresponding figure for a body at 80°K is only about 20 W (these are practical values, the Carnot Specific Work values being much lower).

(c) A severe limitation of high field superconductors is that time varying magnetic fields produce losses which are prohibitive at power frequencies. The status of superconductors in this respect has improved dramatically over the past few years but power frequency operation is most definitely net possible at present.

Consideration of these factors, when our work at IRD commenced about 5 years ag o, produced the following conclusions: 1. The armature of the d.c. machine cannot be made superconducting because of

the losses in the superconductor due to time varying currents and magnetic fields. At the time that this decision was made this factor was even more relevant because superconducting windings could not give a guaranteed performance with direct current.

2. Less fundamental reasons for not adopting a superconducting armature are that the refrigeration power would be high due to heat flow along the shaft and because a vacuum seal on a rotating shaft would be required. These problems are more tractable at the present time than they were 5 years ago but they remain good reasons for avoiding a rotating superconducting winding.

3. The field coils of a heteropolar machine could be made superconducting, the problem areas being the mechanical forces between the coils, the effect of torque

208

reaction on the thermal losses of the coils, the effects of armature reaction and eddy current losses in the armature.

4. A heteropolar machine with superconducting field windings would be subject to similar restrictions in respect of commutation to those of a conventional machine, although the leakage reactance of the armature conductors would be lower in the absence of iron.

5. The disc-type homopolar machine with a superconducting field coil and an ambient temperature armature overcomes all of the problems found with the heteropolar machine. The drum-type homopolar machine, with two windings in opposition is attractive, but there remain the problems associated with the mechanical force between the coils. At the present time these problems may be completely overcome and we have a free choice between the disc and drum-type homopolar machines.

2.2 THE DISC-TYPE SUPERCONDUCTING HOMOPOLAR MACHINE

The basic features of a disc-type superconducting homopolar machine are shown in figure 1. It utilises a single field winding enclosed within a cryostat to minimize

Faraday disc \

Superconducting winding : 4.4 K

Thermal radiation shield : 80°K

Ambient temperature vessel ~ 300°K

Current col lection

Armature field

Machine terminals

Torque reaction disc

Fig. 1 — Basic features of disc type superconducting homopolar machine.

209

the heat leak to the winding by conduction and radiation. The magnetic field produced by the superconducting winding is a maximum at the inside surface of the winding and, in the bore, falls to a minimum on the axis. The armature consists of a Faraday disc with sliprings near the shaft and at the outside of the disc. Means for collecting the current at the sliprings are provided, and a stationary disc takes the current radially between the outer slipring and the machine terminals. The useful magnetic field of the machine is that which cuts the Faraday disc between the sliprings, and figure 2 shows why it is important for the disc to occupy as much of the bore as possible. A small increase in the diameter of the Faraday disc produces an increasing increment of useful flux as the diameter is increased. In the space between the field coil and the Faraday disc there must be accommodated the coil former, a vacuum space, a radiation shield, a further vacuum space, the thickness of the ambient temperature casing of the cryostat and a clearance gap for the rotating armature. Using conventional type graphite brushes for current collection, there would be a substantial loss of useful flux if these were arranged in the conventional manner. To avoid this loss the slipring is designed with an over-hung lip and the slipring is at the underside, i.e. the brush pressure is radially outwards (fig. 1).

Superconduct ing wind ing

20 40 60 80 100$ Percentage of total flux through bore of winding

Fig. 2 — Useful machine flux.

The stationary disc (fig. 1) carries the armature current through the same magnetic field as the Faraday disc and therefore experiences the same torque as the latter. In other words, the stationary disc takes the torque reaction of the machine and none

210

of this force appears on the field winding. It may also be noted from figure 1 that the magnetic field produced by the armature current is in the azimuthal direction between the Faraday disc and the torque reaction disc. This flux is in quadrature with the excitation field and does not give rise to any armature reaction.

Thus, the superconducting winding is not subjected to any mechanical or magnetic influences, precisely the situation required.

2.3 THE DRUM-TYPE HOMOPOLAR MACHINE

The basic features of a drum-type superconducting homopolar machine are shown in figure 3. The working conductor is now a drum and the armature current flows in the axial direction. The sliprings are now at the same diameter and located

Superconducting winding

Radiation shields

N

Current col lection

\ / /

Fig. 3 — Basic features of drum type homopolar machine

near the mid plane of each coil. The torque reaction is taken on a stationary cylinder co-axial with the armature. Means must be provided to accommodate the large force between the windings, and this is most readily achieved if they are contained within a common cryostat.

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2.4 DISCUSSION ON THE CONCEPT OF SUPERCONDUCTING d.c. MACHINES

The choice of the homopolar machine remains the most attractive for super­conducting d.c. motors and generators, and it is most unlikely that the heteropolar type will be able to displace it. There are two major problems associated with homo-polar machines; efficient current collection and low generated voltage. It will be shown later that the performance of the current collection system has a profound effect upon the optimization of a design for a given rating. The problem of low voltage has been overcome in a manner which will be described later in this paper.

A further problem associated with superconducting homopolar machines is that the magnetic field is not contained within iron and the effect of the stray field must be given careful attention. Of major interest are the advantages which superconduc­ting d.c. motors and generators have to offer over conventional d.c. machines and this is discussed later in this paper.

3 — A N ACCOUNT OF THE DEVELOPMENT OF SUPERCONDUCTING d.c. MACHINES AT IRD

3.1 INTRODUCTION

Figure 4 shows the programme on superconducting machines at IRD over the last 5 years.

In 1963, IRD commenced an investigation into the possibility of employing super­conductors in an electric motor with the object, inter alia, of achieving a high power to weight ratio. This work was undertaken on behalf of the Ministry of Defence (Navy). The status of superconductors at that time left much to be desired, the major problem being the well known degradation effects made it very difficult to justify the construc-

FeasibiIity study

Model motor

Research on current collection

Studies of commercial aspects

I Design of Fawley motor

Manufacture of Fawley motor

Testing of Fawley motor at IRD

Testing of Fawley motor at Fawley

I Production

[Studies on dc generators

Studies of commercial aspects

-^Design of dc generator

Manufacture of dc generators

I Product ion

Fig. 4 — Research and development programme on superconducting machines at IRD.

212

1963 1964 1966 1967 1968

I Initial testi ng

tion of large superconducting magnets. The feasibility study was completed in August 1964 and our findings were that it was possible to produce a d.c. motor with a high power to weight ratio through the use of superconductors. Of course, the designs produced assumed that large superconducting windings could be constructed and the decision to proceed to the hardware stage required a careful assessment of the assumptions which had been made. It was decided that it was reasonable to expect that the technology of superconductors would improve and that a small demonstration superconducting motor should be designed and constructed. The work commenced on 1 January 1965 with a programme to be completed in August 1966 and it was anticipated that a machine with a rating of about 2 hp would result. The programme required development work, particularly on current collection, to proceed in parallel with design and construction in order to meet the time scale. On 1 June 1966 the superconducting motor was first operated, and the test programme was completed in July by which time the design improvements were such that the maximum power developed by the motor reached 50 hp at 2000 rev/min. The increased power over that anticipated at the beginning was due to a number of design improvements parti­cularly with the superconducting winding, and the machine came to be known as the model motor.

At this time, we decided to investigate the commercial possibilities of the motor as distinct from its military aspirations and in late 1966 the support of the National Research Development Corporation was sought to design and construct a large machine with particular reference to a drive for a steel rolling mill. The NRDC set up a working party of eminent persons and, after a number of meetings, it was unanimously decided that we should proceed with the construction of a motor of about 1000 hp. However, the NRDC decided that this machine was too small and asked us to consider a machine of about 3000 hp at about 200 rev/min. We were anxious to test the motor in an industrial environment and, after discussions with the CEGB, it was agreed that the motor could be installed and tested at a new power station at Fawley near Southampton. On 1 May 1967 we commenced the design and development of this machine, and manufacture is now nearing completion; it will be installed at Fawley later this year. I must emphasize that the motor is only at Fawley for trials and will be taken out after about 18 months; we are most grateful to the CEGB for their co-operation in this venture.

In the last year or so we have been giving increasing attention to large super­conducting d.c. generators and designs up to 200 MW at 1000 V have been produced. We believe that, in addition to superconducting motors, superconducting generators have an important role to play in industry.

3.2 THE MODEL MOTOR

It is not necessary for me to describe this machine because it is the subject of a separate paper [1] at this conference. Although this machine, shown in figure 5, is now completely obsolete it has played a vital role. Work on it commenced at a time when it was impossible to be certain that large superconducting magnets could achieve engineering status, and much credit is due to the judgement and support of the Chief Scientific Advisor to the Director General, Ships. Had the model motor not been constructed, it would have been impossible to make a sound case to the NRDC for the construction of the Fawley motor.

3.3 THE FAWLEY MOTOR

The cryogenic aspects of the motor and the design of the superconducting winding are covered in two other papers [2, 3] to be presented at this conference and it is therefore unnecessary for me to say very much on these matters.

213

The machine is based upon the disc-type homopolar principle with a single field coil which contains 5^ tons of a fully stabilized Nb-Ti/copper composite manufactured by IMI. The coil is maintained at its operating temperature of 4.4°K by a helium refrigerator arranged in a closed circuit; the refrigerator is manufactured by British Oxygen Cryoproducts Ltd. Figure 6 shows the fully wound coil and figure 7 shows the outer vessel of the cryostat.

The useful magnetic flux of the machine, i.e. the flux which cuts the rotating conductors between the inner and outer sliprings is 6.45 Wb and, therefore, at 200 rev/ min the voltage developed by a simple Faraday disc is about 21.5 V. The model

Fig. 5 — Model superconducting motor 50 hp 2 000 rev/min (first operated at IRD in June 1966).

214

motor was arranged to have a Faraday disc on each side of a stainless steel strength disc and if this arrangement had been adopted for the Fawley motor its power requi­rements would have been 43 V 58,000 A. It was decided that this was unsatisfactory and an alternative design was sought; the possibility of using a large number of discs was rejected because of the severe mechanical complications and maintenance problems. It was required to keep the working conductors as close as possible to the mid-plane of the superconducting winding for two reasons; simplicity of design and to employ the magnetic field to maximum advantage. An obvious step is to divide the Faraday disc into a number of segments which are insulated from each other and each of which has a slipring segment at the inner and outer radii. However, the problem is to find a way of arranging the brushes on the slipring segments so

Fig. 6 — Superconducting coil of Fawley motor (before welding of side plates of coil former).

215

that they do not cause short circuits and such that the voltage between segments does not exceed about 20-30 V. The principle of the method by which this is achieved is shown in figure 8. The brushes are arranged on alternate segments and connected as shown to achieve a series circuit. When one set of brushes are on one slipring segment, the slipring segments on either side are not connected to any brushes and are not working. When the rotor moves such that one set of brushes is making contact with two adjacent slipring segments, the current is shared between the conductors which join the slipring segments at the inner and outer radii. The system involves, therefore, the switching of current from one set of rotor conductors to the next, and the efficiency with which it does this is dependent upon the leakage inductance of the conductors. Theoretical calculations and tests on the model motor (which was modified to the segmented system) have indicated that the leakage inductance is much below the threshold where sparking may be expected. If the conductors were arranged as shown in figure 8, there would be a high voltage between two of the segments, and therefore

Fig. 7 — Outer vessel of cryostat for Fawley motor.

216

Stator bars

Rotor bars

Large diameter "slip ring

SmaI I diameter slip ring

0 e 2e 3e (a) Brushes entirely on one segment

4e 5e

0 e 2e Cb) Brushes on two segments

3e 4e

Fig. 8 — Principle of segmented rotor.

Fig. 9 — Series—Parallel arrangement of segmented rotor.

a modification shown in figure 9 is adopted. This is a series-parallel arrangement where the maximum voltage between adjacent segments is limited to that generated in one conductor. Figure 10 shows the strength disc of the Fawley motor and the milled slots to house the hollow water cooled conductors which connect the slipring segments. To reduce leakage inductance each slot contains two conductors which are joined to adjacent segments. Figure 11 shows the rotor of the Fawley motor partially assembled. A number of variations on this design are possible which improve the utilization of the slipring. The total back e.m.f. of the Fawley motor is 430 V at 200 rev/min and the supply current on full load is 5800 A.

Fig. 10 — Stainless steel strength disc and shaft of Fawley motor.

218

The motor power is supplied from a combination of rectifiers and silicon controlled rectifiers which permit the voltage to be varied over the full range. The motor is equipped with a closed circuit speed control system. It is necessary to say a few words about current collection although this will be covered to some extent by the paper on the model motor. We have been investigating the performance of high copper content graphite brushes for a number of years and these will be employed on the Fawley motor. It is possible to effect considerable reductions in the cost of superconducting machines if the specific current loading of the sliprings can be increased, and to achieve this we are developing a new method of current collection. However, until the devel­opment work is completed it is not possible to publish any details.

Fig. 11 — Partially assembled armature of Fawley motor.

219

It is expected that the Fawley motor will be assembled and ready for test at IRD in June this year and it should be installed and operating at Fawley in October or November. It should be mentioned here that the pace of development is such that, now, about 8000 hp could be obtained from the same frame size as the Fawley motor.

Before proceeding to a discussion of economics it is relevant to make some com­parisons between conventional and superconducting d.c. motors.

4 — OTHER WORK ON SUPERCONDUCTING d.c. MACHINES

A number of establishments are now engaged upon studies involving super­conducting homopolar machines, but I do not know of any which have reached the hardware stage.

In France, at Laboratoire Central des Industries Electriques at Fontenay-aux-Roses, work has been in progress for a number of years on iron-cored homopolar machines using mercury as the means for current collection. At this laboratory it is planned to construct a homopolar machine with a superconducting winding of about 0.5 m diameter. The armature consists of a number of discs which will run totally immersed in mercury.

In the USA, AVCO have conducted studies on superconducting homopolar machines with a number of Faraday discs in series. The current is collected at the sliprings by means?of^ sodium potassium (NaK).

In Japan, at Tokyo Shibaura Electric Co. Ltd., it is planned to build a supercon­ducting homopolar generator with armature discs of 1.2 m diameter and using silver graphite brushes. Dr. Yamamato is presenting a paper [4] at this conference on his work.

5 — DISCUSSION ON THE APPLICATION OF SUPERCONDUCTORS TO d.c. MACHINES

5.1 THE LIMITATIONS OF CONVENTIONAL D.C. MACHINES

(a) Heteropolar

Almost all d.c. machines are heteropolar because of the well known limitations of the homopolar type. To avoid any difficulties which may arise from the varying practices of manufacturers, the following design parameters are based upon published data [5,6},

Consider an armature of diameter D metres, length L metres, maximum air gap flux density B^ Wb/m2

9 armature conductor current, Ia, total armature conductors Zfl, and ratio of average flux density over a pole pitch to maximum air gap flux density Kf.

In one revolution the total flux cut by a conductor is given by c|> = TT-D-L-Ky-B^ (Webers); the work done in one revolution is <\>-Za-Ia (watt seconds).

Defining specific electric loading as q ampere conductors per metre of circumference we have Ifl• Za = TtDq and the work done per revolution becomes D 2 L (n2 • q• K^• B^).

If the machine power is P kW and the N rev/min

- = D 2 L ( 0 . 1 6 4 < r K , B f l 1 0 ~ 3 ) N

The term in brackets is called the Output Coefficient (ESSON) and its numerical value does not vary widely for a given class of machine. The value of B^ is typically about 1 Wb/m2; Kf is about 0.7 and a high value of q for large machines is about

220

50,000. These values yield

- = 6D 2 L N

The limiting values of D and L are determined, inter alia, by peripheral speed and voltage between commutator segments.

The output coefficient of 6 applies to large diameter machines e.g. D = 4 m and falls with smaller diameters. Figure 12 shows the maximum outputs which are possible with conventional machines as a function of speed. The outputs are derived from published data which is a few years out of date, but it is understood that present day

4

/

J /

Conventional d.c. reachines

A / c ^ \ >

s **"

Supercondwct i ng motors and generators ^ \

kD^-L-26 mJ

^Double armat ure macnines used in this regior

Convent iona1 d.c. motors

10 IU0 IU00 10000 Speed, rev/min

Fig. 12 — Maximum range of conventional and superconducting d.c. machines.

capabilities are not much different from that indicated on figure 12. The outputs may be increased perhaps by 60% if duplex lap windings are employed, but some manufacturers regard these as troublesome. The maximum power is in the region of 10 MW at speeds of a few hundred rev/min. At lower speeds the machines become very heavy, i.e. in the hundreds of tons bracket, and double armature machines are employed for certain ratings.

221

5.2 THE LIMITATIONS OF SUPERCONDUCTING HOMOPOLAR MACHINES

It is not possible to produce an output coefficient for a superconducting machine in the relatively simple form of the conventional machine.

The armature conductors of either a disc or drum-type homopolar machine will cut a total flux § in one revolution (corresponding to Tt-D-Ky-B^ for conventional machines). However, there is no such simple relationship between (() and armature diameter in a superconducting machine. The specific electric loading in a conventional machine refers to ampere conductors in the armature, but for superconducting machines it is more useful to define q as the total slipring current per metre of circum­ference.

The work done per revolution becomes

P _ ( f r - t t - D - g x K T 3

N " 60

Since (() is defined as the total flux between the sliprings of either a disc or drum-type machine, this expression applies only to a machine with one pair of sliprings.

A more general expression is

_P _ s - ( | ) - 7 f D - g x l ( r 3

N ~ 60

where s is the number of pairs of sliprings. This expression is unaltered if the machine is segmented, provided that q is properly

calculated to make allowance for that part of the slipring which is not utilized. It is not proposed to give a detailed account of the possible range of values of

the product (|)D, but it varies up to about 40 for large diameter machines. Using conventional high copper content graphite brushes, the value of q may be as large as about 30,000 depending upon slipring velocity and length of slipring which is allowable. If liquid metals are employed, q may exceed 100,000.

The highest practical value of P/N using conventional brushes is probably between 60 and 100, but it is anticipated that this may be increased to about 300 without recourse to liquid metals. Figure 12 shows the range of outputs for superconducting motors and generators assuming P/N to be about 200 for speeds up to about 300 rev/ min and thereafter falling to about 150 at 1000 rev/min.

5.3 DESIGN ASPECTS OF SUPERCONDUCTING MACHINES

Commercial considerations make it impossible to publish design and performance data in any detail but a few general comments may be made.

5.3.1 Superconducting winding The Fawley motor employs a Nb-Ti/copper composite rated at 4000 A/cm2

(about 3000 A/cm2 overall). Magnets for use in high energy physics, e.g. quadrupoles, require rather higher current densities and the design problems are more severe. In the electrical machines described, very high current densities are not essential and in common with most electrical power plant the windings are optimized against minimum cost. Figure 13 shows the relationship between the cost of a superconductor £/A km as a function of magnetic field; this curve reflects the short sample charac­teristic and shows the penalty of working at very high magnetic field strengths. When, say, Nb-Ti is produced in the form of a copper composite, the cost analysis is rather more complicated and is a function of both current and magnetic field. For example,

222

a composite to work at 3 Wb/m2 and carry 100 A would differ from one to work at 6 Wb/m2 carrying the same current in that more superconductor is required, but, to a first approximation, the amount of copper and fabrication cost would be un­changed. There are a number of advantages in using hollow composite conductors which are cooled with helium gas. The newly developed intrinsically stable multi-filament superconductor is important for superconducting generators because it enables the excitation to be changed rapidly. Niobium-tin offers immense possibilities,

Present price

Range to allow much greater expansion

_l_ ~ 2 3 4 5 6

Magnetic f ie ld Wb/m2

Fig. 13 — Specific cost of Nb-ti superconductor.

although it is too expensive at present to be considered for electrical machines; it is very attractive in that it may be operated at 10 or 11°K with a reasonable current density and this reflects in low refrigeration costs. Aluminium stabilized supercon­ductors have the advantage of low weight and the improved designs that are possible because of the higher conductivity of aluminium. However, the bond to the super­conductor is not as good as with a copper composite, although evidence has been put forward that the bond is adequate and the specific cost (£/A km) is very com­petitive.

5.3.2 Refrigeration The capacities of helium refrigerators now available commercially are more than

adequate for the largest superconducting machine. It would be of great benefit if the physical size of the plant could be reduced and there are good reasons to believe

223

that this will be achieved with very compact heat exchangers and high speed rotary compressors. Of utmost importance is the reliability of helium refrigerators, and it is desirable for the design to be such that maintenance can be undertaken in a few hours: if the latter is possible, the electrical plant may be kept operational by means of a storage dewar.

5.3.3 Other design points There is no doubt that improved methods of current collection will provide the

means for achieving reduced costs for superconducting homopolar machines. Research in this area is in progress in a number of laboratories and it is reasonable to expect that advances will be made.

Stray magnetic fields of superconducting machines are of more consequence for some applications than others; it is highly undesirable to use large quantities of iron, and alternative methods of screening have been developed.

6 — ECONOMICS AND APPLICATIONS OF SUPERCONDUCTING d.c. MACHINES

6.1 GENERAL

Economics dictate that an electrical machine must have either an iron magnetic circuit excited by a normal winding or a superconducting winding with the magnetic circuit substantially free of iron. We have shown in the model motor described in another paper [1] that small quantities of iron can be used to advantage, but, in general, superconducting machines have an open magnetic circuit. If an attempt is made to design a machine without iron and without using superconductors, the power consump­tion of the excitation winding is enormous. Consider the Fawley motor with an average current density of 4000 A/cm2; the following table shows the consequence of using a copper winding.

Current density A/cm2

4 000 1 550

620 155

WT. of copper Tons

5.1 13.7 36.7

383.0

Useful flux ((> Wb

6.46 5.43 4.40 2.89

Power loss kW

14,000 5,700 2,400

713

It may be seen that, if the current density is reduced in an attempt to reduce the power consumption, the weight of copper becomes high and the machine flux is reduced to useless levels. If we attempt to use a normal conductor at reduced tempera­tures, for example at the boiling points of nitrogen (77°K) or hydrogen (20 °K), the situation is no better. The total power loss is the joule heating in the winding and that required by the refrigerator, being 30 MW for liquid nitrogen and 20 MW for liquid hydrogen.

Figure 12 shows that superconducting machines may be produced with ratings up to at least 200 MW, far in excess of the capabilities of conventional machines, and new applications may then be considered.

6.2 SUPERCONDUCTING MOTORS

It is not possible for commercial reasons to give a detailed account of the econo­mics of superconducting motors, but clearly there is a lower rating where they cannot

224

compete with conventional machines. If we consider the low speed range say up to 100 rev/min, figure 12 shows that conventional machines may be made with ratings approaching the limits of superconducting machines, but their weigths are large compared with the latter. For example, at 50 rev/min the limit on a conventional machine is about 7.5 MW, but the armature volume is 26 m3 and the total weight of the machine is in the hundreds of tons bracket. A superconducting motor of the same rating would weigh about 30 tons and the capital cost is lower.

In 1965 the total value of d.c. motors sold in the UK, with ratings above 1000 hp, was £ 5.4 M; most of these machines run at less than 500 rev/min and a good propor­tion at about 100 rev/min. The applications which may be considered for super­conducting d.c. motors are as follows:

6.2.1 Drives to steel rolling mills

Large rolling mills are usually driven by separately excited d.c. motors in a Ward-Leonard arrangement, but, increasingly, rectifiers are being employed to supply the motors. This trend is developing to reduce maintenance costs and to increase efficiency. The superconducting motor is now capable of being designed to have the same power supply requirements as a conventional motor and therefore the power supply may be excluded from any cost comparison.

The superconducting motor has a low inertia, low armature inductance and zero armature reaction; furthermore, it is capable of sustaining very severe overload conditions. These factors are distinct advantages for most mill applications. The weight of a large slow speed superconducting d.c. motor is substantially lower than its conventional counterpart, i.e. in the range 20-40 tons compared with several hundred tons; in addition, the superconducting motor is smaller physically.

The conclusion drawn from a study of this application is that the potential of the superconducting d.c. motor is quite large.

6.2.2 Ship propulsion

The superconducting d.c. motor may be produced in ratings which are suitable for ice breakers and ferry boats; for these applications they are expected to show clear advantages over alternative propulsion systems particularly in respect of manoeu­vrability. Other possible applications are deep sea tugs and special purpose ships such as cable layers. Detailed studies are in hand and our findings will become known at a later date.

6.2.3 Large winders

In various parts of the world, mines are being worked at increasing depths, and large winders with a wide range of speed control would be advantageous. Supercon­ducting d.c. motors would be suitable in these applications.

6.2.4 Other applications

Other applications which are expected to be relevant to superconducting d.c. motors are:

aluminium rolling mill drives compressor drives (high and low speeds) special purpose drives, e.g. centrifuges. In the longer term when costs fall it is expected that the range of applications

will be extended to include, for example, large auxiliary drives in power stations.

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6.3 SUPERCONDUCTING GENERATORS

The availability of large d.c. generators for the first time is expected to have a significant impact on certain sections of industry. Important applications for d.c. power are found in the electro-chemical industries, for example, aluminium smelting and chlorine production. The growth of production in both of these industries is about 10% per annum and their combined world requirements for new plant are about 1500 MW/annum. The superconducting generator thus appears as a competitor to transformer rectifier systems and can show substantial savings in capital cost parti­cularly on a green field site, i.e. where mechanical power is provided for generation. There are numerous other applications of superconducting generators.

6.4 COMMENT

Detailed studies of the economics of superconducting machines have shown that there is a growing market where they are competitive with conventional motors and other methods of producing d.c. power. At such an early stage in their develop­ment, this is an exciting situation because it is expected that the next few years will bring substantial reductions in costs.

7 — SUPERCONDUCTING a.c. MACHINES

It is not possible to draw many conclusions at the present time on the potential of superconducting a.c. machines, particularly large generators. There is no prospect in the near future of a superconductor which will permit a superconducting a.c. armature. To achieve the present efficiency of a 500 MW turbo-generator, the a.c. losses in the superconducting armature winding would have to be limited to about 5 kW because of the refrigerator power. This limit is probably orders of magnitude away from the present performance of superconductors at power frequencies.

The only possible alternative is to consider an a.c. machine in which only the excitation winding is superconducting. A number of countries are now engaged in studies on this type of machine, but I do not have clearance to describe the work.

A completely superconducting alternator was described [7] at the Cryogenic Engineering Conference at Stanford in 1967. This employed niobium tin super­conductors for both armature and field windings and was designed to operate at 24,000 rev/min. The machine achieved an output of 6.1 kW, but there was some doubt about the cooling of the armature. To the best of my knowledge there have been no further developments on this type of machine which was developed for airborne duty.

8 — FLUX PUMPS

Flux pumps are based on the flux storage capabilities of superconductors, and depend upon flux transfer for their operation as sources of large d.c. currents. These currents can be generated using either mechanical or a.c. inputs. The main interest in flux pumps has been promoted by the hope for avoidance of thermal losses associated with electrical input leads to power superconducting solenoids in a cryo­genic environment.

The operating principle of a flux pump is illustrated by figure 14. This closed circuit is formed of a high field superconducting coil connected to a plate of low field superconductor. As the permanent magnet is rotated across the low field super-

226

conducting sheet, it drags a normal spot and a portion of trapped flux is added to the superconducting loop. This additional flux increment generates a circulating current increment. During the completion of the cycle, the magnet rotates past the high field superconductor which is not driven normal. If the magnet's direction of rotation is reversed, the sense of the induced current increments are also reversed, while varying the strength of the permanent magnet alters the flux trapped in the normal region and thus the magnitude of the current increments.

Low field superconducting sheet

NormaI spot

Superconducting Joints

High field superconductor

Rotating flux source (permanent magnet)

Fig. 14 — Operating principle of homopolar superconducting d.c. generator.

This model can be taken a step further by connecting a number of plates in series and rotating a number of permanent magnets past them using a mechanical drive. Wipf [8] described such an example (fig. 15) and has constructed a system using electromagnets connected to a three phase supply (fig. 16).

A double switch flux pump is another similar device in which the flux is first admitted to a superconducting ring which is then closed to one circuit and opened up to the second circuit. The purpose of the two circuits is a gradual flux pumping where the second circuit's inductance is much higher than the first. Buchhold [9] has devised an electrical flux pump based on this principle which is a variant of Olsen's [10] balanced rectifier circuits. It uses a power cryotron and saturable reactors to form a superconductive power supply with no moving parts.

Fig. 15 — A multi-segment mechanically driven gliding-spot flux pump (after Wipf).

227

The advantages of Buchhold's design are: (1) large currents are never introduced into the cryogenic region; (2) the pump uses small currents; (3) when pumping stops, the current persists.

This type of system is being marketed with power outputs up to about 50 W and efficiencies of 97%, and is useful for the slow charging of superconducting magnets, while rotating flux pumps have poor efficiencies, but are useful in situations of fine control on high current and low power.

Fig. 16 — A 3-phase a.c. electromagnetically driven gliding-spot flux pump (after Wipf).

Although flux pumps have made good progress and output powers have risen from microwatts to tens of watts, it seems that improvements in both power and efficiency are needed to expand their range of application. A list of references [9-15] is included in this paper for those who wish to explore this subject in greater depth.

9 — CONCLUDING REMARKS

1. In this paper I have attempted to establish the status of superconducting d.c. machines. The fact that I am not free to develop a more detailed account of their commercial viability is an indication that this is substantial.

2. A study of this paper will show the importance of current collection in the cost of the machines; a factor of 2 improvement on the capabilities of conventional brushes on high speed sliprings may effect some 20% saving in capital cost. It is anticipated that improved methods of current collection will be available in the near future.

3. The pace of development of superconducting materials over the past few years has been excellent, and as far as d.c. machines are concerned it is now most important to concentrate on the reduction of production costs.

4. A major theme of my paper is that by making rapid progress with the production of viable d.c. machines we shall advance the development of a.c. machines.

5. It is important to note that the development of superconducting materials for essentially one-off magnets for high energy physics experiments has been of the greatest value to the electrical power industry. As far as the work at IRD is concerned, the importance of the role of MoD (N) in advancing the status of superconducting machines cannot be over-emphasized.

6. It is highly desirable that the physical size of helium refrigerators be reduced; there are a number of indications that this will become possible and there exists a commercial incentive for manufacturers to achieve this as rapidly as possible.

7. Finally, I wish to say that the subjects under discussion at this conference are of major significance to the Electrical Power Industry and that I expect to see an exponential increase in activity in the field.

228

ACKNOWLEDGEMENTS

I wish to thank the MoD (N), the National Research Development Corporation Ltd. and the directors of International Research and Development Co. Ltd. for permission to publish this paper.

I wish to thank the CEGB, particularly the Southern Projects Group, Marchwood Engineering Laboratory and the staff at Fawley Power Station, for their co-operation in respect of the testing of the Fawley motor.

Finally I wish to acknowledge the assistance given by C.A. Parsons and Co. Ltd. in the manufacture of the Fawley motor and their support of our work on generators. Our thanks are also due to the following companies who have been engaged in the manufacture of components for the Fawley motor: Imperial Metal Industries (Kynoch) Ltd.; British Oxygen Cryoproducts Ltd.; Morfax Ltd.; K & L Precision Engineers Ltd.; Thomas Bolton & Sons Ltd.; W.P. Butterfields Eng. Ltd.; System Compu-tors Ltd.; BICC Winding Wires Division; Head Wrightson (Teesdale) Ldt.; Enfield Rolling Mills; Alfa-Laval Ltd.; Mather and Platt Ltd.; Flexibox Ltd.; Morganite Carbon Ltd.; Hackbridge and Hewittic Electric Co. Ltd.

REFERENCES

[1] A. D. APPLETON and R. B. MACNAB, A model superconducting motor. Commission I London, Annex 1969-1 Bull. I.I.R., pp. 261-267.

[2] A. D. APPLETON and J. S. H. Ross, Aspects of a superconducting winding for a 3 250 hp motor. Commission I, London, Annex 1969-1 Bull. I.I.R., pp. 269-275.

[3] F. TINLIN and J. S. H. Ross, The cryostat and refrigerator for a 3 250 hp superconducting motor. Commission I, London, Annex 1969-1 Bull. I.I.R., pp. 277-283.

[4] M. YAMAMOTO, Semi-superconductive rotary machine. Commission I, London, Annex 1969-1 Bull. I.I.R., pp. 285-289.

[5] A. E. CLAYTON, (3rd ed. rev. by N. N. HANCOCK) The performance and design of direct current machines. Pitman (1959).

[6] L. GREENWOOD, Design of direct current machines. MacDonald (1949). [7] G. J. OBERHAUSER and H. R. KINNER, Some considerations in the design of a super­

conducting alternator. Adv. Cryog. Engg, 13, Plenum (1968). [8] S. L. WIPF , A superconducting direct current generator. Adv. Cryog. Engg, 9, Plenum

(1964), pp. 342-348. [9] T. BUCHHOLD, Superconductive power supply and its application for electric flux pumping.

Cryogenics, 4, No. 4 (Aug. 1964) pp. 212-217. [10] J. L. OLSEN, Superconductive rectifier and amplifier. Rev. Sci. Inst., 29, (1958), p. 537. [11] J. VOLGER and P. S. ADMIRAAL, A dynamo for generating a persistent current in a super­

conductive circuit. Phys. Lett., 2, No. 5, (Oct. 1, 1962), pp. 257-259. [12] S.L. WIPF, Flux pumps as power supplies for superconducting coils. Proc. Int. Symp.

on Magnet Technology, CFSTI (1965), pp. 615-624. [13] J. VAN SUCHTELEN, J. VOLGER and D. VAN HOUWELIGEN, The principle and performance

of a superconducting dynamo. Cryogenics, 5, No. 5 (Oct. 1965), pp. 256-266. [14] D. L. ATHERTON, Practical aspects of homopolar superconducting d.c. generators.

Cryog. Engg. News (Oct. 1968). [15] D. L. ATHERTON, Superconducting d.c. generators and motors. fEEE Spectrum, 7, No. 12

(Dec. 1964), pp. 67-71.

DISCUSSION

H. LONDON (United Kingdom) — Do the magnetic field oscillation which occur in the rotor of the Fawley motor when the individual sectors switch from the idle to the current carrying mode have a detrimental effect on the field coil or its container due to eddy current heating?

A.D. APPLETON — The time varying magnetic fields due to the switching of current from one segment to the next is very localized; the current in one slot of the rotor is essentially constant and there is no detrimental effect on the field winding. Large

229

changes in load current do not affect the field coil because the armature field is in quadrature with the excitation field.

P. A. GOEMANS (The Netherlands) — What is the effect of the stray magnetic field of the motor?

A.D. APPLETON — The magnetic field of the motor is not contained as in a conventional machine and its effects must be carefully considered for each application. For a steel rolling mill drive the motors are often located in a separate room and the stray field is not too troublesome. On a ship, however, it is necessary to control the stray field and we have developed methods of meeting this problem.

R.G. SCURLOCK (United Kingdom) — Since a large part of the refrigeration is required to cool the current leads have you considered using a persistent current switch ?

A.D. APPLETON — About one quarter of the refrigerator capacity is employed to cool the current leads. This is significant but it does not mean that the refrigerator is 25% more expensive. The use of a persistent current switch may reduce the cost by about 10% but this is not sufficient to justify the risk of being unable to dump the stored energy of the winding in a safe manner. In my opinion persistent current switches are not relevant to large superconducting windings at the present time. In the event of a situation requiring the stored energy to be dumped the use of a persistent current switch would require the leads to be inserted, the switch must then be opened and hold off the discharge voltage.

D.R. EDWARDS (United Kingdom) — Mr. Appleton referred to future 200 MW d.c. generators and also to a requirement for an optimised design of refrigerator having a capacity of 50 W at 4.2°K. Are these two figures compatible?

A.D. APPLETON — Yes. A 200 MW d.c. generator will not require a refrigerator of a capacity greater than about 50 Watts, remember that generators are high speed machines and the physical size for 200 MW is the same order as the Fawley Super­conducting Motor.

N. KURTI (United Kingdom) — Would Mr. Appleton hazard a guess of the likely market for superconducting motors and generators in 5 or 10 years' time and of the corresponding demand for superconducting materials?

A.D. APPLETON — This is an impossible question to answer but it is important to make an attempt. In 5-10 years time superconducting d.c. motors and generators will be in regular production and I suggest an annual demand of 1700 MW of plant requiring 150 tons of superconductor (excluding copper)

In 10 to 15 years superconducting a.c. generators will be in regular production and the total demand for superconductors could easily reach 1000 tons/annum.

230

ELABORATION ET APPLICATIONS DES SUPRACONDUCTEURS EN NIOBIUM-TITANE

ET EN NIOBIUM-fiTAIN

E. ADAM et J. DOSDAT Thomson C.S.F., Chatou (France)

Elaboration and application of NbTi and Nb3Sn superconductors

SUMMARY: Superconducting materials find widespread applications in the field of intense magnetic field production, where they provide performances that are quite unattainable by conventional means.

One of the main properties required by users comprises critical magnetic fields that are as strong as possible, beyond which the material loses its superconducting properties.

This requirement causes pure materials to be set aside, preference being given to certain alloys and inter metallic compounds, even if the elaboration of the wire or tape entering into the composition of the coil, requires the development of special production techniques.

This paper describes these techniques and the properties obtained regarding niobium-titanium alloys, on the one hand, and Nb 3 Sn composite tapes, on the other hand.

The first ones obtained by conventional metallurgy appear to be well adapted to various applications for fields between 2 and 8 Tesla.

The second ones are much more delicate to apply, but their use makes it possible to reach very high magnetic fields, up to 16 Tesla.

1. INTRODUCTION

La production de champs magnetiques intenses necessite Futilisation de fils ou rubans supraconducteurs bobinables en grandes longueurs apres tous traitements mecaniques et thermiques. Les metaux purs supraconducteurs tels que Hg, Pb, Sn, In, Nb, Ta, V ont des champs magnetiques critiques de Fordre de 0,05 a 0.2 Tesla. En effet lorsque cette valeur critique est depassee on observe un retour a Fetat normal du materiau qui perd sa propriete de transporter des densites de courant electrique elevees sans dissipation d'energie, au-dessous d'une certaine temperature critique comprise entre 0 et 20 degres Kelvin.

Par contre certains alliages (Nb Zr, Nb Ti) et composes intermetalliques (Nb3Sn, V3Ga, Nb3Al, V3Si ...) conservent leurs proprietes supraconductrices pour des champs magnetiques allant de 5 a 25 Tesla. Les alliages Nb Zr et Nb Ti peuvent etre fabriques par les methodes classiques d'elaboration de fils (fusion, martelage, laminage et trefilage). Au contraire le Nb3Sn et le V 3 Ga etant des materiaux tres durs de ductibilite pratiquement nulle, il a ete necessaire d'imaginer les procedes nouveaux permettant d'aboutir a des conducteurs souples ayant une bonne tenue mecanique en bobinage.

Les materiaux supraconducteurs de seconde espece, en particulier le niobium-titane et le niobium-etain, ont trouve un large domaine d'application, leur champ critique Hc2 eleve a permis de realiser des bobinages produisant des champs intenses pouvant aller jusqu'a 6 Tesla dans des volumes de Fordre du metre cube. Cette valeur du champ magnetique peut etre portee a 15 Tesla dans un volume plus restreint.

2. ELABORATION DE CONDUCTEURS SIMPLES ET COMPOSITES EN Nb Ti :

2.1. CHOIX DE L'ALLIAGE

L'alliage niobium-titane est un produit particulierement interessant pour la rea­lisation de produits supraconducteurs tels que des fils, des bandes, car :

233

1. C'est un materiau ductile susceptible d'etre obtenu par les methodes classiques de la metallurgie par opposition au compose fragile Nb3Sn qui necessite une technologie de fabrication particuliere et delicate.

2. C'est un materiau possedant un champ critique Hc2 interessant se situant vers 12 Tesla (a rapprocher des 7 Tesla de l'alliage ductile Nb Zr).

Pour le fabricant de produits supraconducteurs, fils, meplats, conducteurs creux, differents objectifs sont a atteindre.

1° Des considerations economiques conduisent naturellement a chercher a obtenir une densite de courant critique la plus elevee possible.

2° Dans le cas ou le produit est utilise dans la realisation de soleno'ides de grandes dimensions, l'energie magnetique stockee devient tres importante (750 Megajoules dans le cas du bobinage de la grande chambre a bulles du CERN). II est alors neces-saire de stabiliser le materiau. Cette stabilisation consiste en un gainage par un produit a faible resistivite (cuivre tres pur).

3° Dans ces memes bobinages de grandes dimensions les efforts mis en jeu deviennent considerables et un soin tout particulier doit etre pris pour conferer au produit une bonne tenue mecanique.

2.2. PROCEDE UTILISE

Les barres d'alliage supraconducteur sont introduites a l'interieurd'un tube de cuivre a faible resistivite et l'ensemble subit un coetirage qui provoque la liaison metallurgique conduisant a des resistances electriques et thermiques nulles.

Les problemes classiques du trefilage (angle forme et nature des filieres, lubrifi-cation) ont ete ainsi resolus en pratiquant la technique du cotrefilage (procede brevete Thomson). Cette technique, associee avec des traitements thermiques appropries en cours de transformation, conduit a une liaison «metallurgique » entre le cuivre et le niobium-titane.

Par cotrefilage il est ainsi possible de realiser des fils monobrins, des fils multi-brins utilises principalement pour la realisation de petits et moyens bobinages (jusqu'a 50 cm de diametre).

Les fils monobrins ou composites precedents peuvent etre egalement utilises pour la realisation de meplats ou de conducteurs creux a fort courant nominal (1000 a 10 000 A). Les fils sont alors plaques a chaud entre meplats de cuivre : procede dit de colaminage (brevet Thomson).

2 . 3 . EXEMPLES DE REALISATION

1. Conducteur composite rond — Ce conducteur presente un courant critique de 430 A a 5 T. Son courant stable est superpose au courant critique. Ce conducteur peut egalement etre realise en diametre plus petit : — diametre exterieur minimum 0,12 mm — diametre correspondant d'un brin de Nb Ti 0,015 mm a) Structure du conducteur composite — dimensions exterieures de diametre 1,5 mm — alliage supraconducteur Nb Ti

- nombre de brins 10 - diametre d'un brin 0,18 mm - section totale du Nb Ti 0,25 mm2

— cuivre OFHC - section de cuivre 1,50 mm2

— rapport section cuivre/section Nb Ti 5,8 a 6

234

b) Caracteristiques electriques La courbe de la figure 1 represente, en fonction du champ magnetique applique (perpendiculaire a l'axe du fil), le courant critique et le courant stable.

800 Jj

600

400

COURANT EN A

\

: CARACTERISTIQUES DE COURANT CRITIQUE ET DE COURANT STABLE DU F I L COMPOSITE.

\

200

DIAMETRE NOMBRE DE FILS $ D'UN F I L Nb-T i

COURANT CRITIQUE ET COURANT STABLE

\ 9

\

1, 5 m m 10

0# 18 m m

CHAMP MAGNETIQUE P E R P E N D I C U L A I R E EN kG

1 * ■ Z0 40

T -60 80

Fig.l

Un exemple d'utilisation interessant est le suivant : — induction nominate 5 T — courant nominal 420 A

I maximal — densite apparente

section totale du conducteur = 24 000 A/cm^

2. Ruban composite — Intensite moyenne

a) Structure et proprietes geometriques du conducteur composite — dimensions exterieures 10 mm x 1,8 mm — alliage supraconducteur Nb Ti

- nombre de brins 30 - diametre d'un brin 0,29 mm environ - section totale de Nb Ti 2 mm2 environ

— stabilisant cuivre O.F.H.C.

R , ^300°K a O T = 150 ^ 4 » 2 °K

- section cuivre/section Nb Ti 8 environ

235

b) Proprietes mecaniques a la temperature ambiante

— charge de rupture 400 kg — allongement inferieur a 0,01% apres efforts alternes du 200 kg (10 cycles)

c) Caracteristiques electriques

La courbe de la figure 2 represente, en fonction du champ magnetique (perpen-diculaire au plan du ruban) le courant critique et le courant d'impulsion (courant de recuperation IR).

3 000

1 000

INTENSITE EN A RURAN 1 750 A 54 kG 11 x 1,8 mm

H k G 30 40 50 54 60

IcA 2910 2 400 1 950 1 780 1 650

INTENSITE CRITIQUE I

■ INTENSITE REVERSIBLE I r

n I

. IR *C

y

I

V

I

CHAMP MAGNETIQUE A P P L I Q U E « I

20 40

, PLAN RUBAN EN kG

60

Fig. 2

Ce ruban est con^u pour fonctionner a 1500 A pour un champ de 5,1 T.

3. Ruban composite — Haute intensite

a) Structure et proprietes geometriques — dimensions exterieures 88 x 3 mm — longueur unitaire ^ 1 000 m — alliage Nb Ti — nombre de brins 200 a 400 — diametre du brin 0,18 mm — stabilisant cuivre O. F. H. C.

236

b) Resistivite du stabilisant

Champ magnetique applique en Tesla

5,5 4,5 3,5 2,5

Contrainte kg/mm2

10,5 10,5 10,5 10,5

Resistivite a l l ° K x l O ~ 8 Q c m

3,75 3,25 2,60 2,1

c) Proprietes mecaniques a basse temperature (4,2 °K) L'echantillon a ete contraint jusqu'a 15 kg/mm2 sans apporter de modifications a

la resistivite du cuivre et au courant de recuperation.

d) Mesures electriques Les courbes de courant critique et de courant de recuperation sont representees

figure 3 en fonction du champ applique.

10000

I A M P E R E

1 CRITIQUE

\

I E C U P E R A T I O N *%*X^

\

000 \

n mm.

t

'o mmmmmrmmmm

l ■ 2

mmmmmTmmt

3

: CARACTERIST1QUES 1 = f(B) DU RUDAN COMPOSITE F O R T E INTENS1TE

^_ « • 9 ^ ' 4 5 6 TESLA

Fig. 3

La stabilisation de ce ruban est egalement realisee en utilisant de l'aluminium raffine (99,99 de purete).

237

3. ELABORATION DE CONDUCTEURS COMPOSITES EN N b , Sn

3.1. PROPRIETES PHYSIQUES DU Nb 3 Sn

Le Nb 3 Sn est un compose intermetallique qui cristallise dans le reseau cubique P du tungstene. Son parametre cristallin est de 5,29 A. G'est un corps tres dur, stable a la temperature ambiante. Sa densite est legerement plus grande que celle de ses deux composants. Son coefficient de dilatation est de 7.10~6(°C)_ 1, c'est-a-dire assez voisin de celui du niobium. Son taux d'allongement maximal a la rupture est situe entre 0,2 et 0,4%.

Fig. 4

Ces proprietes mecaniques interviennent dans le choix de sa methode d'elaboration. La temperature minimale de formation de ce compos6 est de +850°C environ. La fusion est impossible a cause de l'eloignement des points de fusion : +232°C pour l'etain et + 2468 °C pour le niobium, le point d'ebullition de l'etain etant a + 2270 °C. On est done conduit a employer : soit une methode de frittage in situ d'un melange de poudres de niobium et d'etain, soit une reaction chimique a partir de composes halogenes sur un substrat metallique adequat, soit enfin un procede de diffusion a l'etat solide de l'etain dans un fil ou un ruban de niobium.

238

3.2. DESCRIPTION DU PROCESSUS DE DIFFUSION DE L'ETAIN DANS LE NIOBIUM

L'experience montre que pour obtenir une couche de diffusion epaisse et douee de bonnes proprietes supraconductrices, la temperature optimale se situe entre + 870 °C et +950°C. Or dans cette region on constate l'existence de plusieurs phases selon les proportions respectives des constituants. Le mecanisme de la diffusion peut se resumer ainsi : formation de Nb Sn2 au niveau du bain d'etain, puis de Nb6 Sn5, et enfin de Nb 3 Sn en contact avec Tame de niobium pur. La couche de Nb 3 Sn, une fois formee

Fig. 5

joue le role d'une barriere empechant une diffusion plus profonde de l'etain. Le compose Nb 3 Sn etant fragile et cassant il est necessaire d'utiliser un support meca-nique tel que le niobium pur qui est comparable a un bon acier. Le niobium est pris sous forme de ruban mince pour differentes raisons : — La diffusion se faisant en surface, on augmente l'interface liquide-support, done

la vitesse de formation du compose; — Au moment de la courbure du conducteur, les efforts de traction et de compres­

sion, qui dependent de la distance a la fibre neutre sont reduits au minimum;

239

— Sur chaque face du ruban mince (10 a 20 \i d'epaisseur) on forme une couche de Nb 3 Sn de 2 a 3 |i dont tous les points sont pres de la surface d'echange avec rhelium liquide ou le cuivre de stabilisation dont nous verrons la necessite.

Pour la formation de grandes longueurs de conducteur on utilise un procede de defilement en continu d'une bande de niobium dans un bain d'etain, sous vide pousse. Les traitements thermiques sont effectues avant et apres le passage dans retain. L'epaisseur totale est d'environ 14 u. La largeur du ruban peut varier de 3 a 60 mm.

3.3. PROPRIETES MECANIQUES ET SUPRACONDUCTRICES

— Sous la forme de ruban, le conducteur presente des proprietes d'anisotropie mecanique. II est relativement fragile au dechirement, mais la charge de rupture a la traction en ligne est superieure a 25 kg/mm2. Le rayon de courbure minimum est de 5 mm et le conducteur peut etre bobine et debobine plusieurs fois sans dommage.

— La temperature critique est de 17,8°K en champ magnetique nul et le champ magnetique critique a 4,2 °K est superieur a 22 Tesla.

Pour les applications, l'essentiel est la caracteristique de courant critique en fonc-tion du champ magnetique qui definit les valeurs maximales du courant utile, a un champ magnetique donne. Par exemple a 10 Tesla et 4,2 °K un echantillon court cuivre peut passer 90 A ce qui correspond a une densite de courant de 6.104 A/cm2.

Pour obtenir les performances voisines en bobinage il est necessaire de stabiliser le ruban supraconducteur, c'est-a-dire de le recouvrir par electrolyse ou placage, d'une couche d'un metal normal tres bon conducteur electrique et thermique. Son role est de servir de couche de secours en cas d'instabilites thermiques et magnetiques specifiques des supraconducteurs de 2e espece.

4. APPLICATIONS

L'alliage Nb Ti et le compose Nb 3 Sn permettent de realiser une large gamme de conducteurs adaptes a l'obtention de champs magnetiques intenses entre 2 et 15 Tesla dans les volumes magnetises allant de quelques cm3 a i m 3 .

La figure 4 montre un exemple de bobine homogene pour RMN (6 Tesla, dia-metre utile 60 mm, homogeneite dans 1 cm3 : 10" 6 ) .

La figure 5 montre un ensemble de bobines gigognes en Nb 3 Sn permettant d'obtenir respectivement 7,5; 11; 12,5 Tesla dans des diametres utiles 83; 35; 10 mm.

REMARK

E. J. SAUR (West Germany) — The relatively low overall critical current density of Nb 3 Sn prepared by the diffusion process can be improved by using the following method. Many tin clad niobium ribbons are stapled together by pressing and then rolled down to the original thickness. After a proper heat treatment (e.g. at 1,000°C for 10 hrs) many interior diffusion layers will be produced which raise the overall current density of this " multilayer ribbon " by a factor of the number of the layers.

240

STABLE SUPERCONDUCTORS AND THEIR USES

P.H. MORTON Imperial Metal Industries (Kynoch) Ltd., Birmingham (United Kingdom)

Supraconducteurs stables et leurs applications

RESUME : On examine les diverses methodes employees pour la stabilisation des supracon' ducteurs du type II afin de construire des bobines supraconductrices utiles. Les progres recents de la theorie de la stabilisation ont conduit a la mise au point d'un groupe entierement nouveau de supraconducteurs a tres fins filaments de Nb-Ti dans une matrice metallique. On donne un apercu de la theorie qui conduit a ces conducteurs intrinsequement stables et Von considere les qualites de diverses matieres pour la matrice. Vemploi de ces conducteurs per met des amelio­rations importantes du rendement des bobines et une simplification de la construction des bobines et des cryostats. On cite des exemples pratiques et on discute les developpements possibles de Vavenir.

The stability of a superconducting coil is a measure of the extent to which it is capable of operating at up to the critical current of the superconductor as measured in a short sample. In an incompletely stabilised coil the generation of heat as a result of flux-jumps causes the conductor to go normal at a prematurely low current. Stability can therefore be achieved in two ways; either by minimising the effect of local normal regions, or by preventing them from occuring.

In the first method it is necessary to combine the superconductor with a good normal conductor in a composite and to provide controlled cooling. This method has been used in all large magnets built up to the present time and is discussed in the first part of this paper. Study of the theory of flux-jumps has however now led to the development of the second type of conductor, the intrinsically stable superconductor in which flux-jumps can be virtually eliminated. Consideration of their properties and advantages forms the second part of the paper.

STABILISATION BY MEANS OF A NORMAL CONDUCTOR

Discontinuous changes in the magnetic flux through a superconductor, i.e. flux-jumps, occur principally during the energising or de-energising of a coil. Heat is generated locally and a portion of superconductor may go normal, in which condition it is highly resistive. In a composite conductor the good normal conductor prevents this normal region from spreading by providing (a) a temporary low-resistance electrical shunt and (b) a thermal path for removal of heat until the superconductor is cooled back to its superconducting state.

The materials most commonly used are niobium-titanium alloy as the supercon­ductor and copper as the stabiliser. To obtain stable performance, sufficient copper must be added and a sufficient portion of its surface must be in contact with a coolant, usually liquid helium. The numerical proportions are given by the familiar Stekly equation [1], which may be written in the form:

A

where w is the rate of heat transfer from the surface in watts/cm2, P (in cm) is that portion of the perimeter of the conductor which is cooled, I is the operating current

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(amps), p the resistivity of the copper (ohm-cm) and A the cross-section of the copper (cm2).

When cooling is by means of liquid helium the rate of heat transfer w, is frequently taken somewhat arbitrarily to be 0.4 watts/cm2. The process of heat transfer from copper to liquid helium has been studied by Lyon [2] and Wilson [3], As heat flux increases, heat transfer is by nucleate boiling up to a value of about 0.7 watts/cm2

at which level film boiling commences and heat-transfer rate drops sharply. On reducing the heat flux, film-boiling continues down to a level of about 0.3 watts/cm2

below which steady state conditions are again obtained. With careful control of surface conditions a coil could in theory be operated on the basis of w = 0.7 watts/cm2, but in order to be absolutely safe even in the event of some fault causing local film boiling it would be necessary to use a figure of about w = 0.3 watts/cm2.

The portion of the perimeter of the composite which is exposed to coolant, P, is dependent on the individual coil design and is affected by the size and shape of magnetic field to be produced, the mechanical stresses produced and the method of cooling.

The method of cooling used in large coils varies considerably as can be illustrated by considering some recent examples.

SACLAY ONE METRE COIL

The one metre coil recently made by CEA, Saclay, is oriented with its axis vertical and is layer wound. Vertical cooling channels 0.8 mm wide are provided between the successive layers by means of longitudinal spacers. Since cooling is thereby obtained on most of the wide face of the conductor the value of the cooled perimeter, P, is relatively high and the area of copper, A, correspondingly reduced, resulting in a conductor relatively rich in superconductor as shown in the cross-section in figure \a.

The coil is composed of four sections, each having an inside diameter 1.0 m, outside diameter 1.3 m and height 19 cm. The sections are mounted in pairs with 2 cm separation and with a central gap of 20 cm making a total height of one metre. The conductors have cross-section 10 mm x 1.8 mm and a length of 3,150 m per section. Energised by a current of 1,360 amps this coil has produced a central field of 36.5 kGauss corresponding to a total stored energy of 8.5 megajoules.

In order to withstand the magnetic forces during operation of the coil, one solution adopted was to cold work the composite about 3%, thereby more than doubling its proof stress. This had practically no effect on the critical current of the superconductor and the slight rise in resistivity of the copper could be tolerated. This method of strengthening has been discussed by Taylor et al [4],

Explosive welded joints were used and proven for the first time in two of the sections of this coil. When using a cold-worked composite, the joining of individual pieces of conductor to form a longer length presents some problems. Most types of soldered or welded joint either result in annealing and hence weakening of the copper or else have low mechanical strength and/or high electrical resistance. By explosive welding both these problems were overcome and joints were made which were stronger than the parent conductor, even when dressed back to the same dimensions, and which had an electrical resistance of less than 2 x 10" 9 ohms.

I R D HOMOPOLAR MOTOR

The winding of the field coil for the 3,000 HP homopolar motor being built by International Research and Development Limited, Newcastle-on-tyne, is described

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fully elsewhere in these proceedings. The conductor dimensions, 10 mm x 1.8 mm, are the same as in the Saclay One Metre Coil but the method of winding is very different. The axis of IRD's field coil is horizontal and vertical cooling channels are

Fig. 1 — Cross section of typical conductors. (a) Conductor for CEA Saclay 1 metre coil, 10 x 1.8 mm. (b) Conductor for IRD motor, 10 x 1.8 mm. (c) Conductor for CERN experimental coil, 6.5 x 5.0 mm. {d) Intrinsically stable conductor, 0.50 mm diameter.

produced by using a pancake construction. Cooling of the conductor is therefore on the narrow faces only. Consequently the value of P is less and the area of copper, A, is greater than in the Saclay coil. This is shown in the cross-section shown in figure 16, which is relatively rich in copper. Here again it was necessary to strengthen the con­ductor by cold working it about 3%.

CERL 100 mm COIL

A third method of cooling rectangular composite conductors is illustrated by the Central Electricity Research Laboratory coil shown in figure 2a, in which grooved tape, with transverse grooves across one flat face, was used. When wound in pancake form, this provided cooling channels through the pancakes as well as on the faces of the pancakes, thereby increasing the cooled area. The coil can also be used in either orientation, with axis either vertical or horizontal, without deterioration in cooling due to gas bubble formation.

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This coil is also of interest as it was the first coil to be built with the now familiar integrally-processed composites which were developed jointly by CEGB and IMI. It has been operating for over 18 months with remarkable freedom from trouble of any sort.

OXFORD INSTRUMENTS QUADRUPOLE

Figure 2b shows one pole of a quadrupole being wound by Oxford Instrument Company. The conductor used contains 16 superconducting filaments and has a cross-section of 4 mm x 1.5 mm. The choice of a low aspect ratio rectangular con­ductor permits the bending in two planes which is necessary when winding complex shapes.

Fig. 2a 100 mm bore coil made by Central Electricity Research Laboratory, Leatherhead, Surrey.

T H E USE OF HOLLOW CONDUCTORS

An alternative method of cooling which is now being explored is the use of super­critical rather than liquid helium. A hollow conductor section is used and supercritical

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helium pumped under pressure through the bore. The absence of a second phase in the coolant enables a higher heat transfer coefficient to be reached. Once the engineer­ing problems of handling supercritical helium have been solved, the coil construction becomes much simpler since no additional spacers and cooling channels are needed in the coil. The tubular conductors may be 'potted' in resin after winding, producing a very strong and compact design. Since the outside of the coil may be in vacuum, the cryostat construction is also simplified.

Fig. 2b Quadrupole magnet made by Oxford Instrument Company.

The largest coil of this kind so far built has recently been made and tested by Dr. Morpurgo at CERN, Geneva. The assembly prior to insertion in the cryostat is shown in figure 2c. Supercritical helium is pumped through a heat exchange coil (at left in the picture) to bring it to nearly 4.2 °K and thence into the superconducting coil (at the right of picture). Overall dimensions of the coil winding are; inside diameter 28 cm, outside diameter 60 cm, length 45 cm.

Construction of the coil consists of 32 double pancakes, each containing 76 metres of hollow conductor with outside dimensions 6.5 mm x 5 mm as shown in figure lc. Winding consisted in wrapping the tube with Mylar insulation, winding the double pancake with an interleaving layer of Araldite, binding it with glass tape and potting it in Araldite. The pancakes were stacked between stainless steel plates held in positions by tie rods and the necessary connections made. The coil is inserted in a vacuum container from which it is separated by layers of superinsulation.

Electrically all the pancakes are in series, but for cooling they are divided into eight parallel circuits each containing four double pancakes, i.e. with a coolant path of a little over 300 metres. The input pressure of the helium is 8 to 10 atmospheres and the exit pressure 5 to 7 atmospheres. The cooled weight of the coil is 600 kg. Quenching takes place at 1,340 amps with a central field of 48 kG and has been found

245

to be both well controlled and reproducible. The coil has also been operated con­tinuously for one hour at only 20 amps below the quenching current with complete stability.

Fig. 2c Test magnet made by CERN, Geneva.

PARTIALLY STABILISED COILS

Many small and medium-sized coils are in fact designed to be incompletely stabilised, and indeed with small coils this is almost universal practice. The amount of copper added is deliberately limited and the coil is designed to operate at significantly below short sample current. In a well-designed coil the increase in overall current density as a result of using less copper outweighs the affect of the decrease in operating current due to incomplete stability. The amount of copper used depends on the density of winding, size and shape of coil etc., but the most commonly used conductor con­tains a core of superconductor 0.25 mm diameter in copper of diameter 0.40 mm.

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INTRINSICALLY STABLE CONDUCTORS

The cause of instability in a superconducting wire is believed to be the occurrence of 'flux-jumps'. When an external field is applied to a superconductor, current loops are induced in it at the critical current density Jc. These magnetisation currents may under some circumstances decay suddenly, releasing energy and thereby heating the conductor. This process is referred to as a flux jump and, if sufficiently severe, can lead to a quench. However Hancox [5] has shown that if the diameter of a super­conducting wire is small enough the magnetisation currents are stable and do not change discontinuously, that is to say flux jumps do not occur. Such a conductor can be described as intrinsically stable.

For niobium-titanium the critical diameter for intrinsic stability is calculated to be about 0.05 mm (0.002 in). Wire of such a size is however both expensive and difficult to handle. A more satisfactory solution is to incorporate a number of fine filaments in a composite conductor.

The theory of fine filamentary composites has been summarized by Smith et al. [6]. Although the magnetisation currents within individual fine superconducting filaments are stable, the same is not true of the magnetisation currents formed when the filaments are embedded in a composite. Under the influence of a finite rate of change of field, H, magnetisation currents are driven from one filament to another through the matrix metal, forming complete loops occupying the whole width of the composite. These loops are not stable and the energy released in a flux jump is proportional to the square of the width of the whole composite. Smith showed however that if a high resistance matrix is used the transfer of current through the matrix requires a finite length dependent on H- In the case of niobium-titanium in a cupro-nickel matrix and a field changing at the rate of, say, 10 kG/sec the value of this cross-over length is about 28 cm. By twisting the filaments with a twist pitch of < 28 cm all magnetisation currents should consequently be suppressed. The data [7] shown in figure 3 confirms that by twisting such a composite its stability is indeed improved in accordance with the theory.

A complementary but independent theory of composites discussed by Chester [8] shows that, provided H is small, stability can be conferred by using a matrix of low electrical resistance to slow down flux motion. Again the filaments have to be finer than a critical size and, although the critical value has not been precisely calculated it is again approximately 0.05 mm.

There are thus two ways of making intrinsically stable composites: a) Fine filaments in a copper matrix. These conductors are intrinsically stable provided the rate of change of field does not exceed 1 or 2 kG per second. Coils can therefore be built which operate at short-sample critical current provided they are not energised in less than a minute or so. If Ic should be exceeded, the presence of copper provides a measure of protection. It is not strictly necessary for the filaments in the conductor to be twisted but by doing so the magnetisation currents which may extend over the whole length of the wire are much reduced; this reduces the remanant field and may be an advantage in high homogeneity coils. Consequently it is usual to introduce a small amount of twist but the pitch is not critical. b) Fine filaments in cupro-nickel or other high resistivity matrix. In these conductors it is essential for the pitch of the twist to be related to the rate of change of field. If Ic is exceeded, protection is not built-in and has to be provided externally. Provided these limitations are respected it is then possible to make coils which remain stable even if the magnetic field is changed at a rate of tens of kilogauss per second.

The principal advantages of the copper matrix intrinsically stable conductors are as follows:

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60.

50

40

30

20

10

\ \ 18 FILAMENTS

- \ A \ ^ > A

V

FIXED CURRENT RATE OF CHANGE

0 36 0 71 1 4 0

Nb Ti IN CuNi MATRIX

-

-

**" * ^ S s * , > * ^ "1 "" ""•••v.

" • ^ ^ H

SWEPT FIELD 1 OF FIELD(kG/SEC)

I 1

H 3 X 1 0 3

H 2 X 1 0 3

0 10 20 30

MAGNETIC FIELD(kG) ►

CURRENT/ FIELD CURVES FOR CuNi - Nb Ti COMPOSITE WIRE I T U R N / 8 C M

<i

10 20 30

MAGNETIC FlELD(kG)

3X10°

2 X 1 0 5 £

CURRENT/ FIELD CURVES FOR Cu N i - N b Ti COMPOSITE WIRE TWISTED I TURN / C M

Fig. 3 — Stability of two twisted cupro-nickel composites. Critical currents measured with swept field. The sample with higher rate of twist, b, was stable under rapid rate of change of field whereas the other, a, was not. {Courtesy: Rutherford Laboratory).

1. The designer is given a material which will operate reliably and reproducibly at a known critical current, Ic, and can be raised to that current in a minute or so. Experience shows that it is, however, essential to prevent wire movement in a coil. Several coils have been made which failed to reach Ic until they were 'potted' in grease, wax or resin in order to eliminate completely any trace of wire movement.

2. Coil construction is simpler. There is no need for cooling channels in the coil and by potting in resin one can achieve an extremely strong robust construction. The coil can also be operated in vacuum, simplifying the cryostat.

3. Current density is increased—both because of the intrinsic properties of the con­ductor and because of the avoidance of cooling channels. It is therefore possible to wind quadrupoles with greater field gradients and simpler windings. Devices are also made possible whose design was previously limited by space considerations. Polarised target magnets of improved design are also possible because the increased current density reduces the mass of the winding and means that particle beams can pass through it with less absorption.

4. It is no longer necessary to work at 4.2 °K. The possibility of operating in gas or vacuum enables one to move to either lower or higher temperatures. By going to lower temperatures the critical current density is increased and it may therefore be necessary to use slightly smaller diameter filaments, but it should still be possible to operate at \c. Good and Hudson [9] made a coil which produced 69 kG with a current of 37 A at 4,2 °K. The same coil was operated at 49 A at 1.3 °K to produce a field of 89 kG. By going to higher temperatures, the current density is reduced but in some larger coils the resultant extra cost of conductor may be more than offset by reduced refrigeration cost.

5. No heat is generated in a coil during operation and the refrigeration needed is consequently reduced. The only losses are in the current leads and supports and by virtue of the small amount of thermal radiation. Such a coil would therefore be ideal for use in conjunction with a small self-contained closed-cycle refrigerator.

6. If the filaments are slightly twisted, there are practically no magnetisation currents. This is an advantage in high accuracy or high-homogeneity coils.

7. There is a measure of protection if Ic should inadvertently be exceeded.

Fine filamentary composites in a cupro-nickel matrix have the disadvantage that there is no built-in protection if Ic is exceeded, but this can be provided by other means. Twisted cupro-nickel composites are being seriously considered for low frequency alternating current devices such as synchrotron magnets and in devices such as large superconducting switches where the coil has to be switched on or off very rapidly. The use of cupro-nickel composites will develop more slowly than that of copper composites but their existence makes feasible a range of devices not previously possible.

Although certainly conductors with filaments of diameter about 0.25 mm will continue to be used in many applications, the many advantages of the new intrinsically stable conductors ensure their increasing use in all sorts of superconducting device including many which could not operate successfully without them.

REFERENCES

[1] Z.J.J. STEKLY et al., Annex 1966-5 Bull. I.I.R., pp. 491-503. Com. I, Boulder. [2] D.N. LYON, Adv. Cryog. Engg, 10 (1964). [3] M. N. WILSON, Proceedings of the Second International Conference on Magnet Technology

(1968). [4] M.T. TAYLOR, A. WOOLCOCK and A.C. BARBER, Cryogenics (Oct. 1968). [5] R. HANCOX, IEEE Trans, on Magnetics (Sept. 1968).

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[6] P. F. SMITH, M.N. WILSON, C.R. WALTERS and J. D. LEWIN, Proceedings of 1968 Summer Study on Superconducting Devices.

[7] C.R. WALTERS and M.N. WILSON, Rutherford Laboratory, to be published. [8] P.F. CHESTER, Rep. Prog. Phys., 30, part 2, 361 (1967). [9] J.A. GOOD and P.H. HUDSON, Cryogenics, 9 (1), (Feb. 1969) pp. 64-65.

REMARK

P.F. CHESTER (U.K.) —I t perhaps needs pointing out that the Hancox/ Swartz-Bean criterion for intrinsic (adiabatic) stability only applies to a single wire. If we have a single wire of diameter exceeding the adiabatic limit, then subdividing it into several parallel filaments of the same total area but spaced well apart in a matrix may actually increase the heat generated in a changing external field because of the greater flux intercepted. Quite possibly in those cases where the filaments are of the order 25 um in diameter and are in a copper matrix, the stability is of the " dynamic " variety proposed by the present speaker (Rep. Prog. Phys. 30, 361, 1967).

Transposing the filaments avoids the problem of internal flux linkages and can restore intrinsic stability.

In any case, if all the filaments are to be used, the conductivity of the matrix must not be so high in relation to the rate of change of field that the ordinary "skin effect" prevents full penetration of the field in the time available.

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EFFECT OF TEMPERATURE ON THE CRITICAL CURRENT DENSITY OF Nb - 44 wt % Ti ALLOY

R.G. HAMPSHIRE, J. SUTTON and M.T. TAYLOR Central Electricity Research Laboratories, Leatherhead, Surrey (United Kingdom)

Influence de la temperature sur la densite de courant critique de l'alliage Nb-Ti (44% Ti en poids)

RESUME : Valliage supraconducteur Nb-Ti (44% Ti en poids) est devenu le plus utilise dans la construction des bobines supraconductrices. Tandis que Von connait la variation de la densite de courant critique Jc en fonctiondu champ magnetique Bpour les alliages du commerce a 42 °K, on n'a pas de donnees sur leur comportement a des temperatures plus e levees. De telles donnees sont necessaires pour la construction de bobines supraconductrices qui doivent fonctionner tant dans Vhelium liquide a 4,2 °K que dans Vhelium gazeux a des temperatures plus elevees. V interet s'accrott pour la mise au point de bobines a grande densite de courant globale, dans lesquels on obtient une stabilisation t her mi que au moyen d'une fine sous-division du supraconducteur {sta­bilisation intrinseque). Le degre de sous-division requise depend de la variation de Jc avec la temperature.

La communication comprend des graphiques sur le rapport entre Jc, B (0-10 Tesla) et la temperature (4,2-9,5 °K). La temperature critique au champ zero Tc est de 9,2 °Kpour Niomax S (IMI) et de 9,0 °K pour T48 (Supercon). Pour ces deux materiaux Jc diminue lineairement en fonction de la temperature sur Vetendue utile du champ (2-8T).

The variation of critical current density, Jc, with magnetic field, B, of commercial superconductors is usually specified only at 4.2 °K, the normal boiling point of liquid helium. This data enables magnets to be designed from stabilized superconducting composites for operation in liquid helium. However interest is currently increasing in operating magnets in helium gas at temperatures in excess of 4.2 °K and to design such magnets it is obviously necessary to know JC(B) at the chosen operating temper­ature. The measurements we have made are on Nb-44 wt% Ti which is now the most commonly used superconducting alloy. The samples chosen were the commercial superconducting wires I.M.I. Niomax S and Supercon T48 and we have determined their complete Jc, B, T surfaces from 4.2 °K up to the transition temperatures at approximately 9°K. The data on the I.M.I, alloy has been used by Carter et al. [1] to predict the high temperature critical current of the C. E. R. L. 100 mm bore solenoid, and in their paper they compare this with the actual performance.

The cryostat for measuring Jc as a function of temperature and magnetic field is shown in figure 1. The copper-clad Nb-Ti test wire was soldered to the bottoms of the copper current leads, soldered through the hollow pin glass-metal seals and tensioned into a U shaped groove around the copper block. To obtain good thermal contact it was necessary to solder the test wire into its groove on the copper block. The solder, supplied by Johnson-Matthey, was an In-Sn-Bi alloy chosen for its low melting point of 78 °C. The outer copper can was similarly soldered to the copper flange and inserted in the 25.4 mm bore of a Nb3Sn 10 Tesla solenoid so that the bottom of the U of the test wire was in the maximum-field region. The copper block was heated by heater windings at both ends. The dominant heat leak to the liquid helium bath was by helium exchange gas maintained at a pressure of 19 mm Hg. There was also significant cooling down the copper cladding of the test wire. Typically 130 mW of heater power was required to raise the block temperature to 10°K and less than 30 seconds was required to reach thermal equilibrium. By alternative use of the top and bottom heater coils it was established that there was a negligible temperature difference between the carbon resistor and the bottom of the copper block.

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LIQUID HELIUM

FIELD DIRECTION

( O - I O tesla)

MAXIMUM

FIELD

Vapour cooled Cu current leads

Tube for wires and pumpinq He qas

Glass/meta l hollow pin seal

Copper f lanqe

Low meltinq point solder seal

Copper can

Helium exchanqe qas

Ventinq hole

S.S. support tube

Heated copper block

Heater coils

Ge thermometers

Carbon thermometer

Sample soldered to copper block

Fig. 1 — Apparatus to determine critical current as a function of field and temperature.

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Electrical connections inside the can were made with Eureka wires which were soldered to copper wires thermally anchored with GE 7031 varnish to radial slots in the central post on the copper flange. The copper wires were threaded through the exchange gas pumping tube and connected to glass-metal seals at the top of the cryostat.

An independent determination of the transition temperature of the I.M.I, wire gave Tc = 9.24 °K which is significantly less than the value of 9.4 °K measured with the present apparatus. This is attributed to a temperature difference, AT ~ 0.2 °K between the Cu block and the test wire caused by poor heat conduction across the soldered joint and the high conductivity of the copper cladding of the test wire. The data presented in this paper has been corrected for the temperature difference by assuming that AT is proportional to the difference between the measured temperature and the bath temperature (4.2 °K). The corrected temperatures are probably accurate to ± 0.1 °K.

A calibrated Texas Instruments germanium resistance thermometer was used as the temperature standard in zero field. Because of large magneto-resistance effects in the Ge thermometers we used a J Watt Allen-Bradley carbon resistor to measure temperatures in the magnetic field. The following procedure was used to determine the small magneto-resistance of the carbon thermometer. During the zero-field calibration of the carbon thermometers we determined the relationship between heater power and the block temperature at a constant exchange gas pressure of 19 mm Hg measured at the top of the exchange gas pumping tube. The test sample was removed for these

5 0 0 0

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2 0 0 0

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: \ \ \ \

\ \ V

\ \

\

\ \ V

x \ \ x- \

4 0 SO 6 0 7 0 8 0 TEMPERATURE. °K

9 0 lO-O

Fig. 2 — IMI Niomax S.

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calibrations to eliminate cooling by the copper cladding. Since the only significant cooling was now by the exchange gas we could assume that the relationship between power and temperature was independent of field. With the exchange gas pressure kept at 19±0.2 mm we measured the magnetoresistance of the carbon thermometer at a series of 10 heater powers (covering 4.5 to 10°K in 0.5 °K steps) at 1 Tesla interval for fields up to 10 Tesla. The maximum deviation from the zero field calibration was at 4.2 °K and 10 T and it was equivalent to 0.2 °K.

The critical current Ic was determined using potential probes soldered to the test wire 5 mm above the bottom of the U bend. Current-voltage characteristics were recorded automatically on an XY plotter by ramping the current until a voltage of 500 uV developed across the potential probes when the current was automatically cut off. Ic was defined as that current which produced a detectable voltage (< 5 uV) across the specimen. This is an ohmic potential caused by current, in excess of the true critical current flowing in the copper, so that the values of Jc are too large. However the error is less than 5 amps m m - 2 , which is not significant.

Fig. 3 —1MI Niomax S.

The variation of critical current density, Jc, with temperature at constant transverse field, B, for I.M.I. Niomax S is shown in figure 2. The wire had a core of supercon­ductor of area 6.2 x 10 ~ 2 mm2 (nominal radius 0.12 mm) and a copper sheath 0.17 mm thick. At fields between 2.4 T and 8.0 T, Jc decreases linearly with increasing temper­ature; for the lower fields the slope decreases at high temperature. In zero field, Jc apparently saturated at 5,500 A m m - 2 at temperatures below 5.5°K instead of following the straight line of the high temperature data (not shown in fig. 2). This low value of Jc was probably due to propagation of a normal region from the resistively heated joints between the test wire and the current leads.

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Figure 3, derived from the graphs in figure 2, shows Jc plotted vs B at constant temperature. The results for zero applied field are anomalous because of the self-field of the transport current (0.5 T at the superconductor surface when J = 5000 A mm"2).

At any given temperature (less than Tc) one can define a critical field BJc = 0, at which the critical current is zero. The variation of BJc = 0 with temperature is shown in figure 4. Since Jc varies linearly with temperature over the useful field range, the critical current can be determined at any field and temperature using figure 4 together with the standard 4.2 °K Jc vs B curve.

TEMPERATURE,°K

Fig. 4 — Relationship between temperature and Bjc = 0 for IMI Niomax S.

The critical current of the Supercon T48 wire also varies linearly with temperature, but Jc is lower for fields less than 80% of BJc = 0. The zero field JC(B) curve extra­polates to zero at 9.0°K, but detailed measurement at very low currents show that the critical current remains finite up to 9.5 °K. This was also observed in an inductive measurement. The effect is attributed to inhomogeneity of the alloy.

The design of a superconducting magnet to a given specification becomes imprac­tical and in some cases impossible, if Jc falls below a certain value. This imposes a limit on the usefulness of a given superconductor at high temperatures. We have arbitrarily chosen a minimum acceptable value Jc = 500 A mm - 2 and constructed a graph, figure 5, showing the variation of the highest attainable field as a function of operating temperature for I.M.I. Niomax S.

9

* 8 0

< en UJ

UJ

5

1 1 1 1 1 1 2 3 4 5 6 APPL IED F I E L D . TESLA

Fig. 5 — Relationship between temperature and maximum applied field for Jc = 500 amps. mm-2.

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It is important to design superconducting magnets so that their performance is predictable. Hard type II superconductors are dissipative in a changing field; the heat produced can make the superconductor become momentarily normally conducting. This phenomenon is usually referred to as flux jumping. The normal region so pro­duced is usually localised but in a magnet the transport current causes Joule heating which can make the normal region grow. There are three approaches to this problem: cryostatic, adiabatic and dynamic stabilisation. These methods have been comprehen­sively reviewed by Chester [2]. Cryostatic stabilisation depends upon removing the Joule heating in the conductor to its surroundings faster than it is generated; this is achieved by cladding the superconductor with copper. Adiabatic stabilisation is achieved by never allowing the superconductor to become normal. Dynamic stabiliza­tion depends upon limiting the rate of energy release in the superconductor to become commensurate with the thermal diffusivity of the conductor. Some copper is required for this purpose but much less than for cryostatic stabilization.

For cryostatic stabilization it is necessary to know the dependence of critical current on temperature. In the simplest theoretical approach [3] it is assumed that the critical current decreases linearly with temperature; figure 1 shows that this assumption is very good.

For both the adiabatic and dynamic approaches, stabilization depends on fine sub-division of the superconductor. Neither method is normally used in magnets operating at 4.2 °K because of the expense of producing sufficiently fine wire or sheets of the superconductor. The question is whether the required degree of sub-division is reduced at temperatures in excess of 4.2 °K, thereby making the techniques economi­cally viable. Both methods of stabilization involve the parameter oc(T) = Jc( — SJJST) which can be determined from the data in this paper.

The criterion for a superconducting wire to be adiabatically stable [4] is:

/ 5 n C V w < mm \ 8 o /

where w is the wire diameter. C is the specific heat which has been estimated from the data in Ref. [5] and [6] as:

C = 5 6 0 — 5 — T + 1 3 T 3 J m " 3 o K _ 1

BC2(T)

Using the experimental data to give values of a, the diameter w should be less than 0.05 mm at 4.2 °K and 2.4 T. At 7 °K the permissible thickness increases to 0.22 mm. This is due both to the increase in specific heat and the decrease in the parameter a. The very substantial increase in the allowed size of the wire makes the adiabatic approach much more attractive, but only for low field magnets.

The dynamic stability criterion for a copper/superconductor composite sheet is:

In this expression x is the thickness of the superconductor and K its thermal conduc­tivity (taken a s 0 . 2 W m - 1 °K~*); X is the thickness of the copper and p is resistivity (taken as 3 x 1 0 _ 1 0 Q m ) . Using these values for K and p and the experimental value for oc of 5.9.105 A 2 m m " 4 o K _ 1 at 4.2°K and 4.3 T, then, for X/x = 3, x should be less than 0.13 mm. If the temperature is increased to 5.6°K, a becomes

256

3.2.105 A m m - 4 °K~* and hence the maximum value for x becomes 0.17 mm. There is therefore little advantage in operating a dynamically stabilized magnet at elevated temperatures.

Thus the operation of superconducting magnets at temperatures in excess of 4.2 °K seems advantageous only to those designs using adiabatic stabilization and then only for low field magnets.

ACKNOWLEDGEMENTS

The work presented in this paper was carried out at the Central Electricity Research Laboratories and is published by permission of the Central Electricity Generating Board.

REFERENCES

[1] C.N. CARTER, K.G. LEWIS, B.J. MADDOCK and J.A. NOE. Commission I, London, Annex 1969-1 Bull I.I.R., pp. 331-338.

[2] P .F. CHESTER. Reports on Progress in Physics 30, 561 (1967). [3] Z.J.J. STEKLY and J.L. ZAR, I.E.E.E. trans. NS-12, 367 (1965). [4] P.S. SWARTZ and C.P. BEAN. J.A.P. 39, 4991 (1968). [5] B.Ya. SUKHAREVSKY, A.V. ALAPINA and Yu.A. DUSHECHKIN, Zh. Eksp. Teor. Fiz. 54,

1675 (1968). [6] J. FERREIRA DA SILVA, E.A. BURGEMEISTER and Z. DOKOUPIL, Physics Letters 25A, 354

(1967).

DISCUSSION

J.E.C. WILLIAMS (U.K.) — You use as the criterion for the critical current, the appearance of 5 uV along the sample. Have you considered the current flowing in the copper block owing to this voltage and adjusted the value of critical current accor­dingly ?

J. SUTTON — The measured resistance between the potential probes in zero field when the temperature was just above Tc was 15 uQ. This is just the calculated resistance of the copper sheath. Hence the resistance of the soldered contact must be sufficiently high to prevent a significant current from flowing in the copper block.

If the criterion for the critical current is 5 |iV across the probes, then the error in Jc due to currents in the copper sheath is only 5 amps mm - 2 , and is therefore negli­gible.

257

A MODEL SUPERCONDUCTING MOTOR

A.D. APPLETON and R.M. MacNAB Electrical Engineering Department

International Research and Development Co. Ltd. Newcastle upon Tyne (United Kingdom)

Maquette d'un moteur supraconducteur

RESUME : On decrit un moteur supraconducteur homopolaire qui a developpe 50 ch a 2000 tr/min. On a utilise un cable supraconducteur Nb-Zr a sept fils, partiellement stabilise par cuivre. Le champ magnetique maximal etait de 2,7 Wb/m2 et Valesage a temperature ambiante etait de 0,35 m. On decrit la construction et le rendement du moteur et le developpement associe des balais electriques.

INTRODUCTION

This paper describes the design and performance of a 50h.p. homopolar motor with an ambient temperature rotor operating on the Faraday disc principle and with the field supplied by a stationary superconducting coil. Before building this motor it was established that this configuration of machine could, in very large ratings, provide the basis of designs for industrial drives which would compete with con­ventional machines. The model motor was built to illustrate the feasibility of engineer­ing this type of machine and to establish the design principles; its size and power rating were much too low for it to be considered other than as an experimental machine. In consequence, the design embodied as far as possible techniques which could be extrapolated for use on large machines of the same type.

MOTOR OPERATION

A simple solenoid field winding will generate a flux configuration appropriate to this type of machine and such a solenoid wound with high field superconductor will permit the elimination of iron from the magnetic circuit of the machine. This is the key to the saving in weight which can be made with superconducting machines.

Figure 1 depicts a motor with a superconducting field winding in a cryostat which has room temperature access to the bore. The disc type rotor runs in the bore of the coil, with its axis of rotation coincident with the coil axis, and the motor voltage is generated between the inner and outer sliprings by the rotor disc conductor cutting the field coil flux. The model motor was built with two such discs connected electrically in series to double the armature voltage. Figure 2 is a drawing of the model motor armature.

The geometric proportions of the machine are determined by economics. A given investment in superconducting wire will produce more flux the greater the diameter on which the wire is wound, i.e. flat, large diameter, field coils are the most economic in superconductor. However, the cryostat and rotor rapidly become more expensive with increasing diameter and the actual size is chosen to minimise total cost. In high speed machines, diameter may also be limited by stresses in rotating parts.

THE SUPERCONDUCTING COIL

When the superconductor and coil former were designed, niobium zirconium was the most suitable superconductor, and Nb- 25% Zr in the form of wire of 0.010 in.

261

diameter was chosen for this machine. Stabilisation of superconductors was not then formalised and it was decided to reduce possible degradation of the coil performance by laying up the wires in the form of seven strand cable. When it later became apparent that copper would increase the stability of the coil and help to protect the super-

THIS DISTANCE MINIMISED

MAGNETIC FLUX;

INNER SLIPRING-

CRYOSTAT OUTER VESSEL RADIATION SHIELD

SUPERCONDUCTING WINDING

L - OUTER SLIPRING BRUSHES

STATIONARY DISC CONDUCTOR

ROTOR DISC CONDUCTOR

ARMATURE TERMINALS BRUSHES

AXIS OF

ROTATION

Nr

N SF

Fig. 1

conductor in the event of a quench, an outer wrapping of copper wires was wound on to the cable and the whole bonded together by indium. The cable was then insulated by braided glass fibre with small spaces to permit some helium access. The resulting conductor was only partially stabilised and the coil quenched at 240 A in a maximum field of 2.7 Wb/m2 which is somewhat below the short sample limit. Thereafter, in experiments on the motor the current was limited to 230 A to avoid quenching.

CRYOSTAT DESIGNS

The entire cryostat was of stainless steel with the annular coil vessel suspended from a 12 in. long neck tube of 2 in. diameter and 0.010 in. wall thickness inside the chimney. Figure 3 shows the coil box inside the cryostat which has the end plates

262

removed. A liquid nitrogen cooled radiation shield surrounded the coil box in the space between the latter and the cryostat outer vessel.

I 1 * Ja*

K rrmr ^7?—v sa F iS

-Lilt

Fig. 2

Cooling down of the cryostat was started about 20 h before running the motor, and both the coil and radiation shield were precooled with liquid nitrogen. This was then pumped out of the coil vessel and cooling was continued with liquid helium, 40 litres being required to cool and fill the coil and its vessel whose total weight was around 120 lb.

ARMATURE DESIGN

The diameter of the cryostat bore was fixed by cost considerations at 15 in. and, with the flux available in it (the expected flux was obtained in operation), it was possible to generate a total of about 10 V at 2000 r.p.m. This meant that to attain 50 h.p. something approaching 4000 A armature current was necessary, and this current would have to pass through four sets of sliding contacts in the armature circuit. This

263

problem of high current sliprings and brushes recurs to a similar extent on large superconducting homopolar machines, therefore, early in the design of the model motor a programme of brush tests was started to find the most suitable types for these machines generally and, more urgently, for the model motor itself. It was decided to investigate solid brushes, and the requirements were high current carrying capacity and low contact losses (the latter arise from sliding friction and contact voltage drop).

Fig. 3 — Model motor cryostat.

Several brush grades and slipring materials were tested under conditions similar to those envisaged for the motor (except that brush tests were in zero magnetic field), and brush losses, wear rates and cooling system requirements were assessed and the optimum operating pressures for the brushes were evaluated.

It was clear from available data on brush performance that to supply 4000 A to the 14 in. diameter rotor would entail having a large percentage of the slipring track covered by brushes. It is necessary for satisfactory operation of brushes that the slipring track forms a stable film of oxides and it was thought that a large number of brushes running on the track might inhibit this film formation. It turned out that the brushes could be run with no deterioration in performance and with a 50% brush cover of the slipring track. There were signs that significantly more than 50% cover by brushes led to increased slipring wear.

These brush tests were conducted using a slipring similar to the one on the model motor, but with brushes running on both the underside (as in the motor) and the outer surface of the overhung slipring. Figure 4 shows the brushes on the testing rig with the slipring removed.

The design of the brushgear for the motor involved some consideration of the presence of a high magnetic field. The slipring surface had to be profiled to follow the lines of flux since any flux crossing the brush/slipring interface would drive a circulating current round the brush with losses and deterioration in brush performance. It was also necessary to support the brushes against the side loads which they have to carry and this without altering the brush pressure setting which has to be maintained

264

within a fairly close tolerance to ensure satisfactory operation. To accomplish this a hinged brush was used and this was fairly successful.

Fig. 4 — Brush test rig with slipring removed.

As a result of the tests it was apparent that the sliprings of the model motor would require to be water cooled and that 1 % chromium-copper would be the most suitable slipring material because of its good thermal conductivity, although it was not the best material to minimise brush losses. Cupro-nickel had proved much better for losses, but its inferior thermal conductivity would not permit its use in the cooling configuration of the model.

All of the brush materials tested were basically mixtures of copper and graphite. Morganite grade CM IS proved to be most suitable for inner brushes and CM2 for the outer brushes (because of the difference in slipring speeds). However, because of supply difficulties CM1S was used throughout the machine. This departure of both brush and slipring materials from the ideal caused considerably increased friction at the outer brushes with a reduction in the motor efficiency.

The highest recorded efficiency of the model motor was 69% and with the ideal brush slipring combination this would have approached 80%. The greatest source of losses in the motor is the brushgear and the importance of minimising brush losses in large machines is very clear. The relatively low efficiency of the model motor is due largely to the small scale of this machine. Large motors can, despite their brush losses, achieve efficiencies well above 95%.

Tests were conducted to try to measure an armature reaction effect, but none was detectable as would be expected from the arrangement of the armature conductors. It was also quite apparent that there was no force on the superconducting coil due to the armature currents as the coil was satisfactorily supported from above only by the very thin walled tube.

Some of the experiments on the motor were conducted with iron in the magnetic circuit (placed in or near the cryostat bore and arranged symmetrically about the

265

coil). An increase of about 20% was obtained in the maximum flux using about 400 lb of iron, and maximum power output was, of course, obtained with the iron in use.

The open circuit characteristics of the machine (fig. 5) show the magnetic effect of the iron.

50 100 I 50 F i e l d c u r r e n t , A

200

Fig. 5 — Motor open circuit characteristic.

The performance figures of the motor under typical running conditions are shown in table 1 which also shows the figures obtained at maximum power output.

HIGHER VOLTAGES

Increasing the voltage of a homopolar motor can only be achieved by increasing the number of generating stages. The segmented rotor principle for making a multi­stage, higher voltage machine is presented at this conference in the survey paper « Motors, generators and flux pumps » by A.D. Appleton.

To test this principle the original rotor of the model motor was modified to a single disc 8-segment (4-stage) arrangement.

This modified rotor made possible the generation of 20 V. However, this voltage was not considered high enough to test the intersegment insulation, a critical point of this design.

266

Table 1

MOTOR PERFORMANCE

Typical running conditions Armature current Ifl 3500 A Field current 200 A Speed 1900 r.p.m. Voltage drop in busbars 0.75 V Voltage drop across rotor 0.99 V Voltage drop in armature conductors 0.20 V Total of brush voltage drops 0.79 V Generated voltage (E„) 8.36 V Terminal voltage 9.35 V Total brush electrical loss 2.76 kW Total armature electrical loss (incl. brushes) 3.50 kW Developed power (Eb . Ifl) 29.3 kW

Maximum load conditions Terminal voltage 10.74 V Generated voltage 9.45 V Armature current 3800 A Developed power 35.9 kW

The output power was measured on calibrated generators used as loads for the motor.

Arrangements were therefore made such that test voltages up to > 100 V could be applied across the insulation between adjacent segments while the motor was running.

These tests demonstrated the validity of the segmented rotor principle and achieved a maximum intersegment voltage of 40 V, thus establishing that the maximum stage voltage for this type of machine could not exceed this value—in fact a much lower interstage voltage is at present used in design.

ACKNOWLEDGEMENTS

The authors wish to thank the sponsors of this project, Ministry of Defence (Navy) for their great interest and help in all aspects of the work.

DISCUSSION

P.H. BURNIER (France) — Why are the rotor segments water cooled? There is a lot of space for reducing the current density, and accordingly the joule losses; there is no dB/dr, and subsequently no eddy current losses, because of the cylindrical segmenting of the magnetic field. Is water cooling necessary because of the dl/dt due to commutation from one segment to another?

A. D. APPLETON — To achieve an optimum design the sliprings must be water cooled. (The importance of a high "q" value is shown in my paper). Since water cooling is required it is convenient to use it to cool the rotor conductors to save copper and to use the rotor conductors as a means of supplying water to the slip rings; water cooling is not required for any effects resulting from the dJ/dt due to rotor conductor switching.

B. G. GAYDON (U. K.) — In the segmented rotor design, why was the maximum voltage between segments only about 40 volts before breakdown ?

A. D. APPLETON — Accumulation of debris from the brushes causes this low voltage breakdown.

267

ASPECTS OF A SUPERCONDUCTING WINDING FOR A 3250 hp MOTOR

A.D. APPLETON and J.S.H. ROSS Electrical Engineering Department

International Research and Development Co. Ltd., Newcastle-upon-Tyne {United Kingdom)

Aspects d'un enroulement supraconducteur pour un moteur de 3250 cv

RESUME : On decrit les aspects de la construction d'un enroulement supraconducteur Nb-Ti/ cuivre utilise pour fournir le champ d'un moteur homopolaire supraconducteur de 3 250 cv a 200 tr/mn. On presente les resultats d'essais sur la performance generate du supraconducteur.

INTRODUCTION

The superconducting coil described in this paper represents the first commercial application of the phenomenon first demonstrated by Onnes some 57 years ago; our coil will form the field winding of a 3250 hp homopolar motor which will be installed in the CEGB power station at Fawley later this year. The Fawley motor project was started in May 1967 and is being carried out by the International Research and Development Co. Ltd. under the sponsorship of the NRDC.

The coil design is essentially that of a large toroidal magnet and the configuration of the rotor windings is such that it eliminates all armature and torque reaction from the coil proper. With the sole exception of the rotor shaft, all components of the rotor are non-magnetic to minimise forces on the coil.

The coil takes the form of a rectangular toroid of 2.4 m bore and 2.8 m outside diameter with an axial length of 0.53 m; a total of 5 \ tons of copper-stabilised niobium titanium superconductor is used to produce a total flux of 7.1 Wb. The stabilised superconductor is of rectangular section, 10 x 1.8 mm, and is wound in the form of 18 double discs to obviate making joints at the inner bore of the coil during winding.

Explosively welded joints were made at the manufacturer's works because it was not possible to produce the superconductor in longer lengths than sufficient for single discs; these joints were located at the inner bore of the winding.

The discs are separated axially by rings of radially aligned spacers covering approximately 20% of the faces of each disc.

The coil is cooled by liquid helium in natural convection over the faces of each disc, the design of the spacers and the laths upon which the coil is wound being such that gas bubbles evolved within the coil are ejected to its outside faces thus avoiding parts of the coil being poorly cooled.

M E C H A N I C A L DESIGN

i : The stresses within the coil due to the electromagnetic forces on the conductors are mainly the tensile hoop stresses tending to burst the coil and the compressive axial stresses tending to compact the coil; as the hoop stresses alone were of the same order as the yield stress of annealed copper, it was obvious that a rigorous approach was required.

The field plot within the coil was obtained by a computer programme and, in conjunction with the appropriate current density, the body force on the conductors and the mechanical stresses within the coil were found directly using three-dimensional

269

elastic theory; the analysis developed by Middleton and Trowbridge of the Rutherford Laboratory proved particularly helpful. Up to this point no consideration has been taken of the fact that the axial loads are not transmitted through the coil uniformly, but are concentrated on the radial spacers. The axial stresses within the coil are governed predominantly by the area of each disc which is covered by spacers, but this in turn affects the stabilisation of the conductor on heat transfer grounds.

By choosing a certain degree of cold-work in the copper and knowing the mech­anical properties of the copper at 4.2 °K, it was possible to optimise the covered area of the discs so that the required heat flux from the conductor was minimised throughout the coil without exceeding the design stress. The condition chosen was a 3% cold worked copper with a yield stress at 4.2°K of 17,000 lb/in2 (fig. 1).

Required Heat Flux

wott*/» /cm

50*/ percentage of each disc r '* covered by spacers

0 7

0 6

0 5

0-4 \

0-3

0.2

O.I

2-/ i'L A'/.

Percentage Cold VKbrk in Copper

Fig. 1 — Required heat flux for stable operation v. percentage cold work in copper.

The critical stresses were those near the geometric centre of the coil cross-section where the hoop stress was 6,890 lb/in2, the axial stress 1,110 lb/in2 and the bending stress on the conductor spanning the spacer pitch 500 lb/in2 (fig. 2). The maximum, compounded, effective stress allows a margin of 25% on the specified minimum yield stress of 17,000 lb/in2.

Two particular points should be noted regarding the overall mechanical design:

(1) The allowable design stresses used took no advantage of the margin offered by the presence of the superconducting filaments; these effectively increased the tensile yield stress of the conductor by about 5,000 lb/in2;

(2) the stress analysis is on elastic theory; the use of plastic theory would almost certainly lead to a more efficient coil and may well be used in future designs.

270

It is also worth mentioning that the mechanical and electrical properties of the explosively welded joints at the inner bore of the winding were found to be fully acceptable; in a tensile test, the joint strength exceeded that of the parent material.

Compounded Total StrcM 13,660 p t i

Hoop ttrca.: 6890 pt

Mean axial ttntM. 11IO pv

Max. axial ttrcit — 1110 * Fa

« 6270 pti

(comprestive)

Fig. 2 — Component stresses at critical point in the coil.

STABILISATION ASPECTS

As already mentioned, the coil is fully immersed in liquid helium at a bath tempe­rature of 4.4 °K using a slightly pressurised cryostat in order to reduce refrigeration costs. The conductor is designed for solely edge cooling at a maximum heat flux of 0.325 W/cm2; experimental work showed that a nucleate boiling flux of 0.55 W/cm2

corresponded to a temperature difference of 0.45 °K. This data was obtained under conditions corresponding closely to design conditions—a bath temperature of 4.5 °K and with a Bicelex varnish covering 0.0005 in. thick, the peak nucleate boiling flux under these conditions being 0.61 W/cm2 (fig. 3). The proximity of the adjacent disc in the actual coil was of small importance as work by M. H. Wilson of the Ruther­ford Laboratory had shown that the cooling channels in the coil afforded virtually open-bath conditions.

The figure referred to above of 0.325 W/cm2 relates to only a small portion of the coil; the variation of required heat flux value over the complete coil is considerable, the minimum value being 0.204 W/cm2.

The conductivity of the explosively welded joints was found to be very similar to that of the parent material; it was thought advisable, however, to ensure that the joint was wound in the inner turn of each disc where the enhanced cooling could be advantageous.

271

The transition temperature of the niobium titanium filaments at 3.5 Wb/m is 7.9 °K; this compares with the expected conductor temperature of 4.65 °K when a heat flux of 0.325 W/cm2 is being dissipated from the conductor edges.

HEAT FLUX

Nominal bath temperature:4.4 °K

O l 0-2 0 3 0 4 OS 0 6 TEMPERATlfRE c> r DIFFERENCE K

(Copper to Liquid Helium)

Fig. 3 — Nucleate boiling heat flux from copper to liquid helium v. temperature difference.

TESTING OF SUPERCONDUCTOR SAMPLES

The mechanical properties of the samples cut from each production piece length were verified by establishing a correlation between the room temperature hardness of the sample and the proof stress at liquid helium temperature. By this means, all samples were checked very economically and it was not necessary to carry out tensile test at 4 °K on all the samples supplied. The mean value of proof stress of the samples was 26,000 lb/in2 compared with the specified figure of 17,000 lb/in2.

The resistance ratio of the copper in the finished conductor is difficult to determine, due to presence of the superconductor filaments, and a compromise had to be found. The resistance ratio of annealed copper samples from the lead end of each production piece length was measured and in no sample did the ratio fall below 180. Further test were unnecessary as our requirement in the annealed condition was 150.

A sample from every length used in the coil was checked for its short sample critical current at the rated field of 3.5 Wb/m2 parallel to the broad faces of the conductor. As the anisotropy of the conductor proved to be of the order of 10%, our alternative requirement of a perpendicular field of 2.7 Wb/m2 was nowhere critical.

272

Table 1 BASIC COIL PARAMETERS

Dimensions Inner bore 2.40 m Outer diameter 2.84 m Length 0.535 m Total weight 5 Vi tons No. of discs 36 No. of turns per disc 110

Electrical parameters Conductor filament current density 72,000 A/cm2

Conductor current density 4,000 A/cm2

Overall coil current density 2,550 A/cm2

Inductance 55 H Ampere-turns 2.86 x 106

Stored energy 1.46 x 107 Joules Coil current 725 A

Mechanical parameters Maximum hoop stress in conductor 6,940 lb/in2 tensile Maximum mean axial stress in conductor 1,310 lb/in2 compressive Maximum (axial) stress in spacers 6,840 lb/in2 compressive Minimum proof stress of copper in conductor at 4.2 °K . . 17,000 lb/in2

Degree of cold work in conductor 3%

Materials: Superconductor . . . . Composite conductor of NbTi in H.C. copper; 10mm x 1.8mm

with 5 filaments of 0.020 in. diameter Length: Piece length 1 825 m

No. of pieces 18 Rated Field . . . 3.5 Wb/m2 axial

2.7 Wb/m2 radial Current: 725 A

Laths Cold worked EN58B stainless steel with phenolic resin insul­ation strip

Coil former Stainless steel AI5I 304L Spacers Glass reinforced epoxy-resin mouldings. Varnish insulation . . . Bicelex .0005 in. film thickness. lnterturn insulation . . Glass fibre/isophthalate-polyester tape 0.005 in. thick

Because of the large number of tests, a special rig was devised in which the sample and sample holder could be withdrawn from the cryostat, the sample replaced with a fresh sample and further tests carried out; by this means the cost per test could be considerably reduced. A typical result is shown in figure 4.

The histogram (fig. 5) shows that the critical currents of the samples at the operating magnetic field level vary from 900-1500 A against the specified operating current level of 725 A. The latter is conservative with respect to heat transfer and strength capabilities of the composite superconductor; this value of current could be achieved with only four filaments of superconductor, but it was decided to employ five filaments because inter alia at a late stage in the test programme it is intended to increase the rating of the coil to its mechanical limits. With the present design, the latter will be reached before the critical current becomes limiting.

PROTECTION ASPECTS

The protection system for the coil consists of several individual circuits, all of which can initiate the EMERGENCY-TRIP sequence in which the coil supply

273

circuit breaker operates and the coil current passes through the dump resistor which is permanently connected across the coil.

800 1000 1200 Specimen current, A

moo 1600

Fig. 4 — Typical short sample characteristic.

600 800 1000 1200 Specimen critical current, A

Fig. 5 — Histogram of short sample test results.

Under normal operating conditions, the coil current will be controlled entirely by the small homopolar generator and can be charged at up to 10 V and discharged through the internal resistances of the generator. The protection circuits are:

(1) A bridge system of voltage tappings on the winding whereby a normal region will cause the bridge circuit to become unbalanced and the coil to be tripped. The actual levels at which the alarm and trip circuits will be determined during the test on the coil at IRD;

(2) The helium level in the cryostat will be monitored by two methods: a differential pressure transducer will monitor the hydrostatic head of helium in the vessel, and a bank of carbon resistors will indicate the level in the cryostat neck; if either circuit indicates a low helium level the coil may be tripped.

A pressure transducer on the coil vessel is set to trip the coil if the pressure rises much above 18 lb/in2, abs., or, alternatively, if the rate of change of pressure is excessive.

274

Under EMERGENCY-TRIP conditions, the coil current of 725 A is shunted through the dump resistor of 0.625 Q generating a step voltage across the coil of 450 V.

The air-cooled dump resistor is built from rust proof silicon steel elements mounted on porcelain insulators which are housed in a drip-proof enclosure with louvred walls.

TESTS ON THE WINDING

As each double disc was wound, it was tested to detect shorted turns by an RSO technique. Step pulses were injected in turn at either end of the completed double disc and the waveform compared differentially with that in a previously checked disc. The technique essentially detected an unbalanced condition between the two discs being compared; this could be resistive, capacitive or inductive. In several cases it was found that a fault indication was due to the proximity of temporary clamps on the winding.

This test could only be applied to the individual double discs, and just before the final joints between double discs were made a complete series of tests was carried out as a final check.

The joints on the outer diameter of the winding were designed so that the mechanical and electrical functions were separated as far as possible and X-ray and electrical tests made on each joint to ensure integrity. The resistance of similar joints in liquid helium was found to be 10" 7 Q.

DISCUSSION

R.G. SCURLOCK (U.K.) — A large amount of the refrigeration must surely be absorbed in cooling the current leads to the superconducting windings. Would you consider opemting the coil in the persistent mode, in association with protective devices ?

J. S. H. ROSS — No. Although a useful reduction in required refrigeration capac­ity would be gained by operation in the persistent current mode, we consider that the risks involved would be totally unacceptable in this prototype machine.

B.J. MADDOCK (U.K.) — A comment on stabilisation and heat transfer to boiling liquid helium. The essential function of steady state or cryostatic stabilisation is to ensure that a conductor will recover from unspecified disturbances some of which may cause temperature excursions of several degrees. This implies recovery from film-boiling. The whole of the heat transfer characteristic is therefore important and not just the nucleate-boiling part.

Recovery usually occurs at a flux of 0.1 to 0.2 W/cm2 for an uncoated surface all of which is in the film boiling state. However, if a portion is still in the nucleate boiling state, for example the parts of a conductor at the ends of a normal (resistive) region, then recovery occurs at a higher flux of about 0.3 W/cm2, again for an uncoated sur­face. This value does depend though on the critical temperature of the superconductor and on the ratio of operating to critical current. The value is somewhat higher for thinly coated surfaces.

275

THE CRYOSTAT AND REFRIGERATOR FOR A 3250 hp SUPERCONDUCTING MOTOR

F. TINLIN and J.S.H. ROSS

Electrical Engineering Department International Research and Development Co. Ltd., Newcastle-upon-Tyne {United Kingdom)

Cryostat et refrigerateur pour un moteur supraconducteur de 3250 cv

RESUME : On decrit la construction du cryostat et du refrigerateur pour un moteur supra­conducteur de 3 250 cv. Les processus dans le cryostat et le refrigerateur sont etudies afin d'obtenir un rendement global optimum, tout en tenant compte des exigences quant au re­ft oidissement du systeme.

INTRODUCTION

The overall concept of the 3 250 hp Prototype Superconducting Motor is described in a paper «Mo tors, generators and flux pumps » by A.D. Appleton (p. 207) and the present paper gives an account of the design aspects of the motor cryostat and refrigerator, and shows the necessity of treating both designs together from an early stage.

CRYOSTAT

The basic dimensions of the cryostat are defined by the design of the motor field coil, its method of cooling and manufacturing considerations.

The coil design chosen was of rectangular cross-section, 0.2 m x 0.53 m, with a mean diameter of 2.6 m. The winding was to take the form of discs or pancakes separated by radial spacers, to allow edge cooling of the windings by liquid helium (see paper "Aspects of a superconducting winding for a 3 250 hp motor" by A.D. Appleton and J.S.H. Ross, p. 269).

In order to provide the maximum usable bore area within the cryostat, and also to minimise the amount of liquid helium present, the vessel containing the coil had also to be a rectangular-section toroid, closely fitted to the coil.

The cryostat (fig. 1) is a double-walled vessel with a form of radiation shielding between the two walls and, to reduce heat leakage from ambient to the liquid helium, mechanical supports had to be as slender as possible.

The coil vessel supports could be made relatively light since they are required to carry only the dead-weight of the coil and vessel and small side loads due to stray field effects. In the design of machine chosen, no torque reaction force appears to the coil.

INNER VESSEL

In order to reduce to a minimum the distance from the coil bore to the cryostat bore (fig. 1), the former on which the coil is wound was designed to be the inner cy­linder of the vessel, and various methods of construction of the remainder of the vessel were examined.

The criteria to be met were:

(1) During cool-down the vessel is subjected to a high pressure for a period of a few days;

277

(2) When at low temperature, the differential contraction of coil and former would result in high compressive hoop stresses in the latter;

(3) The vessel must remain helium leak-tight after cycling to 4.4 °K; (4) In the manufacture of the vessel the coil temperature must not exceed 250 °C; (5) The material must have adequate strenght at room temperature and at 4.4 °K.

Following BS 1515, the maximum design stress used was 9.6 ton/in2. In addition the resistivity must be high so as to minimise circulating currents caused on coil discharge and the material should be non-magnetic.

Neck tube

Superinsulant

Radiation shield

Suspension cable

This dimension to be minimal

ffij >""- ™»*

Superinsulant

Outer casing

Fig. 1 — Section of cryostat.

The high pressure requirement results from the need to keep the cool-down time reasonably short, the limiting factor being the heat flux obtaining in the coil, which varies with coolant pressure. It was decided, therefore, to circulate pressurised helium gas through the refrigerator and cryostat to cool both together. The maximum pres­sure in the refrigerator cycle (a turbine expansion Claude cycle) was 9 atmospheres and, allowing for pressure drops, a convient operating pressure for the coil vessel was chosen to be 7 atmospheres. The design pressure was increased to 7.7 atm to allow for the functioning of safety devices.

The most suitable material was stainless steel, AISI 304L being chosen because of its strength and resistivity and availability. This steel is not stabilised, but the low carbon content ensures good weldability. Weld specimens were prepared and tested, the most critical test in this case being a Charpy impact test carried out at 77 °K.

278

A welded construction was considered necessary to meet the requirement for helium tightness over a long period of thermal cycling, but the welding procedure was complicated by the fact that the superconducting coil would be in position inside the vessel during the welding operations. Four circumferential welds were required with full depth penetration so that there was a serious risk of overheating the supercon­ducting material with consequent permanent damage. A compromise had therefore to be reached between the small number of heavy current weld runs required for good impact strength and the larger number of light runs required to reduce the heat input.

The temperature problem was further eased by the device adopted to reduce the differential contraction stresses on cool-down. A coil wound directly onto the inner cylinder/former would produce high local stresses at the ends of the winding in the former adjacent to the welds, and these stresses would be in the same sense as the stresses resulting from the internal pressure. A number of longitudinal spring laths were therefore introduced between the former and the winding which would distribute the load more evenly on the former and, also, introduce an additional thermal resis­tance between the weld area and the winding.

With a correctly designed system of laths, the resulting shrinkage stresses could be treated as a uniform pressure on the inner cylinder so that the effective internal pressure on the inner cylinder was 135 lb/in2 abs. and on the outer cylinder, 115 lb/in2 abs. (7.7 atm). One further load on the inner cylinder was the weight of the winding, resulting in a maximum of 7 lb/in2 additional pressure at the top of the cylinder.

These pressure conditions roughly defined the thickness of plate required for the cylinders, and the main problem became the end closure of the toroidal vessel.

After examining various methods ranging from a thick, flat plate closure to a costly dished end (which proved to be unobtainable), a design was evolved which made use of flat end plates suitably profiled to provide elastic hinges which would relieve the stresses at the welded corners, potentially the critical areas (fig. 2).

J-prep weld

m ; ; ; ; ; ; ; ; ; ; ; ; / ; / ; ; / / / /

Fig. 2 — Inner vessel detail.

The cylinders were then analysed more accurately using standard cylindrical shell theory and the end plates treated from first principles by analysing a elemental wedge-shaped portion of the plate. The final design required f in. plate for the cylinders and 1 in. profiled end plates.

279

Location and suspension of the inner vessel within the outer casing was critrical in respect of heat leakage and of accurate positioning of the coil, it being required to have some control over the position of the magnetic axes of the coil. The system chosen employed four main support cables, so designed that any two would take the total 7-ton weight of the coil vessel, and lighter transverse and longitudinal ties. A system was devised whereby these cable supports could be passed through the outer casing and means of adjustment provided which would not entail breaking the vacuum within.

To obviate movement of the coil in a vertical plane on cool-down, the attachment of the support cables was arranged so that the thermal contraction of the cables and that of the vessel itself cancelled out.

The current leads to the coil enter the coil vessel through a port at the top which is fitted with a vacuum insulated stand pipe or chimney. The liquid helium extends up into the neck tube of this chimney (about 6 in. diameter) to give a more responsive surface for level control measurements.

The leads are quite simple, consisting of copper tubes with internal extended sur­faces for good heat transfer. They are cooled by a stream of boil-off gas which warms up to ambient temperature in passing through the leads.

OUTER CASING

The outer casing of the cryostat was to be basically a simple vacuum vessel, but, in addition, it was required to transmit the full load torque reaction of the motor to the foundations. For reasons explained in Mr. Appleton's paper, this torque appears on the stator frames and it was decided that these should be mounted on the cryostat casing, but in such a way that no forces appeared across the demountable main cover plates of the vacuum vessel.

This meant, therefore, that the cryostat casing was to function also as the main location between the bearing housings at each end of the motor shaft. Problems of deflection and machining were significant in the design and the final design incor­porated location devices which were not affected by the dimensional changes occuring between the machining operations and the finished assembly.

The stator frames were bolted to reinforcing flanges fitted to each end of the outer cylinder of the casing and these transmitted the loads directly to two feet mounted on the machine foundations.

Being a vacuum vessel, the inner cylinder could be made much thinner than the outer cylinder and is in fact made from \ in. plate, whereas the outer cylinder is 1 in. thick. The material throughout is stainless steel, AISI type 321.

For ease of assembly and possible future dismantling, a bolted-up construction was decided upon with circumferential 'O'-rings sealing the annular end cover plates to the cylinder flanges.

The vessel was to be maintained under a vacuum of 10~6 torr by means of a single large diffusion pump and backing pump. A 14 in. diameter port was provided in the outer cylinder for mounting the diffusion pump.

RADIATION SHIELD

Calculations showed that the optimum thermal insulation was achieved by the use of a single refrigerated radiation shield at between 60 and 100°K with multi-layer superinsulation on the warmer side and none on the helium vessel side, the whole inter-vessel space being evacuated to better than 10~5 torr.

To simplify the refrigerator system and its supplies, helium gas at 80°K was chosen as the cooling medium for the shield. This also avoided the two-phase flow

280

problems which would have occurred in a shield of this size with the more common liquid nitrogen coolant, and the temperature fitted in very well with that of the first expansion stage of the refrigerator.

Theoretically, a copper shield would have been marginally superior in thermal performance, but stainless steel was again chosen, primarily because its high resis­tivity eliminated the danger of large circulating currents and collapsing forces in the shield on sudden coil discharge. It is also much easier to maintain the required polish on stainless steel than on copper during manufacture.

THE REFRIGERATOR

For a motor which was to replace an existing conventional motor as a demonstra­tion prototype, it was essential that all the ancillary equipment should be as simple and reliable as possible. With normal power station operation in mind, it was decided that the refrigerator should be fully automatic in operation. The design chosen is in fact capable of fully automatic cool-down and filling of the cryostat as well as auto­matic steady-state operation.

The refrigerator cycle, designed by British Oxygen Cryoproducts Ltd. in close collaboration with IRD, uses a low pressure turbine expansion system which is able to supplygas at maximum cycle pressure to maintain the cryostat radiation shield at 80°K and to provide, in steady state conditions, a continuous supply of liquid helium to the coil vessel. The major part of the boil-off gas from the coil vessel is returned to the refrigerator at or near its saturation temperature, but a portion of it is used to cool the coil current leads and is then returned at ambient temperature.

Cryostat

Transfer lines

Storage dewar

Refrigerator

/ / / / ; ; / / /

Fig. 3 — System layout at Fawley.

7~7

281

The specified refrigeration loads are 27 watts at 4.4°K+100 watts at 80°K + 220 watts lead cooling between 4.4 and 300°K.

As mentioned above, a high pressure supply of refrigerant was required during cool-down of the coil and this was very neatly provided by a diverter valve immediately upstream of the Joule-Thomson (J-T) valve at the cold end of the refrigerator.

Control of this diverter valve is entirely automatic, a temperature sensor causing the valve to switch from cryostat to J-T stream when the coil temperature is down to 5°K. In this condition the high pressure gas is throttled to its saturation pressure and temperature and the wet vapour passed to a separator in a liquid helium reservoir feeding the cryostat.

The operating specification of the motor called for maintenance-free running of ancillaries, including the refrigerator, of up to 5 000 hours. If the refrigerator did require attention within this period, the supply of liquid helium to the coil was to be maintained whilst repairs were carried out.

A liquid reservoir was therefore provided using a standard 1 000 litre storage dewar normally maintained half full to allow 12 hour shut-down of the refrigerator. The reason for the extra capacity was to allow for decanting of the liquid in the coil vessel into the storage dewar should any work on the coil become necessary.

SYSTEM ENGINEERING

Because of the layout of the prototype, involving low temperature transfer lines, and because of various safety margins in the design, the refrigerator is a good deal larger than it theoretically need be. The actual capacity is probably about three times that required for a fully integrated production design, so that production machines would be considerably smaller and cheaper.

This meant that vacuum insulated transfer lines with valves had to be provided between refrigerator, storage dewar and cryostat, thus increasing the required refrig­erator rating by 50%. The particular layout in Fawley power station (fig. 3) allowed the refrigerator, with its top-entry pipework, to be situated below the level of the cryostat which was mounted on the 12 ft high motor plinth.

To reduce as far as possible the length of transfer line exposed to ambient radiation, the liquid feed and gas return lines entered the cryostat at the point closest to refrige­rator terminals, continuing to bottom and top of the coil vessel inside the liquid helium space (fig. 4).

The liquid line from the storage dewar to the cryostat was fitted with a proportional control valve actuated by a level sensing differential pressure cell on the cryostat. This system was designed to control the level to within one in., and was backed up by low and high level alarms actuated by carbon resistors in the cryostat chimney.

A similar level control was fitted to the storage dewar to regulate the output of the refrigerator by controlling the position of the J-T valve.

The gas emerging from the two coil leads require to be maintained at ambient temperature and is therefore passed through a temperature controlled flow valve on each lead outlet before joining a common return to the warm end of the refrigerator.

The cool-down of the whole system would have been best achieved by cooling cryostat and refrigerator together, but a slight variation was called for in order to fill the cryostat with liquid in a reasonable time without having to provide an excessive degree of over-capacity in liquefaction.

The procedure adopted was first to cool-down the refrigerator and dewar and then to fill the dewar to about 98% full. A level sensor on the dewar would then switch the refrigerator diverter valve to the high pressure cryostat position and the cryostat would then proceed to cool down after initially causing a slight temperature rise in

282

the refrigerator which would then cool-down again with the cryostat. On reaching a coil temperature of 5°K, the diverter would again switch to the tank filling position and the liquefaction rate would be automatically reduced. During this period, the valve controlling the supply of liquid to the cryostat from the tank has been manually over-ridden and is now released, allowing liquid to fill the cryostat.

Lead cooling ,

Compressor ♦

Exp. L-j r-+ h turbine 2 L I < £

Diverter valves

Fig. 4 — Refrigerator-cryostat system.

The flow is set up by small pressure differential maintained between dewar and cryostat by a pressure bias valve on the return lines in the refrigerator. The pressure differential can be reversed by a manually operated switch should it be required to transfer the liquid from the cryostat back into the dewar.

ACKNOWLEDGEMENTS

The authors wish to acknowledge that this work is supported under a contract from the National Research Development Corporation

283

SEMI-SUPERCONDUCTIVE ROTARY MACHINE

M. YAMAMOTO Central Research Laboratory, Tokyo Shibaura Electric Co. Ltd., Kawasaki {Japan)

Une machine rotative semi-supraconductrice

RESUME : La demande d'energie electrique s'accroit tous les ans et il est possible qu'on ait besoin en Van 2 000 de generatrices de plus de 10 000 MW. Mais il serait impossible de realiser une generatrice de type classique d'une telle puissance en utilisant le fer et le cuivre. Une bonne facon de resoudre ce probleme est d'utiliser une gendratrice supraconductrice. Par Vemploi des techniques supraconductrices il serait possible d'augmenter dix fois le champ magnetique et la densite de courant, c'est-a-dire que la generatrice supraconductrice donnerait 100 fois plus d'energie qu'une machine classique de semblables dimensions. II n'est cependant pas facile de realiser a present une generatrice completement supraconductrice. La machine qu'on pourrait realiser maintenant serait d'un type semi-supraconducteur avec une bobine inductrice supra­conductrice et un induit de type classique. Nous examinerons ici les facteurs les plus import ants pour la mise au point d'une telle machine tout en decrivant une petite generatrice pilote que les auteurs sont en train de construire.

G E N E R A L CONSIDERATION

The growth of economical activities has been accompanied by an increase in unit size of machines. The annual demand of electrical power is increasing at the rate of 7% per year in the world and 10% in Japan. As a result, the demand in 2000 will be about 10 times the present demand. The rate of increase of the electrical power generating unit is more than the demand for electrical power; for example, the peak electrical loads in the U.S.A. have increased 4 fold in past 20 years. During the same period, the average unit size has increased more than 7 fold from 38 MW in 1947 to 280 MW in 1967 and the largest unit size has increased from 100 MW in 1947 to 1,000 MW today, besides there is a 600 MW unit under operation in Japan. If this same pattern were to continue, the largest unit size to be installed in 1985 might be about 6,000 MW and 30,000 MW in 2000. However it is very difficult to visualize this, because the enlargement rate of conventional generators which utilize iron and copper is decreasing. An answer for such high rate enlargement of unit size is to use super­conductive machines. By use of this kind of machine, 5,000 MW should be possible at least in the 1990 s and 10,000 MW at the beginning of the 2000 s.

The output capacity of an electrical rotary machine is proportional to the product of magnetic field flux density B (B = H in a superconductor) and current density in the armature J. B and J can be increased by the use of superconductive wire. At present, there are two kinds of practical superconductive wires, Nb-Ti alloy and Nb3Sn compound. Nb-Ti wire could produce economically 60~70 kG in a working space as will be described below and Nb3Sn tape could produce twice this. These field strengths are respectively about 5 times and 10 times as large as those in a conventional machine.

The permissible current density of a conventional machine is variable according to the cooling conditions, but its value is generally about 10 2 ~10 3 A/cm2. On the other hand, a stabilized superconductive wire could allow even more than 104 A/cm2

average current density. These two facts mean that a fully superconductive machine which has superconductive field and armature coils could have about 100 times more capacity in about the same size in comparison with the conventioanl one. But it is not easy to develop fully superconductive machine, because we must solve some problems to realize it; for example, a development of a superconductive wire with lower a.c.

285

loss ana higher critical temperature, a bearing of no helium gas leakage, a shaft of low heat transfer, etc. Therefore, it is very difficult to realize the answers to these problems in the near future.

Under these circumstances, the machine which we could now develop is a semi-superconductive machine which has a superconductive field coil and a conventional armature. The technique of large superconductive magnets is established. At the present time, 2 magnets for bubble chambers, as large as 12 and 15 feet, are being constructed in the U.S.A. This type of magnet with Nb-Ti field coils could generate 5,000 MW which might be desired in the 1990s and more than 10,000 MW with Nb3Sn in the 2000 s. Still more, it is better to develop the fully superconductive machine before the 1990 s or the 2000 s. By the way, with steam turbines, the important factors to realize the large unit are a blade length of the last stage and the number of steam flows, the maximum present blade length is 52//. By the use of Ti-alloy, a 65" blade could be made, and then, 8 flows might be possible instead of present 6 flows. A combination of 65" blades and 8 flows could make a 5,000 MW steam turbine. However, it might be difficult to realize a larger one than this, because the vapour density of steam is very low in low temperature. For large units, a combination of a helium gas turbine and a high temperature gas cooled reactor should be aimed.

It is clear that the future large capacity generators are feasible by the use of the superconductivity. But at the present moment, it is appropriate, before trying to study these large ones, to study the semi-superconductive d.c. machine. These machines provide broader and more general information in the field of superconductive tech­nology. I would like to describe in the following the important aspects of the semi-superconductive machine, especially the magnet.

GENERATION OF THE MAGNETIC FIELD

In order to design the high field magnet, it is necessary to understand the charac­teristics of the superconductive wire and the relationship between the maximum field strength in coil and the field strength of working space. These problems are discussed in the following, taking Nb-Ti wire as an example.

a) The Hc-Ic characteristics

The Nb content of commercial NbTi superconductive wire is between 20 and 50 %, and the Hc-Ic characteristic is shown in figure 1 where the Ic value at 50 kG is taken as unity. The higher the Nb content the higher is Ic. Figure 1 also shows that 50% Nb wire allows the high current density even in 100 kG.

^ _o ^ 4-»

c c <D 3

1 s> O CD

■!-» 03 IS

"<t-* CO 'ZL >* -' O

1.

0.

0.

0.

0.

0

8

6

4

2

0

critical field strength (Hc) in kG Fig. 1 — Dependence of the Hc-Ic relation on Nb content.

J I I I I L

286

In order to measure the Ic of a wire in a high field magnet, we made a small superconductive magnet and tested it as shown in table 1. T-48 wire of 14 mil core diameter is used on the inside part of the coil and wire made by our laboratory on the outer part. The maximum field strength obtained in the centre of the bore is 94 kG and the critical current is 35 A. In a small magnet like this, the field strength in the bore is about the same as that in the inside layer of the coil. From this test result we see that NbTi wire with high Nb content could carry a current density of nearly 5 x 104 Acm~2 in the core proper and in a field as high as 90-95 kG, and that an average current density of 104 Acm- 2 could be possible over the whole conductor cross section.

Table 1 DIMENSIONS AND TEST RESULTS OF TRIAL SMALL MAGNET

Coil inside diameter 23 mm Coil outside diameter 113 mm Coil length 100 mm Used wire: 14 mil core dia Nb-Ti Field strength obtained 94 kG Critical current 35 A Coil type Solenoid

b) The maximum field strength in the coil There are two kinds of coils for rotary machines, that is, saddle type coil and

solenoid coil, the former being used for the conventional machine and the latter for the homopolar machine. For both coils, the maximum field strength in the coil, Hm, is generally 1.3~1.5 times the average field strength in the working space, Hw. If practical maximum field strength used in the coil is taken 90~95 kG, the average field strength in working space might become 60~ 70 kG. In the following, this relationship will be described, taking a solenoid type coil as an example.

m i n i m u m volume line

1. 0 1. 2 1. 4 1.6 1. 8

a j x C u r r e n t Densi ty

Fig. 2 — Hm/Ho against a, (3.

The relationship between the maximum field strength in coil, Hm, and that at the coil center H0, is shown in figure 2. The diameter of the coil for the actual machine might be W 2 m . In such a configuration, the Fabry factor F should be between 0.05 and 0.1. For these two F values a and P are 1.12 and 0.37, and 1.21 and 0.47 respectively.

287

The corresponding values of Hm/H0 are 2.0 and 1.5. H(«), the field strength at a distance n/100 x a t from the axis is given by

H(,)-"-[Hi5)1 + K(l5i)4+-] e 2 , £4 ... are coil factors. In approximate calculation, terms higher than 2nd term, can be neglected, so

H(w) is shown as

If a working space is taken in the range of n = 20~90%, from these relationships Hm/Hw becomes about 1.3~1.5.

It is difficult to calculate the general solution of the field strength distribution of saddle type coil which has a complicated coil shape. But we can find the relation­ships between Hm and Hw, by simplified calculations or experiments. These relationships are reported in case of the saddle type magnet for MHD power gene­ration. In these reports, we can find about same relationship between Hm and Hw as that of solenoid one.

SEMI-SUPERCONDUCTIVE D.C. MACHINE

Most of the conventional d.c. machines belong to shunt, series or compound type, and one seldom finds a homopolar machine. However, the homopolar machine has many advantages, for example, the simple and strong construction of armature, no commutation problems, less armature reaction, higher efficiency, etc. But, its characteristics are limited in the high current and low voltage region. If this limitation of the characteristics is overcome by the superconductive technique, the homopolar machine might find a wide application. In the following, we would like to discuss this problem.

The application range of the actual semi-superconductive machine might be more than 1,000 kW. In such a large machine the terminal voltage is generally 200—750V. The induced voltage is given by E = B x / x V , and taking B = 7T, / = .75 m and V = 75 ms~ *, we find that for two discs, 2E = 800 V. It means that the semi-superconductive homopolar machine could be used for general purposes by the superconductive technology. But, sometimes we have to design the machine of low r.p.m. and high voltage, or small size and high voltage. In such cases, the same considerations are necessary. For example, segmented discs, split discs, etc. In the segmented disc type, the discs are radially segmented into several conductors and each conductor is connected in series by the low voltage drop brushes. In the split disc type, several discs are split, the induced voltage in each disc is connected like segmented one.

In a semi-superconductive generator, we must endeavour to make the magnet volume minimum to save superconductive wire and liquid helium. As described above, in the coil for an actual machine, the Fabry factor is generally 0.1-0.05, and from this, P = 0.35-0.5. That is, coil length is 0.35-0.5 times coil diameter. By the way, in the segmented discs, it is difficult to use more than 2 discs, because of the construction of the brushgear. Therefore, the coil length with P = 0.35-0.5 is too long and the magnetic field space is not effectively used, but this type of machine has the advantage of simple armature construction, especially easy exchange and checking of brushes.

In view of an effective use of field space, the split disc type is better. This type is often used in the homopolar machine with NaK brushes. The problem in this machine

288

is the construction of the brushes, including their checkings and exchanges. In order to study the characteristics of the semi-superconductive machine, including these pro­blems, we are now making a split disc type small semi-superconductive machine. Its rating is shown in table 2.

Table 2 TRIAL SEMI-SUPERCONDUCTIVE D.C. GENERATOR

Superconductive coil of field Inside diameter 215 mm Outside diameter 300 mm Field strength 45 kG Wire: copper clad multi-core Nb-Ti

wire 2.0 x 3.0 13 strand Coil type continuous discs

Armature Split discs Number of discs 8 Diameter of disc 120 mm r.p.m 3 000

Brush High silver content brush Voltage drop approx. 0.05 V at 15 A/cm2

In a superconductive coil, one should avoid connections as far as possible. Our superconductive coil was wounded in a continuous discs coil by the use of a trans­former winding technique, In spaces between each disc; 2 mm thickness, epoxy resin bonded glass fibre spacers are inserted to keep a flow channel and to hold tightly the coil. The superconductive wire is a composite conductor with 13 superconductive strands in 2.0 x3.0 mm OFHC substrate. The insulation between turns is 0.1 mm thickness mylar tape stuck on substrate. The critical current of this wire is about 850 A at 45 kG. The coil is designed to give 45 kG at 750 A. The cryostat which contains this coil, is made from stainless steel and has a room temperature bore of 150 mm diameter for the armature. This machine is not yet in operation. But we expect to run the machine in the near future with a load of several kW and thus get valuable information about semi-superconductive rotary machines.

CONCLUSION

We have described above the semi-superconductive motor as a machine of the present and discussed the future potentialities of fully superconductive machines. Next we have to go to the a.c. machines for example alternator and transformer, etc. For them, it is necessary to develop a low a.c. loss and high Tc wire. At present Nb-Ti single wire with a Tc of about 10 °K has about mW/cm3 of a.c. loss at 1 kG or 104

A/cm2 and 50 Hz at 4.2 °K. It is important to improve both characteristics. The compound type wire might give a good answer. Nb3Sn tape has a Tc of 18.1 °K and the measured a.c. loss is smaller than in Nb-Ti wire. The superconductive machine is now in its cradle, but we believe that it should grow up to a giant of the future electrical industry.

REFERENCES

[1] R. R. BENNETT, Planning for power—a look at tomorrow's station sizes. IEEE Spectrum (Sept. 1968).

[2] D.B. MONTGOMERY, Some useful information for the design of air-core solenoids. The report of National Magnet Laboratory M.I.T. (1961).

289

TRAVAUX DE DEVELOPPEMENT EN VUE DE LA CONSTRUCTION D'UN PROTOTYPE

DE CRYOTRANSFORMATEUR

P. LAIR Alsthom-Savoisienne, Saint-Ouen

S. L E H O N G R E

Centre a"Etudes Cryogeniques de L'Air Liquide, Sassenage

et

J. B O N M A R I N Groupe Pechiney (France)

Development work for the construction of a prototype cryo-transformer

SUMMARY: The work undertaken jointly by the companies Alsthom-Savoisienne, VAir Liquide and Pechiney are concerned with the study of the possibilities of using for transformer windings refined aluminium maintained at 20 °K by immersion in liquid hydrogen.

In its first stage this work has for its aim the study and the realisation of a prototype medium tension cryo-transformer of a few MVA.

The authors deal with the following aspects: the general arrangement envisaged for this prototype, new shapes of aluminium conductors, structure they propose to adopt for the windings, electric insulation and the fabricated containers.

They will finally describe some of the specialised cryogenic devices for maintaining in this apparatus the windings at a low temperature and for providing the necessary cooling power.

I . OBJET DE L'ETUDE

Des travaux de developpement sont menes en France conjointement par les Societes Alsthom-Savoisienne, L'Air Liquide et Pechiney. Ils concernent 1'etude des possibility de mise en ceuvre dans les tranformateurs d'enroulements en aluminium super-raffine portes a 20K par immersion dans l'hydrogene liquide. La Societe Alsthom-Savoisienne est maitre d'ceuvre et est chargee en particulier des etudes generates et de la mise au point de la partie active et des enveloppes; la Societe L'Air Liquide des etudes et realisations relatives aux dispositifs cryogeniques et le groupe Pechiney de la production, de la purification et de la transformation du metal conduc-teur dans les formes directement applicables aux cryotransformateurs. En premiere etape, ces travaux avaient fait I'objet, en octobre 1964, d'une journee d'etudes organisee par la l r c Section de la Societe Francaise des Electriciens [1].

Depuis 1966, l'action de recherche et de developpement a pris une orientation plus precise dans le cadre du programme de la Delegation Generate a la Recherche Scientifique et Technique. Cette orientation consiste en la definition d'un objectif : Etudier et realiser pour essais un cryotransformateur prototype monophase de 15 MVA a 63 kV. L'adoption d'une puissance et d'un echelon de tension certes limites pourra permettre, dans une certaine mesure, de juger de l'avenir de cette technique et, eventuellement, de definir par la suite des objectifs plus ambitieux en cas de succes. Ce niveau de puissance a ete determine de facon que les flux de fuite dans les enroulements du prototype soient representatifs de ceux des transformateurs de grande puissance. La tension de l'enroulement primaire a ete choisie de l'ordre de 15 kV tension voisine des reseaux de distribution ou de celle des generateurs de puissan­ce electrique alternative. La tension de l'enroulement secondaire 63\/3 kV a ete

291

determined de fagon a aborder les problemes dielectriques a des niveaux de tension relativement eleves.

Cet appareil comportera done :

— un circuit magnetique fonctionnant dans Fair ambiant et dont la technique est parfaitement connue;

— deux enroulements en aluminium raffine maintenus aux environs de 20 K, qui sont le siege de pertes electriques;

— un systeme d'enveloppes en materiau isolant avec superisolation thermique sous vide pousse;

— des structures realisant l'isolation electrique; — des traversees assumant un double role :

— transfert de puissance electrique (courant — tension) — passage de Fambiante a 20 K;

— d'un systeme de refrigeration permettant l'evacuation des pertes et le maintien des enroulements a la temperature prescrite.

Les differents composants que nous venons d'enumerer feront chacun l'objet d'un chapitre precisant les problemes poses et certaines des solutions que nous y avons apportees dans l'etat actuel de nos travaux.

II. ENROULEMENTS EN ALUMINIUM SUPER-RAFFINE

L'aluminium super-raffine est produit par le procede de raffinage electrolytique a 3 couches. La selection des productions des cuves d'electrolyse permet de ramener ces impuretes aux valeurs suivantes : Fe : 3 ppm, Si: 2 ppm, Cu : 5 ppm (titre 99,999 %)

Cet aluminium super-raffine (10 ppm) nous a permis de mesurer a basse tempera­ture (20 K) sur des enroulements des resistivites en courant continu 500 a 800 fois inferieures a celles des enroulements en aluminium de conception classique a tempe­rature ambiante (compte tenu des phenomenes parasites majorant la resistivite) [1],

Pour etre utilisables avec profit en courant alternatif dans les cryotransformateurs, ces conducteurs doivent satisfaire a des caracteristiques physiques precises et repon-dre a certains imperatifs technologiques.

Pour le metal que nous utilisons, la profondeur de penetration en courant alternatif est 25 fois plus petite, environ a 20 K pour une frequence donnee. Si nous voulons done maintenir a une valeur satisfaisante les pertes en courant alternatif, les dimensions a retenir pour la section de conducteurs elementaires doivent etre sensiblement homo-thetiques des dimensions classiques dans le rapport 25. Ainsi l'ordre de grandeur des dimensions requises est environ de 50 | ix400u.

Etant donne le nombre eleve des brins elementaires a disposer en parallele, il y a lieu de rechercher les technologies qui permettent de realiser un conducteur global directement utilisable dans la fabrication d'un enroulement.

Les formes de conducteur global actuellement developpees sont les suivantes : — Assemblage sous forme de tissu, dont la chaine est constitute alternativement de

brins meplats d'aluminium et de fils de verre et la trame de fibres de verre.

Dans le but de diminuer l'epaisseur du conducteur global par rapport a celle du tissu de brins d'aluminium et de iils de verre, tout en conservant a 1'ensemble la dispo­sition parallele des bandes minces entre elles, nous nous sommes efforces de developper une solution industrielle deduite de la technique de realisation des circuits imprimes.

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Celle-ci consiste a deposer un cache sur une feuille d'aluminium super-raffine contre-collee avec du mylar et a attaquer chimiquement l'aluminium afin de dissoudre les parties conductrices non cachees.

— Assemblage sous forme de torons de flls d'aluminium super-raffine de diametre de 100 \i. Le trefilage de tels fils s'est avere jusqu'a present difficile sur une longueur appreciable par suite des faibles caracteristiques de l'aluminium super-raffine. Ces fils ecrouis par le trefilage, l'emaillage et le toronnage doivent etre recuits a temperature convenable apres ces operations.

Ces problemes semblent actuellement avoir trouve une solution satisfaisante.

La realisation d'enroulements a partir de tels conducteurs globaux se ramene alors a des techniques connues de bobinage qu'il s'agisse de tissu Verre-Aluminium ou d'une bande continue « imprimee » la technique releve de celles actuellement connues d'enroulement en feuilles. Neanmoins, dans ce dernier cas, compte tenu du fait de la faible resistivite mise en jeu, il y a lieu d'apporter un soin particulier aux transposi­tions des conducteurs dans la hauteur des bobines.

Avec les torons, ces enroulements peuvent etre soit « en galettes » soit en « cou­ches » tels qu'ils le seraient actuellement avec des cables transposes par exemple.

III. LES ENVELOPPES

Les enroulements sont isoles du milieu ambiant et du circuit magnetique au moyen d'une enveloppe a double paroi de forme annulaire. La paroi froide ou cuve interieure contient les enroulements et l'hydrogene liquide. La paroi chaude ou cuve exterieure contient la precedente. Un vide pousse et une super-isolation thermique appropriee entre les deux cuves assure l'isolement thermique de le cuve interieure. Chaque cuve realisee en stratifie verre-epoxy est constitute par un assemblage de 2 cylindres concen-triques fermes par 2 flasques plans. Les liaisons entre cylindres et flasques se font soit par collages, soit par joints demontables. En regime de fonctionnement permanent, la cuve interieure est soumise a la pression interne necessaire a la mise en oeuvre de la refrigeration. En regime de court-circuit, l'energie dissipee peut produire une vapori­sation partielle de l'hydrogene liquide contenu dans les enroulements donnant une surpression qui se combine avec les forces electro-dynamiques radiales provenant des enroulements.

Ces phenomenes impliquaient l'etude d'un systeme de calage susceptible de trans-mettre les efforts au circuit magnetique. Ce calage est dispose entre le tube central de la cuve interieure, tube repere 2, et le tube central de la cuve exterieure, tube repere 1, (fig. 1). Son role est double, en regime normal il centre la cuve interieure par rapport a la cuve exterieure, en regime de court-circuit il transmet les efforts radiaux du tube 2 qui est lui-meme cale sur le circuit magnetique.

L'ensemble du dispositif adopte se presente sous forme d'une « cage d'ecureuil» qui peut etre emboitee entre les cylindres 1 et 2. Les cales sont constitutes par des anneaux de stratifie espaces regulierement au moyen d'entretoises disposees paralle-lement aux generatrices. Chaque cale annulaire porte des dents qui appuient sur le cylindre n° 1. L'experimentation realisee sur maquettes cylindriques et planes de petites dimensions a permis de definir la nature des stratifies et les valeurs des contrain-tes de compression admissibles sur les surfaces d'appui. Les raccordements entre parois cylindriques et couvercles s'effectuent au moyen de joints colles demontables. Le stratifie plan et le stratifie cylindrique sont assembles en construisant, sur la ligne de joint, un stratifie appele frette. Ce stratifie porte sur les 2 pieces a assembler et assure la fonction d'etancheite.

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L'etancheite au vide de ces types de joint a ete controlee avec succes a la tempera­ture de Thelium liquide sur des eprouvettes de dimensions reduites. L'etancheite a la temperature de l'azote liquide a ete verifiee avec succes sur une cuve maquette sensible-

TubeJ

Fig. 1

ment a l'echelle !/2 du prototype envisage. Le passage des conduits cryogeniques, ou des connexions electriques a travers la paroi de la cuve interieure est realise de la maniere suivante :

— un cylindre en acier inoxydable, d'un diametre compris entre 100 et 250 mm, est pique sur le couvercle de la cuve. La base de ce cylindre est raccordee a la paroi de facon etanche et non demontable, l'autre extremite du cylindre porte une bride metallique annulaire munie d'un joint de type classique (joint indium par exemple). Ce joint permet de raccorder, de facon etanche et demontable, soit une autre bride

294

metallique, solidaire du fuseau isolant d'une traversee electrique, soit une plaque d'obturation metallique rassemblant les passages des tuyauteries cryogeniques.

IV. ISOLATION ELECTRIQUE

Le probleme de l'isolement electrique du cryotransformateur est lie tres etroitement a la conception meme des enroulements et par consequent a celle de l'appareil. Compte tenu du fait que Ton doit concevoir et realiser une isolation haute tension pour le prototype, il est necessaire de resoudre les problemes poses par la tenue des structures isolantes et non pas seulement l'etude des materiaux eux-memes [2].

En fonction de l'experience acquise avec l'isolation papier huile, il a semble indis­pensable d'essayer des maquettes dont les dimensions sont du meme ordre de gran­deur que celles des parties du cryotransformateur qu'elles doivent figurer et concues pour des tensions voisines de la tension reelle a tenir.

Dans une premiere etape, l'isolation prevue est adaptee a la solution qui utilise pour les enroulements les conducteurs en tissu de verre/aluminium. Le materiau de base est le papier impregne de resine epoxy a l'image des enroulements maquettes en tissu. L'isolation principale du cryotransformateur ne comporte pas de larges canaux d'hydrogene liquide. Dans l'etat actuel de nos travaux, nous disposons d'une isolation interne d'enroulements qui donne sastisfaction. L'isolation entre enroulements et entre enroulement et masse est en cours de mise au point. Dans l'hydrogene liquide, une premiere approche jusqu'a 60 kV a donne des resultats tres corrects, les maquettes prevues pour 140 kV sont actuellement en cours d'essais. Ces essais sont accompagnes de mesures de decharges partielles.

V. CRYOTRAVERSEE

Les connexions electriques sont soumises, en plus des contraintes electriques classiques, a des imperatifs d'ordre mecanique et thermique : — Tenue mecanique lors des contractions thermiques et transmission d'un flux de

chaleur minimal entre l'ambiante et le milieu cryogenique. Les entrees de chaleur par les connexions doivent etre reduites le plus possible et il convient de de dimensionner correctement les passages de courant. Les resultats obtenus ont fait I'objet d'une communication au XIP Congres International du Froid a Madrid en 1967 [3]. Des essais dielectriques entrepris avec des traversees conden-sateurs en papier phenoplaste semblent, jusqu'a present, apporter une solution satisfaisante.

Deux traversees de cette conception dont la tension nominale etait de 36 kV et la tension d'essai sous la frequence industrielle de 75 kV ont donne entiere satisfaction lors d'essais prolonges. Deux autres traversees (72 kV et 140 kV respectivement) ont fait I'objet d'essais preliminaires dans 1'azote liquide et se sont comportees de fagon encourageante.

VI. REFRIGERATION DU PROTOTYPE [4]

Principe :

Les pertes totales du cryotransformateur (somme de la puissance dissipee dans les enroulements et de celle absorbee par les moteurs d'entrainement des compresseurs du refrigerateur et eventuellement de la pompe de circulation) sont, non seulement fonc­tion de la purete de l'aluminium et de la structure retenue pour les enroulements, mais

295

encore de la maniere dont la refrigeration est produite et transmise a ceux-ci. II faut produire les frigories a un niveau de temperature adequat a l'aide d'une installation dont le facteur de merite (quotient de la consommation d'energie d'un cycle reversible et de la consommation d'energie du cycle reel) soit le plus eleve possible.

II faut aussi transferor cette puissance frigoriflque a la machine dans les meilleures conditions. Ce transfert est effectue par circulation d'un fluide d'echange entre le refri­gerateur et le transformateur. Dans 1'evaluation des performances globales, il faut tenir compte des ecarts de temperature entre le fluide d'echange et les conducteurs et entre le fluide d'echange et la source froide. II est done absolument essentiel d'assurer de tres bons echanges de chaleur, en particulier dans les enroulements du cryotrans­formateur. La mise en circulation du fluide d'echange necessite par ailleurs une depense d'energie mecanique. Un compromis devra etre trouve entre : — les pertes par ecart de temperature; — les pertes par circulation du fluide et la diminution de puissance specifique des

enroulements due a l'encombrement des canaux de circulation.

L'analyse de revolution des pertes electriques d'un cryotransformateur de puis­sance donnee, ou, plus precisement du parametre de pertes correspondant, fait appa-raitre une plage de temperature allant de 17 K a 23 K pour lesquelles les pertes reelles varient tres peu. Le fluide d'echange peut done etre, soit de l'helium gazeux, soit un reliquefacteur d'hydrogene, soit un refrigerateur a helium. La recherche des conditions optimales de refrigeration a implique l'etude simultanee des performances de ces deux types de refrigerateur et du compromis defini ci-dessus.

En pratique, les problemes technologiques sont plus facilement resolus avec l'helium qu'avec l'hydrogene; et il existe des maintenant des elements du cycle par-faitement au point tels que les turbines de detente d'helium. En plus, pour les unites de refrigeration de tres grosse puissance, les compresseurs rotatifs ont des caracteris-tiques plus interessantes en helium qu'en hydrogene. II a done ete retenu d'etudier un cycle de refrigeration a helium pour le prototype. Celui-ci devant etre refroidi a l'hydro­gene liquide, une boucle de circulation de ce fluide a alors ete prevue pour trans­porter la puissance de refrigeration depuis le cycle vers le cryotransformateur. L'en­semble refrigeration et boucle de circulation d'hydrogene est represents figure 2. L'hydrogene liquide est mis en circulation dans le circuit ferme cryotransformateur echangeur froid du refrigerateur par une pompe.

Echanges de chaleur :

Normalement, la circulation d'hydrogene liquide dans les canaux de refroidisse-ment des enroulements est un ecoulement en phase unique (liquide sous-refroidi pressurise). Neanmoins, une surcharge du cryotransformateur ou une mauvaise repartition des debits entre les canaux de circulation alimentes en parallele est suscep­tible de faire apparaitre un ecoulement en double phase. Pour l'hydrogene liquide, les ecoulements en phase unique sont bien connus et obeissent aux lois clasiques en ce qui concerne tant les echanges de chaleur que les pertes de charge. Par contre, pour ce qui concerne les ecoulements en double phase, peu de travaux ont ete effectues pour l'hydrogene liquide; notamment dans des geometries comparables aux canaux de cryotransformateur (faible diametre et grande longueur). Un dispositif experimental a ete realise pour mesurer les coefficients d'echange de chaleur et les pertes de charge avec l'hydrogene en double phase.

Mise en froid :

Le fait que la temperature de fonctionnement d'un cryotransformateur soit tres differente de la temperature a laquelle il est construit et assemble implique de porter

296

une tres grande attention aux problemes de contraction thermique qui se produisent pendant sa mise en froid. II faut veiller, en particulier a ce que les gradients de tempe-

d T u r b i r\e\

-/t\-

Echanqeur 1

Echanqeur 2.

Echancjeur 3

Tu r b i ne I I

Pom pie L-fyyyyi-J

I—iH—WWH : ryo trans* I \^^y I f o r m a t e u r I . I

R o u c l e nyd roqene ) i

Echancjeur 4

B o u c l e hyd roaene l i c j u i o l e

Fig. 2

rature dans la machine soient faibles de maniere qu'il n'y ait pas de contraintes dues aux dilatations differentielles, susceptibles de provoquer des deteriorations.

II faut prevoir une mise en froid lente et homogene.

297

Sur maquette, ce refroidissement est obtenu par injection dans la cuve d'azote gazeux de plus en plus froid, le remplissage en hydrogene liquide etant realise a partir du moment ou la temperature atteinte est voisine de 80 K.

Pour le prototype une methode plus elaboree sera adaptee a la fois au refrigerateur et a la boucle de transfert choisis.

VII. CONCLUSION

Tous les problemes devant permettre la realisation d'un prototype ont, d'ores et deja, ete abordes. Des points particuliers restent cependant a resoudre pour pouvoir le preciser completement avant de le mettre en fabrication. Le developpement des formes nouvelles de conducteur reste lie aux possibilites actuelles de l'industrie de production et de transformation de 1' aluminium.Les structures dielectriques adoptees doivent etre confirmees par de nombreux essais appropries dans l'hydrogene liquide permettant de maitriser le niveau d'isolement requis. En outre, l'optimisation des structures envisagee actuellement ne pourra etre entreprise qu'a partir du moment oil tous les problemes technologiques poses seront resolus ou pres de l'etre. Cette optimi­sation permettra alors de faire le point des avantages et des inconvenients de Tapport cryogenique a la technique et a la construction des transformateurs de grande puis­sance.

Les premiers travaux entrepris jusqu'a present ont permis d'obtenir des resultats encourageants.

REFERENCES

[1] S.F.E. lre Section — Gros Materiel Electrique. « Journee d'Etudes consacree aux cryo-machines Electriques » RGE (juin 65), Tome 74; (juil. aout 65), Tome 74.

[2] « Contribution a l'Stude des dielectriques aux temperatures cryogSniques » RGE (juin 68), Tome 77, pp. 593-609.

[3] P. BERARD, « Connexions electriques cryogEniques — Aspects theoriques» I.I.F.-XIIe

Congres International du Froid, Madrid (1967) I, pp. 157-171. [4] P. BURNIER et A. de la HARPE, « Applications de l'hydrogene liquide en Electrotechnique »

Cours sur l'hydrogene liquide. I.I.F., Grenoble (1965), pp. 339-373.

DISCUSSION

N. KURTI (U.K.) — Pourriez-vous preciser les avantages des cryotransformateurs en comparaison avec les transformateurs classiques?

P. LAIR — La mise en oeuvre dans les transformateurs de puissance d'enroule-ments en aluminium super-raffine portes a 20 °K apporte les avantages theoriques suivants:

1) Augmentation des densites de courant dans les enroulements (20 a 50 A/mm2), ce qui amene une reduction appreciable de rencombrement de ceux-ci done de ce que les constructeurs de transformateur appellent le « dimensionnement»;

2) Compte tenu de cette augmentation de densite de courant et de la puissance necessaire a la refrigeration, les pertes totales du cryotransformateur sont alors divisees par un facteur k (k = 3 a 4).

3) L'emploi d'hydrogene liquide et d'isolants electriques (solides) adaptes a ce fluide et a la temperature de 20CK devrait amener la reduction des distances d'isole­ment correspondantes par comparaison avec celles relatives a l'isolation classique papier-huile.

298

Mais cette mise en oeuvre pose des problemes technologiques complexes dont certains furent decrits dans notre communication. Apres resolution de ces problemes technologiques, on devrait pouvoir realiser dans un avenir qu'il ne nous est pas possible de preciser actuellement des cryotransformateurs de tres grande puissance et a tres haute tension disposant de pertes electriques et d'un encombrement reduits.

P.H. MELVILLE (R.U.) — Pour comparer ce systeme avec celui qui emploie les supraconducteurs, pouvez-vous me dire la valeur des champs magnetiques dans le cryotransformateur?

P. LAIR — La valeur du champ magnetique de fuite retenu impose les dimensions du conducteur unitaire compte tenu de la resistivite du metal utilise et de la densite de courant qui le traverse. Avec les dimensions que nous avons a l'heure actuelle a notre disposition, nous pouvons admettre des champs magnetiques de fuite de valeur analogue a ceux des transformateurs classiques. Ces champs magnetiques se calculent par les methodes completes que nous employons actuellement et que les constructeurs utilisent normalement. Pour fixer les idees, les maquettes d'enroule-ments actuellement experimentees ont permis de mettre en oeuvre et de mesurer des champs magnetiques de fuite correspondant a des inductions allant jusqu'a 0,25 Tesla.

299

SUPERCONDUCTING RECTIFIERS

R. FASEL and J.L. OLSEN Laboratorium fur Festkorperphysik

Eidgenossische Technische Hochschule, Zurich {Switzerland)

Redresseurs supraconducteurs RESUME : Vutilisation de solenoides supraconducteurs dans un champ magnetique longitu­

dinal exterieur comme elements redresseurs a ete proposee il y a un certain nombre d'annees. Lespertes dans ces redresseurs sont si elevees qu'ils paraissent peu applicables en pratique malgre leur simplicite. On decrit un type ameliore de redresseur lid etroitement au cryotron et Von etudie les pertes dans des prototypes. Ces redresseurs semblent susceptibles d'etre mis au point de facon a pouvoir entrer en concurrence avec les pompes a flux pour le reglage precis et la compensation des pertes des grandes bobines.

The fact that leads into cryogenic apparatus conduct not only electricity but also a proportionate amount of heat makes it seem very desirable to reduce the electrical currents to be carried in and out of a cryostat. In many applications large currents are needed at low temperature, and a number of alternatives to a direct input from room temperature have been proposed. They include the fluxpumps described by Volger and Admiraal [1], Buchhold's [2] switched cryotron rectifier or commutator and a passive rectifier system proposed by the present authors [3, 4]. Both these latter systems are for use in conjunction with a low temperature transformer.

In practice all the devices have losses, and although Atherton [5] has shown that in principle these losses can be made to vanish it has also been pointed out by Wipf [6] that this is only possible for thermally stable devices if they are made infinitely large.

Optimized electrical leads that are not cooled by the evaporating gas will introduce approximately 40 milliwatts of heat per ampere of current into a helium cryostat. With such large losses fluxpumps and rectifier devices of relatively low efficiency have advantages over leads. It has recently become clear, however, that this heat input can be reduced to less than one milliwatt per ampere by good thermal exchange between evaporating gas and lead. This, of course, means that the efficiency require­ments for fluxpumps and rectifiers are greatly increased if they are to compete with cooled leads.

The fluxpumps and rectifiers built to date have the further disadvantage of low output voltages lying in the range of 10-100 millivolts. Thus it is clear that a super­conducting generator or rectifier is unlikely to be useful at the ends of cryogenic transmission lines which have to be operated at several thousands of volts to be efficient.

In spite of this it would seem that there may be a number of special applications for such devices. An example of such an application is that of overcoming small resistive losses in large superconducting coils with local contact losses.

The properties of a superconducting solenoid in an axial field have been used previously in a rectifier element [3, 4, 7]. It was found, however, that the voltage current characteristic of this element caused unacceptable rectifier losses. We have recently studied an improved rectifier element [8]. This is a cryotron-like device where a wire or a band of superconductor is placed along the axis of a solenoid that is connected in parallel with it. An external field HE parallel to this axis acts as control field (see fig. la).

Rinderer [9] has investigated the behaviour of wires in longitudinal fields. It is seen from this work that there is a sudden onset of resistance when the vector [sum of

303

the field at the surface equals the critical field, Hc. The behaviour is best shown in a vector diagram having as axes the longitudinal field, H L , and the field, H,, caused at the wire surface by the current I. Here the first sudden increase of resistance from

Fig. 1 — a) Rectifier element. b) Vector diagram for field at wire surface from a longitudinal magnetic field, HL,

and a field I^ due to the current I. The line through BE A represents a typical working path, and the resistance along this path is indicated.

zero to 1/2 R„ occurs on a circle of radius Hc as indicated in figure lb, and the material has its full normal resistance , R„, when HL ^ Hc. In the ideal case the resistance, R, in a wire can be shown to be given by

R/R„ = i [ l + ( l - o c ) * ] , f o r a < 1,

where a = [ ( H c

2 - H * ) / H ? ] * .

The lines of constant resistance are therefore a set of ellipses which reduce to a circle forR = 1/2 R„.

For the element described above consisting of a wire of radius a surrounded by a coil of n turns per unit length we have

HL = 2 I / a ,

and H L = H e + 4 x n l .

With changing current I the working point in the vector diagram of figure lb moves along a line AEB. In the absence of current the condition of the element is given by E, and on applying current A or B are reached according to direction. In

304

ideal wire there is a sudden jump of resistance to half the normal value at A and B. The voltage, current characteristic for HE ^ 0, is clearly asymmetric, and rectifying properties can be obtained. One way of using such elements is indicated in figure 2.

Load coil

rectifier element

Fig. 2 — Use of a transformer and two rectifier elements to supply current to the load coil.

a E <

3

o

500

Magnetic field H L , Gauss

Fig. 3 — Onset of resistivity near Hc. a) Pb 1% Bi band 0.3 mm x 3.3 mm. b) Copper coated Pb 1% Bi evaporated film, 2 000 A x 10 mm, showing quenching

of surface superconductivity. Hc for this alloy is 620 Gauss.

305

We have pointed out elsewhere that if the circuit shown in figure 2 is used to supply a superconducting coil with current the ratio r| of energy dissipated resistively to that stored in the coil is given by

i l = 2 I R / I L >

where IL is the final load current and IR is the current at which resistance appears for currents in the direction opposite to IL.

It is clear that in the ideal case we can decrease IR/IL indefinitely by increasing HE to a value very close to Hc. In practice this is limited by the spread of the supercon­ducting transition in real materials.

The maximum currents that can be carried are limited by H, < Hc, so that

I L < i f l H c .

Thus a 1 mm 0 wire of lead at 4.2 °K gives a maximum IL = 140 Amperes. Unfor­tunately, it seems impossible to increase the maximum IL by increasing Hc because the available high critical field materials are superconductors of the second kind. In these spread out transitions and hysteresis effects are to be expected. The transition to a superconductor of the second kind also limits the extent to which the resistivity of the element may be increased by alloying.

Even dilute lead alloys raise problems because of surface superconductivity, and we have referred to this in a previous note [8]. It is, however, possible to eliminate surface superconductivity by coating with a normal metal. In figure 3, we show the critical currents for an uncoated Pb Bi band and for a coated Pb Bi film. It is seen that the surface coating eliminates most of the current carrying capacity above Hc.

We conclude that rectifiers having IR/IL < 0.03 can be built wit such elements, and that they can be used for currents up to about 1000 Amperes if fairly wide films are accepted. For higher currents more complicated arrangements than a single wire of film in a coil will be required. In these cases we believe that Buchhold's switched rectifiers [10] will be preferable to the simple circuits described here.

ACKNOWLEDGEMENT

We are grateful to the Maschinenfabrik Oerlikon for providing financial support.

REFERENCES

[1] J. VOLGER and P. A. ADMIRAAL, Phys. Letters, 2, 257 (1962). [2] T.A. BUCHHOLD, Cryogenics, 4, 212 (1964). [3] J.L. OLSEN, Rev. Sci. Instr., 29, 537 (1958). [4] R. FASEL and J.L. OLSEN, Kdltetechnik und Klimatisierung, 19, 274 (1967). [5] D.L. ATHERTON, Cryogenics, 7, 51 (1967). [6] S.L. WIPF, Proc. ICEC. 1 (1967), p. 137 (Heywood Temple, London, 1968). [7] R. FASEL and J.L. OLSEN, I.I.F.—Xllth Int. Cong. Refrign, Madrid (1967)I, pp. 149-155 [8] R. FASEL and J.L. OLSEN, Proc. ICEC, 2, p. 204 (Iliffe, 1968). [9] L. RINDERER, Helv. Phys. Ada, 29, 339 (1956). [10] T.A. BUCHHOLD, Proc. ICEC, 1, p. 133 (Heywood Temple, London, 1968).

DISCUSSION

F. MOISSON (France) — Les limitations d'emploi des redresseurs decrits sont liees a la variation d'une resistance et non d'une inductance. II est possible d'envisager un redresseur constitue par un enroulement supraconducteur et un circuit magnetique

306

saturable commande par un enroulement auxiliaire; un tel systeme permettrait par exemple, avec un montage en pont comprenant 4 inductances, de redresser comple-tement un courant monophase — et des systemes analogues sont concevables pour le courant polyphase. Dans un tel montage, les pertes sont dues au circuit magnetique et les tensions importantes aux bornes des inductances sont peut-etre compatibles avec l'emploi aux extremites de cryoliaisons.

T. BUCHHOLD (W. Germany) — With the system proposed by Prof. Olsen one can charge a coil but not take out energy. In order to take out energy the solution of Professor Volger or a system with switched cryotrons has to be used.

J. SAM MAN (France) — What is the limit of switching speed?

J.L. OLSEN — The switching speed is certainly high enough to deal with a.c. of frequencies of some hundreds of Hertz. Above this one may expect a phase shift in the onset of resistance will introduce undesirable losses.

J. S. H. ROSS (U. K.) — Dr. Olsen mentioned a current limitation of about 1 000 amps. Would he care to mention a similar limitation on field?

J. L. OLSEN — At present we believe that sharp switching characteristics can only be obtained with Type I superconductors. For this reason there is a field limitation somewhere between 500 and 1 500 Gauss.

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MAGNETIC SUSPENSION AND GUIDANCE OF HIGH SPEED VEHICLES

H.T. COFFEY, T.W. BARBEE Jr. and F. CHILTON Stanford Research Institute, Menlo Park (U.S.A.)

Suspension et direction magnetiques des vehicules a grande vitesse

RESUME : On a etudie Vemploi de bobines supraconductrices pour la suspension et la direction des vehicules a grande vitesse. Un support magnetique est obtenu par le moyen de Vinteraction du champ magnetique de bobines monties sur le vehicule avec des courants de Foucault induits dans un rebord guide fixe. On obtient des forces de guidage lateral par le meme mecanisme en fournissant le rebord guide d'embases verticales. II est possible de rendre un tel systeme dyna-miquement stable. Pour les tres grandes vitesses il est possible de calculer ces forces par la methode des images magnetiques, parce que le champ magnetique est limite par la haute frequence efficace a Vepaisseur de peau. Les images magnetiques sont formees comme si le rebord guide etait supraconducteur. On a fait Vanalyse theorique d'un modele a deux dimensions en utilisant la methode de transformations de Fourier. Ces calculs demontrent que la conductivity finie fait entrer un rapport de la vitesse dans la force d'appui magnetique et determine une force de perte de resistance ou de freinage magnitique. Ces expressions ont ete modifiees d'une maniere semi phenomenologique pour mieux s'adapter aux resultats experimentaux. On discute ces resultats et quelques-unes des implications pour les systemes rapides de transport de masse.

INTRODUCTION

The use of superconducting magnets in the suspension and guidance of trains was first proposed by Powell and Danby [1] and subsequent studies have been performed by Guderjahn, et al. [2] and Barbee, et al. [3]. In this paper, the basic principles of the application of superconducting magnets are outlined, and some of the pertinent performance characteristics are discussed.

The motivation for these studies is the proposed construction of passenger trains operating at speeds up to 300 mph between major urban areas 300 to 500 miles apart in the United States. This high speed link would constitute only a part of a more extensive transportation system providing rapid and reliable transportation within the cities and efficiently coupled to the high speed interurban trains. Since the terminals for these trains can be located in the centers of metropolitan areas, the total transit time for such a system can be significantly less than that provided by jet aircraft operating between airports located on the perimeters. Among the auxiliary benefits to be derived from such a system are: (1) more reliable scheduling of tranportation in inclement weather, (2) a significant reduction in air pollutants in densely populated areas, and (3) a significant reduction in the noise level in the cities.

There is ample reason to question the reliability of conventional suspension systems using steel wheels on steel rails operating at 300 mph for prolonged periods. Conse­quently, alternative means of suspension and guidance of these trains are being sought. Two of the more promising alternatives are air cushions and magnetic suspensions.

PRINCIPLES OF MAGNETIC LEVITATION

The principles involved in a magnetic suspension or magnetic levitation (M AGLE V) system are quite simple; complications arise only when they are applied to a complex geometry. The basic principle used here is the repulsive force experienced by magnetic

311

fields of like sign; the sources of the magnetic fields are irrelevant. They may arise from permanent sources such as ferromagnets, quasi-permanent sources such as superconducting magnets, or induced sources such as eddy currents in metallic conductors. The last source is best illustrated by reference to plate I [4] which indi­cates that persistent eddy currents are induced in a lossless superconducting plate and consequently remain stable in time. It is well known that the effect of these diamagnetic currents is to simulate a mirror image magnet below the surface of the superconductor, resulting in a force on the real permanent magnet. If such a magnet were to be placed

Plate I — The floating magnet. The short bar magnet (slightly larger than life size) is floating in helium gas nearly half an inch above the bottom of the superconducting lead bowl; the shadow of the magnet is visible on the right-hand side of the bowl. The bowl is painted white with black lines on it to bring out the perspective and is standing on three copper legs which dip into liquid helium to keep it superconducting (the liquid helium is not visible in the picture). The white specks on the magnet and in the bowl are small pieces of solid air (the air is a slight impurity in the helium). By permission of Cambridge University Press [4].

over a normally conducting surface, eddy currents would again be generated opposing the motion of the magnet. Unlike the superconducting case, however, the currents would soon be dissipated through joule heating and the magnet would fall to the surface of the conductor. Thus, maintenance of a repulsive force between a magnet and a normal conductor requires that they remain in constant relative motion to produce a d(p/dt in the surface of the conductor.

At high relative speeds, the forces developed between a magnet and a ground plane are the same as those developed by a magnet over a superconducting ground plane or between a magnet and its mirror image. Consequently, the high (infinite) velocity limit of the force can be readily calculated if the ground plane is assumed to be flat and to have infinite thickness and width. In the case of a rectangular coil of length 2L and width 2W suspended at a distance Z0 above the ground plane, the image force Ft is

312

given by:

= Mnif l 2 Z - [L 2 +Z 2 ]*- [W 2 + Z2]*_ rcZ I

_ [W2 + Z2]*-[L2+W2 + Z2;]* z 2 _

w2+z2

[L2+Z2]^-[L2+W2 + Z2]^ 2\

fTT^ z { (1)

In this expression, n is the number of turns in the magnet and i the current in each turn. It is readily seen that the system is stable since the force increases as the magnet approaches the ground plane. The forces actually vary somewhat faster with suspension height if the magnet is a superconducting magnet operating in the persistent mode. The magnet current then becomes a function of suspension height, varying as i= i0 (l+k) where i0 is the current in the magnet when completely removed from the conducting surface and k is the ratio of the mutual to the self-inductance of the real and image magnets.

If the guideway has a finite resistivity, as will be the casein any practical application, this formulation provides the absolute maximum lift force attainable only at infinite velocities. At zero velocity, the ground plane experiences a constant magnetic field and no currents are induced, therefore there is no lift force. Approximations must be introduced into the calculations to handle finite velocities or guideway s of finite dimen­sions. The case of finite dimensions has not yet been attempted analytically; however, the velocity dependence of the lift and drag forces at high velocities has been cal­culated. The effect of finite guideway resistivities is to introduce a velocity dependence into the lift force and to produce a magnetic drag force arising from the dissipation of energy in the normally conducting guideway. From the differen­tial equations in this limit, it is found that the important parameter is £ = (4n/\i<jvz)±, where v is the velocity. To obtain an approximate expression for all velocities, it is recognized that the solution obtained is the first term of the exponential expansion. If the exponential is taken as the proper solution, the correct dependence at high veloc­ities and the proper limit at low velocities are obtained. The resulting expressions for the lift (FL) and drag (Fd) forces are:

FL/F, = exp { - CD + 2 C Z / 2 L ) - ] - 1 } (2)

F , /F L = i [1 - exp { - C[ l + 3 ( Z / 2 L ) * ; r } ] (3)

The magnitude of the dimensionless parameter £ determines the meaning of high velocity; the velocity is high when £ is less than one. These formulas are compared with experimental data obtained in tests of a 1-J x 4^" electromagnet containing 38 turns and positioned over a rotating aluminum plate as shown in figure 1. We do not mean to imply that these solutions are exact, but rather that they are semiphenomenological expressions with the proper limits that approximate the experimental results. An important consequence of both the analytic results and other data [3], which space does not permit us to reproduce, is that the magnetic drag force at high velocities decreases as v ~*. Thus, the system performance improves as the velocity increases. The magnetic lift force approaches an asymptotic value equal to the image lift force of equation 1.

Just as a magnetic lift force is obtained by moving a magnet over a horizontal ground plane, a similar repulsive force will be exerted on any other metallic conductor located near the path of the magnet. This feature is easily used for guidance by bending

313

the outer portions of the metallic track into vertical positions to obtain lateral forces. If the vehicle travels in a path equidistant between the two vertical conductors, no net force will be experienced. If the vehicle moves toward one of the vertical conductors, however, the repulsive force on that side will increase and a corresponding decrease will

I .5

o I ',0

< en o

Q

< tZ 0.5

0 0 100 200 300

VELOCITY fps

Fig. 1 — Magnetic lift and drag forces for small test magnet at suspension height of 0.375 inches and current of 14.14 amperes.

be experienced by the opposite side. Both of these effects tend to return the vehicle to its equilibrium position. This gives the system both vertical and horizontal stability. Using the expression above, the performance of a practical superconducting magnet suitable for a MAGLEV vehicle can be predicted.

TYPICAL MAGNET PERFORMANCE

The parameters to be used in designing the superconducting magnets required for a MAGLEV transportation vehicle are not well defined at present. These parameters depend on the vehicle's weight and physical requirements, as well as the usual cryo­genic and magnetic requirements. The size of the vehicle will obviously depend on the anticipated passenger demand, optimum scheduling, and other parameters. Conse­quently, we will assume a passenger compartment resembling a jet aircraft fuselage with a capacity of 100 passengers and a weight of 50,000 pounds. We will further assume that the load is borne by six superconducting magnets positioned three to a side and underneath the vehicle but protruding beyond the sides of the vehicle into the guid-eway.

Assume that these magnets are square, 0.5 m on a side, and produce 3 x 105

ampere-turns. Further, let the winding be contained in a circular cross section 7.6 cm in diameter. Such a magnet produces a maximum magnetic field of 16 kOe in the windings and has an overall current density of about 7x 103 amperes/cm2. The dissipation of power in the event that all of the current were to be conducted in the copper sheathing of the assumed Nb-Ti conductor would be 0.27 watts/cm2, well below the film boiling power density level. With the assumed dimensions, channels 0.45 mm wide would be provided between conductors to avoid accumulation of helium bub-

Z0

I

= 0.375"

= 14.14 amp.

EXPERIMENTAL r T ^ ^ ^ '

—

Ly

LIFT - \ r/y' ~ ^ ^ V c A L C U L A T E D LIFT

// ^EXPERIMENTAL DRAG ~ |

/^^CALCULATED DRAG

1 !

314

bles. A total of 85 pounds of conductor would be required per magnet. These magnets would be fully stabilized and could not become normally conducting without the loss of the liquid helium in which they would be immersed.

The calculated magnetic lift and drag forces for each of these magnets are shown in figure 2 as fractions of the image lift force, which is somewhat larger than 11,000 pounds. Eighty percent of the image lift force would be achieved at a velocity of 50 mph so conventional rubber tires can be used for suspension and guidance at low speeds

i.o SUSPENSION HEIGHT = 0.15m (5.9"

LIFT/IMAGE LIFT

-DRAG/IMAGE LIFT

50 100 I5C 200 250 300 350 400 SPEED — mph

Fig. 2 — Magnetic lift and drag forces for proposed MAGLEV magnet as a function of speed. Image lift force is 5.5 tons.

until lift-off occurs. A suspension height of 0.15 m has been assumed in these calcula­tions. The lift force triples before a dewar 7.5 cm in radius can touch the guideway as shown in figure 3. This increased suspension height over that attainable with practical air cushion vehicles is a significant advantage of magnetically suspended trains. The necessity of using superconducting magnets in this application is best illustrated by noting that an equivalent copper magnet operating at room temperature would require 1.3 x io 6 watts of power.

14

</> 12

10

u cr. o

hr i

i U

i [^-VEHICLE

^ | - j - TOUCHES j GUIDEWAY

2K I

4 5 6 7 8 <

SUSPENSION HEIGHT inches

10

Fig. 3 — Magnetic lift force as a function of suspension height for proposed MAGLEV magnet. Equilibrium suspension height is 5.9 inches.

315

The energy consumed by excursions through minor hysteresis loops as the train traverses irregularities in the track will cause a small decrease in the magnet current. The recent introduction of twisted and stranded cables will significantly reduce these losses.

Operating at a speed of 300 mph, the magnetic drag would be approximately 5 percent of the total vehicle weight or 2,500 pounds, requiring approximately 2,000 hp. Including aerodynamic drag and additional power for .05 gs acceleration at this speed, the overall power requirements are about 5,500 hp.

ADVANTAGES OF MAGNETIC SUSPENSION

The rather high suspension height possible in a MAGLEV system compared with an air cushion vehicle has several advantages. One of these is the reduced probability of contact with the guideway or debris on the guideway. The magnet described previously provides a suspension height of 0.15 m. Although it is modest in size, it could be made still smaller if a lower suspension height were desirable. As the sus­pension height is decreased, however, the tolerances of the guideway become more stringent. The magnet presented here can traverse a W vertical offset in the guideway while experiencing an acceleration of only 0.1 gs. Larger offsets could easily be toler­ated if an auxiliary shock absorbing system were used to isolate the passenger com­partment from the suspension. These characteristics make the structural design of the guideway considerably more flexible.

If it should be desirable to construct a hybrid system composed of two or more suspension systems, MAGLEV should be completely compatible. Since this system requires no power source to provide levitation (the power dissipated in the track comes from the propulsion system), it continues to levitate in the event of power failure. The system is compatible with almost any means of propulsion whether it is jet powered, pneumatically propelled, or powered by a linear induction motor. Since there are no moving parts, little or no maintenance should be required. Finally, the passive nature of the levitation system assures its absolute silence, a major consideration when operating in populated areas.

ACKNOWLEDGMENTS

The experimental measurements reported here were made at the Sandia Corporation in cooperation with Messrs. Thomas Downey and Donald Williams. Their cooperation and the financial support of the Sandia Corporation for parts of this work are grate­fully acknowledged.

REFERENCES [1] J.R. POWELL and G.R. DANBY, "High Speed Transport by Magnetically Suspended

Trains," ASME Winter Annual Meeting, New York, Railroad Div., 66-WA/RR-5 (Nov. 1966).

[2] C.A. GUDERJAHN, S.L. WIPF, H.j. FINK, R.W. BOOM, K.R. MACKENZIE, D. WILLIAMS and T. DOWNEY, "Magnetic Suspension and Guidance for High Speed Rockets by Superconducting Magnets," presented at the Applied Superconductivity Conference and Exhibition, Gatlinburg, Tenn. (Nov. 1968).

[3] T.W. BARBEE, G.N. BYCROFT, E.G. CHILTON, F.M. CHILTON and H.T. COFFEY, "The Hypervelocity Rocket Sled—A Design Analysis," performed for Sandia Corporation under Contract AT(04-3)-115. See also F.M. CHILTON, H.T. COFFEY and T.W. BARBEE, "A Magnetic Support System," presented at the Applied Superconductivity Conference and Exhibition, Gatlinburg, Tenn. (Nov. 1968).

[4] D. SHOENBERG, Superconductivity, Plate I, p. 21, Cambridge University Press, London (1960).

316

DISCUSSION

P.F. CHESTER (U.K.) — With such high ratios of lift-force to drag-force what is the origin of the damping necessary to give good lateral and vertical dynamic stability at high speed ?

H. T. COFFEY — I should have been more explicit in my previous remarks in that I was referring to the static rather than the dynamic stability. Nevertheless, from experiments with small magnets placed over or between rotating copper discs we know that a considerable amount of damping is present in the system. We have not made any calculation of the damping.

P. H. MELVILLE (U. K.) — Have you considered the possibility of using magnet expulsion as the driving force. This can be done by making the tunnel get wider as you travel along it ?

H.T. COFFEY — N o .

317

AN INDUSTRIAL GAS REFRIGERATING MACHINE FOR THE TEMPERATURE RANGE FROM ROOM TEMPERATURE

DOWN TO 20 °K AND LOWER

HJ. VERBEEK Cryogenic Development Department, N. V. Philips Gloeilampenfabrieken

Eindhoven {The Netherlands)

Une machine frigorifique industrielle a gaz pour le domaine de temperatures allant de la temperature ambiante a une temperature inferieure ou egale a 20 °K.

RESUME : On donne une explication sur une machine frigorifique a gaz a deux etages derivee d'une machine frigorifique industrielle a gaz a un etage existant deja. On etudie la conception, le fonctionnement et les proprietes de cette machine. On traite plus en detail de la conception des echangeurs de chaleur. On explique le fonctionnement d'un systeme frigorifique dans lequel la machine frigorifique a gaz est incorporee dans un circuit a helium Joule- Thomson et avec lequel on atteint le domaine de temperature de Vhelium liquide.

On decrit une methode de transfert du froid par des tuyaux sur une longue distance. On etudie brievement le refroidissement en grande masse, depuis la temperature ambiante

jusqu'a la temperature de Vhelium liquide. Finalement on evalue le cout du refroidissement en fonction de la temperature a laquelle le

froid est fourni et en fonction de la distance de transport.

INTRODUCTION

An ever-growing number of processes requires refrigeration at decreasingly lower temperatures and in increasingly larger quantities. According as our knowledge of technologies, requiring refrigeration at low temperatures, increases, the related problems are transferred from the research stage to the stage of large-scale application. The development of the auxiliaries used for carrying these new technologies into practice should keep step with this progress and is in fact strongly stimulated by it.

In the research stage the need is mainly for small cold sources. Dependent on preference and circumstances, these may be cryogenic liquids or refrigerators. The use of test arrangements still offers considerable latitude for improvisation. In the develop­ment stage, however, the demands made on the required quantities of cold and the reliability governing the continuity of the tests usually become more stringent.

It will be clear that for large-scale utilisation of cryogenics in processes whose function is closely related with society, the degree of economic desirability and the reliability to be guaranteed are decisive factors.

The gas refrigerating machine described in this paper may satisfy the need for such an auxiliary in a number of processes. Herewith we should like to contribute to the exchange of data concerning the requirements for the applications and the availability of cryogenic equipment.

DESCRIPTION OF THE GAS REFRIGERATING MACHINE

The design of the two-stage gas refrigerating machine, type C20, now being developed, is based on an existing machine, type C, previously described by Dros [1]. The modifications introduced into the C machine to arrive at the C20 machine are of a secondary nature as regards the mechanical part. However, the design of the thermodynamically operating part, in which the cooling cycle is performed, differs

321

considerably, see figure 1. In contrast to the two-stage cooling process described by Prast [2] in which one piston and one stepped displacer are used, this refrigerating machine operates with one compression and one stepped expansion piston per working space. Each of the four main cylinders or working spaces is limited on either side by one compression and one expansion piston. The working space is formed by one compression and two expansion cylinders, one cooler, two regenerators and two so-termed freezers. Thanks to the stepped shape of the expansion piston it performs the function of two expansion pistons.

i

4)©®©

m, 5 \ 1X1M \

Q COMPRESSION PISTON

( 7 ) COOLER

( 7 ) REGCNATOR-INTERMEDIATE STAGE

(7) REGENATOR-LOWER STAGE

( 7 ) FREEZER -WTERMEDIATE STAOf

( 7 ) FREEZER-LOWER »TAGE

0' EXPANSION PISTON

Fig. 1 — C20 — Working principle.

The functioning of the two-stage cooling process can be explained in stylised form with the aid of figure 1. The cycle is performed in four phases between the situations 1, 2, 3, and 4 in a working space filled with helium gas.

Phase I from 1 to 2: The compression piston moves to the right. The pressure of the gas in the entire working space increases.

Phase II from 2 to 3: The compression and expansion pistons simultaneously move to the right. The gas flows through the cooler, regenerators and freezers from the compression cylinder to the two expansion cylinders. The pressure decreases.

Phase III from 3 to 4: The expansion piston moves to the right, the pressure decreases even more.

Phase IV from 4 to 1: The compression and expansion pistons simultaneously move to the left. The pressure in the working space increases.

In reality these four phases are not separated but they merge into each other. The cold is generated in two expansion cylinders and carried off to the outside in the two freezers. The compression heat is dissipated in the cooler. The most important function of the regenerators always situated between two temperature levels was already described by Kohler and Prast [3], [2]. It is of paramount importance that the volume variation of the expansion cylinders lead by approximately 90° with respect to that of the compression space. This phase difference is ensured by the shape of the drive gear.

The presence of two expansion cylinders and two freezers makes it possible to generate cold at two different temperature levels and to transfer it to a cold-transfer medium, for instance at 70 °K and 20 °K.

322

Helium gas flows through the pipes of the freezers and on the outside of these pipes the cold is transferred to the cold-transfer medium. The freezers have been designed for a pressure of 30 atm of the cold-transfer medium and are enclosed in a high-vacuum-tight insulating jacket. Their construction is such that cold can be transferred both through condensation, for instance of nitrogen, neon and hydrogen, and through convective heat transfer to—preferably—helium gas at high pressure. They are fitted with flanges to which the supply and drain pipes of the cold-transfer medium are connected. These flanges are located in the insulating space and ensure hermetical sealing in high vacuum.

The machine, which has a V-drive, is driven by an electric motor directly flanged onto the crankcase, the rotor of the electric motor being mounted on the crankshaft. By means of a crank-connecting rod mechanism two plunger rods, each fitted with two plungers, are moved to and for in hydraulic cylinders. In this way eight oil columns are driven which in turn move eight pistons correspondingly, see figure 2.

As was already stated, the C20 machine was derived from the C machine. There is little difference between the appearances of the two machines. That is why we show in figure 3 the C machine in a factory where it is used for the liquefaction of oxygen and nitrogen.

SPECIFIC DATA OF THE MACHINE

Electric power consumption Pel « 200 kW Refrigerating capacity of the intermediate stage (70 °K) PM « 5000 W Refrigerating capacity of the lower stage (20 °K) PE « 1750 W

The machine has been designed for heavy-duty operation. Its speed is 730 r.p.m. It contains approximately 1 Nm3 helium gas as working agent. The maximum pressure of this working agent is 60 atm. The freezers are suitable for a maximum pressure of the cold-transfer medium of 30 atm.

323

The machine uses approximately 20 m3 of cooling water per hour. Cooling by means of a closed-cooling-water circuit, cooled in turn by a freon refrigerating machine, should be preferred, however. The freon machine can be entirely air-cooled.

Fig. 3.

Figure 4 shows the refrigerating capacity PE as a function of the lower-stage temperature TE, the temperature of the intermediate stage cold exchanger TM being shown as parameter. The fact that this refrigerating machine can attain a temperature of as low as 12°K does not mean that it is justifiable to use it for refrigeration at this temperature, because the cost of refrigeration would be very high at this temper­ature. To attain temperatures lower than 12°K it would be necessary to combine the C20 machine with a Joule-Thomson cooling circuit.

Table 1 SURVEY OF GAS REFRIGERATION MACHINES (PHILIPS)

Type

C A20 B20 C20 (dev)*

Number of stages

1 2 2 2

Refrigerating capacity

Watt

25 000 80

320 1 750

at°K

11 20 20 20

Lowest refr. temp.

°K

35 12 12 12

Shaft power

kW

182 10 40

182

Application field

industry lab. and ind. lab. and ind.

industry

*(dev): machine under development.

For comparison we give in table 1 a survey of the main properties of a few types of our gas refrigerating machines that may be of importance with respect to the subject of this conference.

324

REFRIGERATION FOR THE TEMPERATURE RANGE OF LIQUID HELIUM

Figure 5 schematically shows how the C20 machine can be combined with a Joule-Thomson cooling circuit. Helium gas flows through the freezers of the C20 machine at a pressure of approximately 20 atm and cools down to approximately 16°K in

4000

3000

2000

1000

WATT

I

' 12 U 16 11 2 b 3

y r M- 7 0 « K / PM « f * 0 0 W

^ Tt 10 A >0 «K

Fig. 4 — C20 — Lower stage refrigerating capacity PE as a function of lower stage freezer temperature TE.

the last freezer. The helium then flows through a counter flow heat exchanger, thereby cooling down to approximately 6°K. After that the gas expands through a throttle valve to a pressure of approximately 1 atm. The expanded mixture of helium gas and liquid flows to a heat exchanger, where the liquid evaporates entirely and where the cold is transferred to an object or to a medium to be refrigerated. The helium gas then flows through a number of counter flow heat exchangers and reaches room temperature, after which a compressor compresses the gas from 1 to 20 atm. This compressor pumps approximately 70 grams of helium per second through the Joule-Thomson circuit at a power consumption of approximately 200 kW. The refrigerating capacity of this installation is approximately 600 W at 4.2 °K.

325

This refrigeration system can be improved by using an ejector-loop at low temper­ature [4].

m ffc 70 gr H « / i l i t m .

( 7 ) COMPRESSOR

(T) COOLER

( T ) FREEZER UNIT C20

( 7 ) COLO RECOVERY COLUMN

© HEAT EXCHAN9CR J.T. CIRCUIT

( 7 ) THROTTLE VALVE

I st INTERMEDIATE STAGE 130* K

2nd INTERMEDIATE STAGE 70 *K

1 §t LOWER STAGE 35»K

• 2 nd LOWER STAGE 15»K

Fig. 5 — C20 — Combined with J.T. refrigerating circuit-flow diagram.

C O L D TRANSFER SYSTEMS

For the transfer of cold from the C20 machine to an object to be refrigerated use can be made of an evaporation-condensation system operating on, for instance, nitrogen and hydrogen. In this paper we shall refrain from giving a description of this type of cold transfer system. Helium gas at high pressure, for instance 25 atm, offers more scope as a cold transfer medium. The gas then flows in a closed circuit through the freezers of the refrigerating machine to the object to be refrigerated. Subsequently it returns to the machine propelled by a centrifugal pump operating at a low temperature. The cold transfer line, figure 6, consists of a high-vacuum-tight insulating jacket containing four pipes for the cold-transfer medium. Two pipes contain the medium transferring cold at the lowest temperature level, for instance at 20 °K, while in the other two pipes the cold is transferred at the inter­mediate level, for instance at 70 °K. To avoid losses due to radiation in the lowest temperature stage a radiation screen has been connected to one of the pipes transferring cold at the intermediate-stage level. This screen completely encloses the two coldest pipes. It is thus possible to limit the cold losses to less than 0.5 W per running metre of transfer distance at the temperature level of the lower stage. Thanks to this arrange­ment cold can be transported at 20 °K over a long distance. It is advisable to use the cold of the intermediate-stage level for cooling a radiation screen enclosing an object to be refrigerated.

326

The principle described above can also be used for cold transfer at the temperature level of liquid helium over a long distance. The pipes at the lowest temperature level are enclosed by two radiation screens, see figure 7. The cold losses per running metre of transfer distance at the various temperature levels are of the order of magnitude of:

outer radiation screen (70 .. inner radiation screen (20 .. cold transfer pipe ( 4 ..

100°K) high vacuum + superinsulation < 2 W/m 30° K) high vacuum < 0.5 W/m

5°K) high vacuum < 0.1 W/m

SUPERINSULATION VACUUM TIGHT OUTER WALL 300*K

(LOWER STAGE) t g 20-30*K

RADIATION SHIELD e g 90*K

Fig. 6 — Cross section of cold transfer line (15 to 20 °K level).

From this we can derive the distance that can be bridged for cold transfer. This distance is determined by the available refrigerating capacity and the required refriger­ating capacity at a certain temperature.

VAKUUM TIGHT OUTER WALL 300'K

SUPERINSJJLATJOJL

MAIN COLD TRANSFER LINES,

COLD TRANSFER LINES (INTERMEDIATE STAGE) • g. 70-100»K

(LOWER STAGE) e.g. 25* K e.g. 4 -5»K

Fig. 7 — Cross section of cold transfer line (4 to 5 °K level).

Cooling down a heavy object, for instance a superconductive object, is a time-taking job. At room temperature the C20 machine has a refrigerating capacity of approx. 10 kW at the lower stage. According as the refrigerating capacity decreases at lower temperatures, the specific heat of the mass to be refrigerated also decreases. The machine refrigerates a mass of 10 tons of copper at the rate of approximately 10°K per hour. If during the cooling down period the refrigerating capacity of the intermediate stage is also used, a considerable saving of time can be effected. In the latter case provisions have to be made for a change-over of the flow of the cold-transfer medium through the intermediate-stage and lower-stage freezers, viz.: in series in the case of cooling-down and in parallel in the case of refrigeration.

Naturally this would only be a useful proposition if the reduction of the cooling-down time is considered an important factor in the entire process. After the mass has been refrigerated to approximately 15°K with the aid of the C20 machine, a change­over can be effected to the Joule-Thomson cooling circuit which cools the mass further down to the required temperature of 4-5 °K.

327

THE COST OF REFRIGERATION AS A FUNCTION OF TEMPERATURE AND COLD-TRANSFER DISTANCE

Our calculation of the cost of refrigeration is based on the following points: depreciation in 10 years, interest 6% per annum, maintenance and servicing 4% per annum, everything related to the total capital investment in the cooling installation. The cost of electricity charged is U.S. $ 0.011/kWh. This calculation relates to con­tinuous operation.

For 77 °K we have based the calculation on the C machine, for the temperature range from 40°K to 15°K on the C20 machine and for 4.5 °K on a combination of the C20 machine with a Joule-Thomson circuit. The results of such a calculation are shown in table 2.

Table 2

EVALUATED COST OF REFRIGERATION AS A FUNCTION OF TEMPERATURE

Temperature °K Machine type

Cost of refrigeration Depreciation,

interest $/kWh refr. Running

cost of energy $/kWh refr. Total cost U.S. $/kWh refr.

77 C

0.08

0.10 0.18

40 C20

0.6

0.6 1.2

30 C20

0.8

0.8 1.6

Factor of cost increase for cold transfer over a certain distance Distance

10 metres 100 metres 500 metres

1 000 metres

1.01 1.1 1.5 2

1.01 1.1 1.7 2.5

1.01 1.1 1.7 2.5

20 C20

1.5

1.5 3

1.02 1.2 1.9 3

15 C20

4

4 8

1.02 1.2 2.6 9

4.5 C20 + J.T.

7.4

8.6 16

1.01 1.1 1.7 2.5

We shall not deal here with the basis of calculation of the cost increasing factor in the case of long-distance cold transfer. The cost of the cold-transfer pipe and the cold losses are of course included in this calculation. However, the values given at page 327 have not been optimized and are only meant as an indication of the order of magnitude.

As regards the figures of table 2 we should like to state that a "large-scale applica­tion of the relevant refrigerating machines" has also been of the factors governing the cost price. The cost price per kWh of refrigerating shown in table 2 tallies reason­ably with the cost of refrigeration published by Kurti [5] in 1967.

CONCLUDING REMARK

In this paper we have tried to restrain ourselves in the technical description of the refrigerating machine and to consider it from the viewpoint of the users who will regard this machine as a tool with which various processes can be realised.

For this reason the explanation of the typically cryogenic aspect has also been kept brief. The specific properties of the refrigerating machine for cooling down from room temperature and for refrigerating at low temperatures have been empha­sised. In consequence of this a few possibilities for cold transfer have been dealt with. This refrigerating machine will represent a compact and universal cold source whose range of application has not yet been developed to the full.

328

REFERENCES

[1] A. A. DROS, An industrial gas refrigerating machine with hydraulic piston drive. Philips Technical Review, 26 (1965), No 10.

[2] G. PRAST, A gas refrigerating machine for temperatures down to 20 °K and lower. Philips Technical Review, 26 (1965), No 1.

[3] J.W. L. KOHLER and C O . JONKERS, Fundamentals and construction of a gas refrigerating machine. Philips Technical Review, 16 (1954), Nos 3 and 4.

[4] J.A. RIETDIJK, The expansion-ejector, a new device for liquefaction and refrigeration at 4°K and lower. Annex 1966-5 Bull. I.I.R., pp. 241-249, Com. I, Boulder.

[5] N. KURTI, Low temperatures in the generation and transmission of electric power—1.1. R. Xllth Int. Cong. Refrign, Madrid (1967), I, pp. 1-13.

329

CLOSED CYCLE REFRIGERATION OF A SUPERCONDUCTING MAGNET

C.N. CARTER, K.G. LEWIS, B.J. MADDOCK and J. A. NOE Central Electricity Research Laboratories, Leatherhead {United Kingdom)

Refroidissement en cycle ferine d'un aimant supraconducteur

RESUME : Pour rendre efficaces le refroidissement et le fonctionnement des grandes machines supraconductrices, il faut se servir de refrigerateurs. La temperature de regime rfest alors pas forcement liee au point d'ebullition de Vhelium liquide. Les auteurs rendent compte de details sur leur experience du cycle frigorifique ferme de Vaimant supraconducteur de CERL, soit avec liquide soit avec gaz.

Le refrigerateur qui a ete concu pour diverses utilisations presente done une puissance substantiellement superieure a celle exigee pour cet aimant et peut faire face a des besoins de 150 W a 4,4 °K. Vaimant est une bobine solenoide (100 mm d'alesage, 400 mm de diametre exterieur) avec enroulement d'un conducteur multiple et produit un champ de plus de 6 T. / / pese 210 kg et a une capacite calorifique de 18 MJ environ (entre 300 et 4°K).

On donne des details sur le refroidissement et le fonctionnement en regime permanent ainsi que des mesures des courants stables et critiques et Von presente d'autres aspects du fonctionne­ment lorsque Vaimant est refroidi par un gaz a des temperatures de 4,5 a 9°K. On examine les avantages et les inconvenients du fonctionnement dans un gaz ou dans un liquide.

1. INTRODUCTION

Large superconducting machines will need some form of closed cycle refrigeration for cool down and probably also for steady operation. The operating temperature is then not necessarily tied to the boiling point of liquid helium. We present some results for the CERL superconducting magnet which were obtained to gain experience of closed cycle refrigeration and of the operation of a magnet at various temperatures in helium gas close to atmospheric pressure. Brief mention is made of the performance of the magnet in liquid helium: a more detailed account will be given elsewhere.

2. MAGNET

The magnet [1] is a solenoid with a bore of 100 mm, an overall diameter of 400 mm and a length of 300 mm. It is wound with Niomax-M composite conductor [2] devel­oped jointly by the Central Electricity Research Laboratories and Imperial Metal Industries Ltd., which has a 10 mm x 2.5 mm section and is grooved transversely. The central field at the critical current is 6.4 T. There are 8 double pancakes stacked in two groups of four (fig. 1). With 140 kg of conductor and approximately 70 kg of stainless steel and plastic materials, the thermal capacity between 300 K and 4K of the magnet and lower support structure is 18 MJ. The seven 100 mm lap soldered joints between the pancakes together with the two connections to the 25 mm tubular gas cooled current leads have a total resistance of about 0.5 |iQ at operating temper­atures.

The coil is suspended with its axis vertical in a super-insulated cryostat. Vacuum-jacketed super-insulated transfer lines connect this to the refrigerator cold box: the sections at the magnet end can be seen in figure 1 running from 600 mm above the cryostat top plate through the radiation baffles to a little above the coil. The return line (at the back) has a fine mesh filter whilst the inlet line (at the front) has a liquid separator and a diverter valve which enables the helium flow to be admitted either

331

above or below the coil. It was expected that when the cryostat was full of liquid, this valve would be needed to avoid blowing the full refrigerator flow through the liquid. However, this proved unnecessary unless an undisturbed liquid surface was

Fig. 1 — Magnet and support structure (centre) with cryostat (left). A 500 litre liquid helium storage vessel can also be seen (right).

required. To force most of the gas through the magnet and particularly through the small passages formed by the grooves across the conductor, the coil is shrouded in styrofoam which fills the annular space between it and the cryostat wall (fig. 2). The bore is also blocked with styrofoam. Some gas is allowed to flow past the joints which are grouped on the outside of the coil.

Four methods of temperature measurement are used: (i) from room temperature to 20 K, copperxonstantan thermocouples, (ii) below 20 K and in positions where the magnetic field is low ( < 0.3 T), Cryocal germanium resistors, (iii) below 10K but in any field, Allen Bradley 1/10 watt 100 ohm nominal carbon resistors. The calibration of these is less reproducible than that of germanium resistors but the magnetoresistance is much lower. Within the range 4.2 K to 10 K, the maximum

332

apparent temperature error occurs at 4.2 K and is only 150 mK in a field of 7 T, (iv) in liquid for measuring the conductor temperature relative to the bath temperature, gold/iron: "silver normal" thermocouples.

COMPRESSOR FROM |

VACUUM INSULATED LINES)

MAGNET piVERTEPl VALVE

ICRYOSTAli

OLD! -COLD BOX

rtTl MAGNET m

'v-* > >

Fig. 2 — Refrigerator flow diagram.

For studying the stability of the coil and the recovery of resistive regions, all the pancakes have voltage taps at strategic positions. Also, three pancakes contain small heaters (20 mm long) placed close to the bore where the recovery current is lowest, each of which can be used to raise a length of the conductor above its transition temperature.

3. REFRIGERATOR

The refrigerator [3], which was designed for a variety of duties and therefore has a capacity substantially greater than is required for this magnet, can sustain a load of up to 150 W at 4.4 K. It was built by the British Oxygen Company and uses a modified Claude cycle (fig. 2). Helium at 8.3 atm pressure is pre-cooled by liquid nitrogen and approximately 75% of the total flow of 60 g/s passes through a gas bearing expansion turbine which drives a paddle wheel. Work is absorbed by a separate helium brake gas circuit in which this wheel rotates. The turbine runs at 4200 rps. The cold gas from the turbine outlet is used to cool the remaining 25% of the flow, which gives up further heat to the return flow in heat exchanger D before passing through a Joule-Thomson expansion valve VI10. This stream circulates through the magnet via 36 mm bore transfer lines each 7 m long.

During the first stage of the cool down the restriction of the small bore J-T valve and the high pressure side of heat exchangers B, C and D may be by-passed by opening valve VI15, thus allowing gas at liquid nitrogen temperature to be delivered directly to the magnet: the refrigerator heat exchangers are cooled by the return flow.

The diverter valve VI09 allows any fraction of the returning gas to by-pass exchanger D. Whenever the temperature of this gas is higher than that of the turbine

333

exhaust, full by-pass is used. At lower temperatures this valve is used to regulate the refrigeration produced. It may be automatically controlled so that either a constant outlet temperature or a constant liquid level in the cryostat is maintained.

In the same way the diverter valve VI07 is used to stabilise the turbine inlet temperature automatically, usually to 17K, by "dumping" a small fraction of the cold exhaust at the junction of exchangers A and B.

The rate of cool down to around 50 K of the complete system is limited by the total mass of heat exchangers A and B (530 kg) rather than by that of the magnet (210 kg). Exchangers C and D total 60 kg.

4. COOL DOWN

The main compressor circulates helium gas at or a little above room temperature through the cold box, transfer lines and cryostat. Any contamination present, particu­larly air or water vapour, is removed by purging and by the purifier. When the required purity is attained (< 500 ppm of all impurities and < 100 ppm of water vapour) the cool down, which has three main stages, is started.

S INLET TEMPERATURE S OUTLET TEMPERATURE GNET TEMPERATURE

TIME. HOURS

Fig. 3 — Temperatures and flow rate during cool down.

4.1 300 K to 80 K. Liquid nitrogen is supplied to the cold box boiler (fig. 2) and valve VI15 is part opened. The permissible flow is limited by the pressure which can be tolerated in the cryostat. This is 1100 mm at present because of a safety valve set to this pressure. As the temperature falls, the mass flow is increased by further opening VI15 in steps until the return helium gas reaches about 80K when the flow is nearly 30 g/s (fig. 3). During this stage the magnet temperature is considerably higher than that of the return gas, probably because not all the gas passes through the magnet winding. Some passes over the joints and some round the styrofoam which shrinks noticeably on cooling. The curve |of magnet temperature in figure 3 is for a pancake at the middle of the coil; the lower pancakes cool down somewhat sooner, the higher

334

pancakes later. The time for the whole magnet to reach 80 K is about 2 hours. The reduction in rate of fall of temperature towards the end of this stage is simply due to the diminishing temperature difference between gas and coil. The rate of refriger­ation in the cryostat is, for example, about 5 kW when the gas return (outlet) temper­ature is 150K.

4.2 80 K to 9K. Provided the oxygen content (which is taken as an indication of air content) is now less than 10 ppm, the turbine is started and the by-pass VI15 closed. This reduces the gas flow through the magnet to about 12 g/s. The temperature falls fairly slowly at first but then more rapidly because of the large fall in specific heat of the magnet and lower heat exchangers. During this period the flow rises because of the increasing gas density until the maximum flow of the compressor is reached. As the turbine inlet temperature continues to fall, the turbine requires more gas and so the flow through the cryostat has to be reduced until the turbine reaches its normal operating temperature, after which the flow stays nearly constant. It takes about 2 hours for this stage and the rate of refrigeration in the cryostat is, for example, about 300 W when the gas return temperature is 50 K.

4.3 9 K to 4.3 K. The diverter valve VI09 is now set for full flow through exchanger D. At these low temperatures the volume specific heat of the helium gas is greater than that of the coil. The rate of fall of temperature is therefore limited because the refrigeration power is used to produce the dense cold gas which is accumulated in the cryostat. It takes about \\ hours from 9K until some liquid is produced at 4.3 K (cryostat pressure 835 mm).

4.4 During the early stages of the cool down the bottom of the magnet is consider­ably colder than the top, but below about 10 K its temperature is uniform to within ±100 mK. Throughout the cool down a flow of about 0.15 g/s of gas is taken up the current leads and 0.06 g/s frornjthe top of the cryostat. This gas is returned to the circuit at room temperature.

5. OPERATION OF MAGNET

There are two important currents which describe the operation of a superconducting magnet. First, the critical current, which is the least current for which some part of the winding reaches a critical combination of current, magnetic field and temperature and above which the superconductor ceases to have zero resistance. Second, the recovery current which we define as the maximum current for which a substantial resistive region will disappear when the source of heat which produced it is switched off. For this magnet the recovery current is lowest close to the bore where the heaters are also placed.

5.1 Operation in Liquid. The recovery current as just defined is 1650 A which corresponds to an average heat flux of 0.35 W/cm2 from the exposed surface of the conductor. It is likely that the true value of heat flux direct to the liquid is lower than this with some of the heat, perhaps a quarter, being lost to adjacent superconducting material through the interturn insulation. Several turns are driven normal in this test and film boiling occurs.

It is possible to exceed the recovery current, without any restriction on the rate of change of field (0.1 T/s is the maximum with the present power supply), and reach the critical current of 2180 A. This corresponds to a central field of 6.4 T and a stored energy of 190 kJ. Further increase of current forces some of it to flow in the copper with the superconductor becoming slightly resistive in the high field region. The behaviour is reversible up to 2340 A, with a corresponding resistive voltage of 80 mV,

335

at which the limit of nucleate boiling is reached and the coil has to be discharged through an external resistor.

The current leads are cooled by a total flow of 0.2 g/s of gas. The liquid level in the cryostat is sensed by a carbon resistor and a signal derived from it controls the refrigerator diverter valve VI09 and hence the rate of liquefaction so as to maintain the level.

5.2 Operation in Gas. The recovery current, which was determined in a way similar to that already described, is shown as a function of the magnet operating temperature in figure 4. The gas pressure was 830 to 840 mm. On the basis of a constant heat transfer coefficient of the order of 10" 3 W cm" 2 K~ \ the recovery current line was expected to merge with the critical current line close to the critical temperature as observed, but at lower temperatures less variation with temperature and values somewhat lower than found were expected. Clearly, heat conduction to adjacent turns complicates the simple stabilisation calculations even more so than for the magnet in liquid. Expressed in terms of the area of conductor directly exposed to the gas, the apparent heat flux at 5K is 0.07 W/cm2. Contrary to the behaviour in liquid, resistive regions created by a heater at currents greater than a little above the recovery current propagated readily and spread to the pancakes immediately above. This latter effect was caused by rising hot gas rather than by conduction of heat along the conductor.

2000

1000

0

_ _ j

A N s

J

1 1 1 1

O — O MEASURED CRITICAL CURRENT J V — • _ « ■ • — P R P n i P T P H r O I T I T A I r i l D D C U T 1

^ ^ C L D — D RECOVERY CURRENT IN GAS > ^ v A RECOVERY CURRENT IN LIQUID |

N ^ V —

^ ^ \ ^ n > ^ \ 1

—1 1 J Ts \ l 1 7 « 8

TEMPERATURE, °K 10

Fig. 4 — Dependence of critical and recovery currents on temperature of magnet.

A typical size for a stable resistive region formed just above the recovery current is six turns of the winding. Growth inwards is limited at the bore by the much increased area of the conductor surface exposed to the gas (there is no mandrel or coil former). Growth radially outwards is checked by the falling field which reduces the magneto-resistance of the copper and raises the current capacity of the superconductor.

336

The recovery currents given in figure 4 are virtually independent of the gas flow rate within the range available. For example, at 5K a change from 11 to 18g/s increased the recovery current by only ^%. Such a small effect again suggests that heat transfer to adjacent turns dominates that directly to the gas.

Practically no restriction on the rate of energising the coil has been found provided the current is less than the recovery value. Above this, great difficulty is experienced and even at a rate of only 1 mT/s the behaviour is unpredictable. We conclude that electromagnetic and probably also mechanical disturbances within the coil are frequently severe enough to generate normal regions which cannot recover. When the magnet was energised (in gas) for the first time after it had been reassembled at room temperature, a very large disturbance occurred sufficient to produce a resistive region which did not disappear, although the current was 12% below the recovery current for the temperature of operation. This disturbance was almost certainly a sudden bedding down in the axial direction of the pancakes caused by the electromagnetic forces.

Because of the difficulty of obtaining currents above the recovery values, the critical current of the magnet as a function of temperature (fig. 4) was determined by setting the chosen current at a temperature where it lay below the recovery current line and then, with the current kept constant, by gradually increasing the temperature until the critical condition was reached. When the chosen current was greater than the recovery value in gas at any temperature, the coil was first energised in liquid. This was then siphoned out before the temperature was increased.

The third curve in figure 4 is the critical current predicted from the data of Hamp­shire et al [4] on the variation of critical current density for niobium—44 wt% titanium with field and temperature, together with some measurements of the short sample performance at 4.2 K of the magnet conductor. Agreement with our observed values is reasonable. The discrepancy seen at the higher temperatures is largely due to a measured difference of about 0.2 K between the zero field zero current transition temperatures of the magnet conductor and the material used by Hampshire et al.

At the lower temperatures, rapid propagation of a resistive region occurs immedi­ately the critical condition is reached, whereas above about 8.5 K the transition is less abrupt. Stable sharing of the current between the copper and the superconductor can then be detected but only within a temperature increase of less than 20 mK.

A measure of the dissipation caused by changing the magnet current and hence field was obtained at 5.4K by oscillating the current between 300 A and 700 A at 12A/s until the gas inlet and outlet temperatures were steady. The temperature difference between them was then 90 mK. When the current was fixed at 500 A this difference was 40 mK. Since the gas flow rate was 14 g/s, the extra dissipation was about 4 W, while that due to the cryostat, current leads and magnet joints was about 3W.

Throughout all these experiments with the magnet operating in gas, sufficient flow was taken through the current leads to keep the temperature at their lower ends within 150 mK of that of the magnet. A typical flow was 0.2 g/s.

6. DISCUSSION

Closed cycle refrigeration of this superconducting magnet has proved perfectly satisfactory both for cool down and for steady conditions. Operation in liquid with automatic control of level and in gas with automatic regulation of gas temperature have been achieved.

Reliable operation at currents greater than the recovery current proved almost impossible with gas cooling, in marked contrast to liquid cooling. This difficulty can

337

probably be overcome without having to have more copper stabiliser by using con­ductors with finer filaments of superconductor. This technique promises to eliminate the electromagnetic disturbances which lead to resistive regions, while mechanical disturbances can be prevented by thorough bonding together of the turns. The other anticipated problems all proved of little consequence. No large temperature disturb­ances were created by the steady dissipation which occurs while energising the magnet. The lower ends of the currents leads could easily be held close to the temperature of the magnet. No difficulty was experienced in obtaining a uniform temperature through­out the winding. However, there appear to be no particular advantages to operating in gas at pressures close to atmospheric, rather than in liquid, except that the potential rise in pressure in an emergency is much less and that there is freedom in the choice of temperature. To exploit this feature and obtain a reduction in refrigerator power consumption, a superconductor with a high transition temperature is necessary. The obvious disadvantages are the low heat transfer coefficient, which makes cryostatic stabilisation difficult, and the larger temperature fluctuations which are liable to occur because of the absence of a latent heat. Gas at a higher pressure (supercritical) and with a greater flow rate would be better.

ACKNOWLEDGEMENTS

We are very grateful to M.C. Brunning, K.W. Cannon, F.R. Gillepsie, M.C.A. Hookey and A. A. Humphreys for their help in preparing and running this system. The work presented in this paper was carried out at the Central Electricity Research Laboratories and is published by permission of the Central Electricity Generating Board.

REFERENCES

[1] B.J. MADDOCK, C.N. CARTER and P.B. BARRATT, Second International Conference on Magnet Technology (Oxford), pp. 533-536 (1967).

[2] M.T. TAYLOR, ibid, pp. 229-232 (1967). [3] A. P. STOLL, Proceedings of the Second International Cryogenic Engineering Conference,

(Brighton) pp. 36-38 (1968). [4] R.G. HAMPSHIRE, J. SUTTON and M.T. TAYLOR, Commission I, London, Annex 1969-1 Bull. I.I.R., pp. 251-257.

REMARK

N. KURTI (U.K.) — Both Mr Maddock's and Dr. Verbeek's papers emphasized the fact that it will be both desirable and economical to cool large pieces of equipment by running the refrigerator attached to it at continuously descending temperature levels. In other words, the difficulties associated with cooling by means of liquefied gases outlined by Dr. Scurlock (pp. 195-201) will not be met in the ultimate large-scale appli­cations with which this conference is mainly concerned. However his findings are very useful for pilot-scale experiments where for convenience liquefied gases are used.

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SURVEY PAPER

MATERIALS

P.H. BURNIER Dept. Cryoelectrotechnique, ALSTHOM, Massy {France)

Materiaux

RESUME : Vexpose general considere les proprietes physiques aux basses temperatures des metaux purs, des materiaux structuraux et des materiaux non-metalliques, en particulier les polymeres : les proprietes des supraconducteurs sont exclues de V expose, du fait qu'elles font Vobjet d'autres contributions presentees au cours de la reunion.

En ce qui concerne les metaux purs, on etudiera principalement la resistivite electrique et la conductibilite thermique, avec quelques considerations sur le nombre de Lorenz. Les effets parasites dus au champ magnetique et aux dimensions sont discutes. On discute brievement des proprietes electriques et thermiques des contacts.

En ce qui concerne les metaux purs et les materiaux structuraux, Vexpose considere par-ticulierement les proprietes suivantes : chaleur specifique, contraction thermique, proprietes mecaniques (limite elastique, resistance a la rupture et module d'elasticite sous des contraintes de tension, deflexion, de cisaillement et de choc). Quelques indications sont donnees sur la soudure et le coefficient de frottement.

En ce qui concerne les materiaux non metalliques, les principales proprietes prises en con­sideration sont la chaleur specifique, la conduction thermique, les proprietes dielectriques (constante dielectrique, angle de pertes, tensions de claquage) des fluides cryogeniques, des poly­meres, des materiaux composites {materiaux stratifies et poreux) et de quelques materiaux inorganiques. On mentionne les problemes poses par Vassemblage de materiaux composites.

En conclusion, revolution generate des proprietes des materiaux aux basses temperatures est discutee, de maniere a mettre en lumiere la vaste gamme de valeurs que peut couvrir chaque propriete selon le type de materiau considere.

INTRODUCTION

The properties of materials used in electrotechnics are so numerous, and their behaviour at low temperature presents so many peculiarities, that it seems not possible in a survey to give anything else than a general impression on such a wide subject. Some guide-lines could be looked after in the general theories predicting this beha­viour, but unfortunately, the theories are in a widespread range of advancement: if some of them may be usefully followed, other are very limited or even completely lacking. That is the reason why it will be tried to start with discussing the specific heat of materials, before tackling the transport properties and finally reaching the mechanical and dielectric properties.

I — SPECIFIC HEAT

The specific heat of materials results from the storage of thermal energy under the form of translational or rotational vibrations of any part of matter able to enter into motion. These parts can be atoms (lattice specific heat) in any type of materials, they can also be free electrons in metals (electronic specific heat) or more or less large parts of molecules in polymers, or even elementary magnetic moments in ferromagnetic materials. At high temperatures, the Dulong and Petit rule (1819) states that one gram-atom of any material has the same heat capacity of 6 calories per degree Kelvin (or 25 joules per degree Kelvin). We must bear in mind that a part of the heat supplied to a substance may contribute not only to increase its

339

internal energy, but also to perform an external work if the volume of the substance is kept constant. This leads to consider two heat capacities: Cp, measured under a constant pressure, and Cv, measured under a constant volume. The difference Cp - Cv is equal to the amount of work done in expanding against the pressure of the system. The difference Cp — Cv is small at low temperatures because the expansion coefficient is negligible under these conditions.

It was only after the discovery of the application of the quantum theory that the evolution of the lattice specific heat at low temperatures could be explained. Since the atomic vibrations in a solid have frequencies of the order of 1013 Hz, and by equating the energies hv and kT, it may be shown that at a temperature of 150°K or so, the atomic vibrations cannot be fully excited. This leads to a progressive reduction in lattice specific heat CV L as the temperature is decreased, reaching CV L = 0 at 0°K.

In 1907, Einstein calculated the evolution of CVL with temperature. He considered the material as being made of a series of independent identical oscillators, vibrating at a frequency vE characteristic of the material considered:

CV L = 3R * 2 e X p ( x ) , J = 3 R E ( ^ (1)

with:* = 0E/T = hvE/kT,E being the Einstein fuction, R the gas constant (R = 1.987 cal or 8.31 joule/mole.°K), h and k the Planck and Boltzman constants). Debye improved this model by considering, not a single frequency for the oscillators, but a distribution of frequencies of a parabolic form up to a maximum frequency vD, reached when the wave length is of the order of the lattice spacing of the atoms (a few angstroms). This distribution is explained through a coupling between the Einsteinian oscillators:

- ^ = 3 R - D ( 1 ° [ e x p ( x ) - l ] [ l - e x p ( - x ) ] \ l

CvL = ^ jo [e Z r — - = 3 R D l ^ ' (2)

with x = 0D/T = h.vD/kT and D being the Debye function. The Einstein and Debye functions are tabelled in a good review-monograph: " Specific heats at low temperatures" by E.S.R. Gopal (Heywood, London, 1966). At very low temperatures, a good approximation of these functions is:

CVL(Einstein) = 2 4 . 9 4 2 - x 2 - e x p ( - x ) x > 16 (3)

CVL(Debye) = 1 ^ 6 6 = ^ x > ^ ( 4 )

x 3

with 12**R

59£

CVL is there expressed in Joule per mole and per degree Kelvin. Practically the Einstein function underestimates CVL at low temperatures, and the Debye model is generally used, with 0D tabelled for various materials in numerous papers (see Gopal's book and also "Properties of materials at low temperatures—part I" by R. J. Corruccini—ChemicalEngineering Progress, June 1957, vol. 53, no. 6, pp. 262-267).

340

The Debye model enables very precise calculations of lattice specific heat, provided that the variation of 0D with temperature is taken into account, since the main short coming of this theory is the evolution of the so-called "Debye-constant" with temp­erature. As a rule, 0D keeps its high temperature value down to a temperature equal to one fourth of this value, and afterwards goes through a minimum before reaching near 0°K a plateau of values 0DO which can be lower or higher than the high tempera­ture value. It should also be borne in mind that 0D depends on the way it is measured: the value got from specific heat measurements may slightly differ from those obtained through other experiments: (elastic behaviour, compressibility, melting point, thermal expansion, infra-red data, electrical resistivity, scattering of X rays, y rays and neutrons).

Other models more sophisticated than Einstein's and Debye's models, such as those calculated by Blackman or Born Von Karman are not well fitted for practical calculations of CV L.

The theoretical calculation of the lattice specific heat of alloys and compounds is more complicated. Near Debye temperature, CV L can be approximated quite well by linear combination of the constituent elements (Kopp-Joule rule). For compounds with only one type of bonding, the Debye function is found to work properly with a 0D value bearing no relation with the corresponding values of the constituents. It is necessary to multiply the Debye function by 3.R./2 for a n-atomic molecule in order to get the molar heat capacity.

For more complicated structures, the real specific heat curves can apparently be represented empirically by a sum of suitable functions (the foresaid Debye and Einstein functions, and the Tarasso function: Zh-Fiz. khim., 24, III, 1950 and 39, 2 077, 1965). This is particularly the case of polymers, rubbers and glasses.

In the case of a lack of experimental data, several formulas give some means of getting approximate values of specific heat. They are described in the foremen-tioned Corruccini's paper.

In metals, the contribution of free electrons to specific heat has to be considered* It is shown to depend in pure metals on the density of states at the Fermi surface dN/dE, on the molar volume V and on temperature T following the formula:

Ce = f 7 i 2 f c 2 V — T = y T (5) dE

= 1 .36 .10" 4 -V*-n jT

Ce being expressed in joule per mole and per degree Kelvin, V being expressed in cm3/mole and ne being the number of free electrons per atom. The main feature of this expression is that Ce varies linearly with temperature. Consequently, at room temperature, Ce is small (generally negligible) compared to lattice specific heat, when Ce may be quite appreciable at low temperatures. For temperatures lower than 0D/24, the total specific heat may be written:

C = CV L + C e = | 3 T 3 + y T

so that a plot of C/T against T2 should therefore be a straight line. The experimental values of y and 0 D O (deduced from equation (4): p-O3^ = 1943.66) are given by Gopal (forementioned book), by Corruccini and Gniewek (Specific heats and enthalpies of technical solids at low temperatures, NBS Monograph 21, Oct. 3, 1960). Precise values for copper, silver and gold are given by Furukawa, Saba and Reilly (NSRDS — NBS note no. 18, 1968).

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For alloys, the situation is much more complicated, and the experimental results can be explained by the change in the density of states, in the numbers of free electrons and perhaps, in the strength of interaction (Low Temperature specific heat of Ni and some F.C.C. Ni-based alloys, by Ehrat, Ehrlich and Rivier, / . Phys. Chem. Solids, Pergamon Press 1968, London, vol. 29, pp. 799-806).

In superconductors, the electronic specific heat takes a different form than that given in equation (5) for normal metals:

Ces = Y T c a e x p ( - ^ ) (6)

Tc being the transition temperature, and a and b being temperature independent parameters (Low temperature specific heat of BCC titanium-molybdenum alloys, by Sinha, J. Phys. Chem. Solids, Pergamon Press 1968, London, vol. 29, pp. 749-754). Under such conditions, Cnormal is greater than Csuperconducting at low temperatures, but the curves cross each other when temperature is increased, up to a sharp discon­tinuity given by the Rutger's relation:

d-C.- -£(•£) ' (7) 4 7 r \ STJ

where Hc is the critical magnetic field. In ferromagnetic or ferrimagnetic materials, the excitation of electron spins

gives rise to another component proportional to T 3 / 2 , so that the total specific heat becomes:

Cv = yT + PT3 + 5T3/2

y being of course zero for ferrimagnets, which are insulators. Some more complicated formulas have been given to take into account particular cases (Theory of anomalous specific heat of nickel and copper-nickel alloys at low temperatures, by Bennemann, Phys. Rev., vol. 167, no. 2, 10 March 1968, pp. 564-572).

For paramagnetic salts, the energy of the magnetic moment corresponds to a term of specific heat:

C M = (Nu 2 H2//cT2) • sech2 (uH//cT) (8)

which is very important at very low temperatures. It is used in the so-called paramagnetic-cooling for obtaining very low temperatures, far below 1 °K. It may finally be interesting to observe that, in finely divided powders, a surface effect introduces in the total specific heat formula a term proportional to T2 and to the value of the surface.

If we now want to compare the values obtained, the electrical engineer will be more interested in knowing the volumic specific heat Cvv» expressed in joule per cubic centimeter and per degree Kelvin, than in the values given by unit weight or by gram-atom. For pure metals, table 1 shows the values of the corresponding coeffi­cient y v for the electronic component and p v for the lattice component, together with the values calculated at 4.5 °K for both terms and for the total volumic specific heat.

It may be seen on this table that the electronic term of the volumic specific heat varies in a wide range of values (from 14 microjoules per cubic centimeter and per degree centigrade for bismuth up to 8,400 same units for a-manganese both at 4.5 °K —i.e. within a range of 1 to 600). A similar remark can be made for the lattice com­ponent (from 108 same units for chromium up to 10,970 for lead—i.e. within a range of 1 to 100). The total volumic specific heat varies from 432 units for beryllium to

342

11,760 units for lead (range of 1 to 27), but similar values may be obtained with quite different contributions of the electronic and lattice terms; for example high values are met with in indium (9,230 units, 520 coming from the electronic part, and 8,710 coming from the lattice part) and in a-manganese (8,622 units, 8,400 coming from the electronic part, and 222 coming from the lattice part).

Table 1 SPECIFIC HEAT OF METALS AT 4.5 °K

Metal Yv OD PV CVe C V L C V Units J/cm3 °K2 °K J/cm3 K4 J/cm3 J/cm3 J/cm3

4.5 °K 4.5 °K 4.5 °K

Pb In Mn Ni Bi Cu Cr Al Be

173 115.5

1 868 1 069

3.14 97

204 136 45.5

96 109 476 440 118 344.8 610 426

1 160

120.3 95.7

2.44 3.47

55.5 6.69 1.19 2.51 2.51

788 520

8 400 4 800

14.1 436 917 612 205

10 970 8 710

222 316

5 051 609 108 228 228

11 760 9 230 8 622 5 116 5 065 1 045 1 025

840 432

From the values (not very numerous) got from technical literature for organic materials, one can see that the scope of the specific heat curve versus temperature varies by a lesser amount than for metals, though the values at room temperature or at very low temperature respectively lie in the same orders of magnitude (table 2 and fig. 1).

Table 2 VOLUMIC SPECIFIC HEAT (joules/cm3. °K)

T 300 °K 80 °K 20 °K 4,5 °K

1. Cryogenic fluids 2. Metals Pb

Cu Al Fe Mn Be

3. Organic a) thermosetting epoxy resins

bakelite resins (phenol—formol) b) rubbers

natural rubber butadiene styrene rubber

c) plastics polyethylene polyethylene + propylene polyvinyl chloride polyvinylidene-chloride polytetrafluoroethylene polystyrene polymethylmetacrylate

— 1.47 3.46 2.44 3.51 3.56 3.58

1.31

2.12

1.78 1.79

2.16 1.62 1.37 1.33 2.28 1.43 1.72

LN2 = 1.82 1.30 1.82 0.967 1.21 1.59 0.165

0.090

0.524

0.627 0.575

0.52 0.51 0.51 0.486 0.685

LH2 = 0.68 0.603 0.069 0.024 0.035 0.067 0.0029

0.088

0.110 0.106

0.104 0.040 0.045 0.167

LHe = 0.56 0.0118 0.00105 0.00084 0.0034 0.0086 0.00043

0.0027 to 0.0046 0.0087

0.003

0.0078 0.0061 0.0038

343

Fig. 1

It has to be noted that specific heats do not vary very much with the chemical or physical state of the material. This is not at all the case with the transport properties which are very sensitive to the variations of these parameters.

II — TRANSPORT PROPERTIES

The transport properties of materials are essentially the thermal and electrical properties. Electrical conductivity is a characteristic of metals only, but thermal conductivity has to be considered for any kind of materials. Nevertheless their behaviour at low temperatures will be seen here at the same time, as it is affected by the same parasitic effects.

Thermal conductivity can be estimated by applying the theory of perfect gas to phonons:

K = i C - V p - L , (1)

with the following notations: C = specific heat; \ p = velocity of phonons (or sound); Lp = mean free path of phonons. In non-metallic materials, the mean free path is only limited by the interaction of phonons with the lattice, when in metals it depends on the interaction of phonons both with lattice and with conducting electrons. In dielectric crystals, Lp is roughly inversely proportional to temperature (Debye hypo­thesis), but Peierls, taking into account more complicated phenomena (Umklapp process) showed that:

KL = A(T/eD)"-exp(0D/ tT) (2)

for temperatures lower than 0D/1O. The constants n and b are not very different from unity, and A is a characteristic of the material considered.

Reese and Tucker (/. Chem. Phys., 43,105, 1965) gave a more complicated formula for polymeric materials:

K = *-L (-) f °° *4e*p(*) dx (3 6n2k \a)Jo exp(x)-l'(x2 + A/aXbl)kYp/kT

where (A/a) is an experimental parameter of dimension cm" 1 and Xbl is the tempe­rature independent scattering length comparable to the dimensions of spherulites in the polymer. This formula seems to fit the experimental data obtained with partially crystalline materials, such as nylon and polyethylene (Kolouch and Brown, J.A.P., vol. 39, no. 8, July 1968, pp. 3999-4003).

In practice, the behaviour of non-metallic materials shows that the phenomena involved are much more intricate than these theories can suggest it. High-purity mineral crystals present a sharp maximum in thermal conductivity at temperatures in the range of 20 °K to 50° K, reaching values of KL up to 150 W/cm. °K, and falling down to values of the order of 1 W/cm. °K both at liquid helium temperatures and around 100°K (fig. 2). Much lower values are met with in amorphous materials, such as glasses or polymers (fig. 3). These values lie in the range of 10~3 to 10~4

W/cm. °K at cryogenic temperatures. Three remarks have to be made:

— Amorphous materials, such as glass or quartz or polymethylmetacrylate or teflon, reach a plateau between 4 and 20 °K. On the contrary partially crystalline polymers, such as polyethylene or nylon, have a lattice thermal conductivity continuously decreasing when lowering the temperature. (Compendium on properties of materials at low temperature, by V. J. Johnson - NBS Pergamon Press, 1961).

345

For such polymers, thermal conductivity increases with increasing density (Kolouch-Brown. J.A.P., vol. 39, no. 8, July 1968, pp. 3999-4003): this pheno­menon is probably related to the greater crystallinity of the higher density grades.

When stretching these materials, heat conductivity increases along the direction of stretching, and decreases in the perpendicular direction, giving rise to a marked anisotropy (Fox and Imber, / . of applied Polymer Science, vol. 12, 1968, pp. 571-579).

346

In metallic materials, the lattice thermal conductivity becomes mostly oversha­dowed by the electronic term Ke, because the diffusion of phonons on conducting electrons is much larger than that on lattice atoms:

K = KL + K e # # K . (4)

01

JUUO

O06

005

O04

•^ flfl3

J ^ b 001

^ ^ 001

\ 0006

s \ 0004

K 000%

OOOZ

.0001

8 9 10

^/ / •

/

''/ A

It 14

/ / / /

/

/ /

16 18 20 30

/ / *— / _,

• /

y r / /

'

/ / / / / / / / / / /

f /

— "^^""^z^

*S^A/y/nn

A

Temperature 40 60 80 100

/ '

*r*

^*£P£ 0,956 s

/k7 ^ ^^TtFlon

^P£0,9U

y '^•'Perspex

^

120 160 200 300

1 ' 1 ^ Quartz glass ^ s ^

/ S</?/s7Se>

Silicc vye r ub^P^"

/Voturat

400 50C

-_

-

hN rubber

55 .5

4

.3 fc

t ./ «}

08 ^ \

06 ^

04 C: 0 0

03 N . 5 $

01 \

01

.008

006 6 8 C

Thermal conductivity of o/os&es, on at P/a&t*

2D 40 60 80 100 Temperature °R

200 300

Fig. 3

The electronic thermal resistivity 1/Ke is itself the sum of two terms: the ideal thermal resistivity 1/Kf, due to the scattering of phonons on electrons, and a residual thermal resistivity 1/Kr, due to impurity effects:

1/Ke = l/K, + 1/Kr

It has been shown that, at low temperatures the first term is proportional to T2, and the second inversely proportional to temperature:

1/Ke = ocT2 + p/T (6) Consequently, a plot of T/Ke versus T3 is a straight line with a slope. High

purity metals show a marked maximum of thermal conductivity at a temperature generally below 30 °K (all the lower as Debye temperature is lower ). The value of the maximum can reach 200 W/cm. °K, when the room temperature values lie around

347

1 W/cm.°K (fig. 4). Experimental values of thermal conductivities can be found in the NBS Compendium of properties of materials at low temperatures (1961) and in the NBS circular 556 (1954). A law quite similar to relation (3) is met with when consider­ing the electrical resistivity (Matthiessen's rule-1864):

P = Pi+Pr (7)

10 100 1000

Thermal conductivity oF metals and alloys

Fig. 4

The electrical resistivity is the sum of the ideal electrical resistivity, due to the interaction of electrons with lattice phonons, and of the residual electrical resistivity,

348

due to the scattering of electrons by impurities and lattice defects (dislocations, vacancies, interstitial atoms). But in this latter case, the residual resistivity is inde­pendent of temperature, and the ideal resistivity is a rather complex function of temperature, given by the Bloch-Gruneisen relation:

/•0r - B (T

JD/T v 5 x -dx o [exp(x) - 1 ] [1 - exp( - *)] (8)

B, M, 0D being respectively a constant, the mass of an atom and Debye temperature It can be shown that pt is proportional to T at high temperatures (T > 0D), and proportional to T5 at low temperatures (T < 0D/10).

A good survey giving experimental values of electrical resistivity for many metals has been made by Hall (NBS Technical note 365, 1968). The whole subject of "Elec­trical resistance of metals" is given in Meaden's book (Plenum Press, 1965).

It is interesting to see in Bloch Gruneisen formula that the ideal resistivity depends on temperature only through the ratio T/0D. Accordingly, a given value of resis­tivity is found at a high temperature if the Debye temperature of the material is high. This is particularly the case of beryllium, which has a Debye temperature three times higher than that of aluminium, and four times higher than that of copper. Beryllium is the only metal which is worth being used at liquid nitrogen temperature, every other metal having to be cooled down to liquid hydrogen level. Under such conditions, beryllium is a very attractive hyperconducting (or cryoresistive) cryoelec-trotechnic material. Its high price may be compensated for by the lower price and the better efficiency of liquid nitrogen refrigerators. Its interest is still more attractive when the thermal leaks of the apparatus considered are important, such as in trans­mission lines or in rotating machinery. We are currently developing a heat treatment which could give commercial grades of beryllium wires a very high resistivity ratio at liquid nitrogen temperature.

One can also find a relation between the mean free path of electrons Le and the electrical resistivity, which can be compared to relation (1) given for thermal conduc­tivity:

ne Le

m, n, e being the apparent mass, number per atom and electrical charge of electrons in the theory of "quasi-free " electrons. Ve is their velocity, of the order of 108 cm/sec. These four parameters are essentially temperature independent, so that the product p-Le should be a constant for a given metal.

The Wiedemann-Franz relation connects with each other the electrical resistivity and the thermal conductivity:

P . K e = L . T (10)

The factor L is called the Lorenz number.

L = in2 f-Y = 2.45.10-8(V/°K)2

but its experimental determination clearly shows an evolution of this constant at low temperatures (fig. 5) with a marked minimum at a low temperature, and sometimes a maximum at an intermediate temperature. The minimum value is lower when the metal is purer and better annealed. For copper with a resistivity ratio of 1:450, a

349

minimum of L = 0.85.10"8 (V/°K)2 is reached at 25 °K, and the plateau value expected at very low temperatures goes up back to a value of 2.2.10"8 (V/K) much closer to the theoretical value, which at high temperatures has not yet reached this theoretical value at 600 °K (Moore - Elroy-Graves, Canadian Journal of Physics, vol. 45, 1967, pp. 3 849-3 865, and Roder-Powell-Hall, Proceedings of the 5th Inter­national Conference on Low Temperature Physics and Chemistry, 1957, pp. 364-367). The Lorenz number may be increased by magnetic field (de Nobel, Physica XV, nos. 5-6, July 1949, pp. 532-540).

A2,5

10 100 Vor/ct£/on of cne. Lor en 3 number with T. for para annealed copper

Fig. 5

Using the Wiedemann-Franz relation, and supposing the Matthiessen's rule and Bloch Gruneisen formula are valid, and neglecting the variation of L with T in the temperature range considered, we have shown that the maximum of thermal conduc­tivity is found when the ideal electrical resistivity is equal to one fourth of the residual resistivity (fig. 6). This remark may give an easy means of choosing the metal which has its maximal thermal conductivity at a given value of temperature. As the value of this maximum is the higher when the material is the purer, i.e. when the residual resistivity is the lower, this generally results in the choice of a metal with a low Debye temperature.

350

Both thermal and electrical conductivities are indeed more or less markedly affected by the influence of magnetic fields. Mendelsohn and Rosenberg (Proc. Royal Society, A 218, 1953, pp 190-205) found that thermal magnetoresistance is very high in cadmium, appreciable in zinc, tin, lead and gallium, slight in some metals

W

flcm

/(max ^ p =. 4 Q

\lCurves given for pure copper)

fGmp&raCure

100 1000

flax/mum of thermal conducc/v/Cy

Fig. 6

such as copper and aluminium, and negligible in a good number of other materials. A field of 18.5 koe divides the thermal conductivity of cadmium by a factor of one thousand at 2.3 °K, that of tin, zinc and thallium by a factor of 10 to 80, depending upon the purity of the sample.

351

to Aer,« \B*

/ V

AL(20/4°K)

Cu tlAGMETORESISTANCE OF & * - Al.Cu.

IN A SCOHLEU'S PLOT

jBe*w/>e *0,5mm P3°° -/ft *4,7

AL = 80pm bands % 4.2

Cu = W/AC <fi 0,5mm 273 ^ W-272-855 #4,2

AL(4t2°K)

J 15 45 60 75

Fig. 7

Electrical resistivity of metal undergoes a similar magnetoresistive effect. Kohler found that the relative increase in resistivity ApH/p0 is a single function of the product of the field strength H by the resistivity ratio PeD/p0, whatever the temperature and the purity of the sample may be. Under a contract from INCRA, we have verified this rule in copper, and, in a cooperative work made with L'AIR LIQUIDE and PECHINEY, we have shown that this rule is verified by beryllium, but not by alumi­nium thin films (fig. 7). This magnetoresistive effect presents a saturation at high fields in metals such as aluminium, or continues to increase steadily without saturation in metals such as copper or beryllium. In ferromagnetic metals, such as iron or nickel, this magnetoresistive increase follows a marked decrease of resistivity, which takes place during the first few hundred oersteds of the curve (Dimitrov, Annales des Mines, April 1968, p. 75). This anomaly may be due either to the vanishing of Bloch Walls or to the orientation of magnetic domains.

10

I

<H 0.1

.Ol

I I I

/ I s\

'-\-\fi —IcW | ^y W§ w\

' I I I

4& Wfr^-Wvi? >n

I

jjJJ

*

LHfl LTHI

UJ

.Ol 0.1

Fig. 8

10

Another parasitic effect found both in thermal conductivity and in electrical resistivity is the size-effect, occuring when the mean free path of phonons or respec­tively electrons reach the same order of magnitude as the thinnest dimension of the sample. This size effect is practically found only at low temperatures on highly conductive materials (high-purity mineral-crystals or metals), when the mean free path can reach high values, of the order of a fraction of millimeter or even more. Sondheimer made a theoretical study of the electrical size effect (Adv. Phys., 1, vol. 1, 1962), which Corruccini (NBS Technical note no. 218) expressed under the form of

353

simple curves which can be approximated as (fig. 8) :

Apd = c Le (11)

Po d

C being a constant, equal to 0.4 for films and 0.8 for circular wires, d being the thickness of the film or the diameter of the wire. Andrews, West and Robeson (to be published in Phil. Mag.) showed that a similar effect is also found in a polycrystalline copper and aluminum: the residual resistivity markedly depends on grain size, being proportional to the inverse boundary intercept distance (fig. 9).

p 1

1

1 — I w i ret HSU

i

vtty

over

to' oh 1 77 C/l

i i

se i

~ r

>our idar

|

1 i

/ ".

A C

?/er

L-u. - J

cept

r

A.B

D

.C.

\tan CB r i 1

"C

D-

A\

r

f 5 9 10 11 It 13 14 15 16 1?

Fig. 9

354

VH "* \ Hoo

oe o

GALVANOMAGNETOMORPHIC EFFECT

1L

0,5±

x THICKNESS : D

T = 4,a°K

53,4- jj 34,4 p

f o o o : 1.x 10" A cm f ooo = 1,5x10" H e m

H.d oersted.cm 4 _ ^

Fig. 10

Size and magnetoresistive effects can be mixed in the so-called galvanomagneto-morphic effect. Mc Donald and Sarginson tried to establish a theory of this compli­cated effect. Together with L'AIR LIQUIDE and PECHINEY, we have shown that in aluminium films, this effect corresponds to an increase of size effect up to a maximum value, followed by a decrease which finally causes the vanishing of size effect at high magnetic fields (fig. 10) (unpublished paper presented at the Low Temperature Electrical Conductivity Symposium, London, May 1965).

14

10'

K)-2\

to'*i

W/cm3

For J-JOA/mm1

IV=/°7V

d*0,3cmy a

c/;0,£pJ7

/

V

fy< 10 -3 W K)'7 k?-{

AC ZOSS£S DU£ TO JOULE AND SKJN EFFECTS

Fig. 11

Every result presented above is related to steady state conditions (for instance DC conditions for electrical values). Under AC conditions, it should not be forgotten that other phenomena can become very important, such as skin-effect (fig. 11) and

356

10 .13

10 .12 -10 10" 10"" 10" 10" 10

RESISTIVITY (ohm-cm)

. 7 10_ 10"

Fig. 12

especially eddy current effect (fig. 12). In practice, when designing an apparatus using high purity metals, one has to find a compromise, generally between size effects and eddy current effect. This leads to a complicated technology using hundreds or thousands of very thin wires or flat bands, electrically insulated against each other and thoroughly transposed, for building a single high-intensity conductor.

We have not devoted a great attention to alloys, in which the mean free path, remains very small ; thermal conductivity and electrical resistivity of alloys accordingly decrease steadily with temperature, by a quite moderate amount, without any complicated parasitic effect.

We consider the subject of " thermal insulation systems " as being out of the scope of this survey, since it concerns systems and not purely materials. Nevertheless a very good book has been recently published under this title by Glaser, Black, Linds-trom, Ruccia and Wechsler by the NASA (NASA SP 5027, 1967).

Ill — MECHANICAL PROPERTIES

It is very important to study the mechanical properties at low temperature because some materials become very brittle under these conditions and also because in complex structures the thermal contraction may cause very high stresses to appear on some parts.

As a rule, the face centered cubic metals remain ductile up to the lowest tempera­tures, and body centered cubic metals and most organic materials become very brittle at temperatures which may be above the liquid nitrogen temperatures. The brittleness of b c c materials, which is met with both in polycrystalline substances and in monocrystals, is not yet clearly explained: it may result from the difficulties of dislocations movements inherent to this type of lattice or from the interaction of dislocations with interstitial atomic defects (solid solution hardening). (Fleischer, Ada Metallurgical vol. 15, Sept. 1967, pp. 1513-1519). The brittleness of most organic material (Teflon seems to be the only exception) is probably due to the great strength of secondary bindings at low temperatures, and to the interference of high internal stresses due to thermal contraction. This is the reason why a short review of the expansivity phenomena will be given before trying to get a general outline of the evolution of mechanical properties of materials at low temperatures.

As regards thermal contraction of materials at low temperatures, Gruneisen has shown that the coefficient of thermal expansion varies similarly to the specific heat:

Y P I C V _ I B £ C £

PT and ps being the isothermal and adiabatic compressibilities and y the Gruneisen constant. The volume V and compressibilities being approximately temperature-independent, a is essentially proportional to the specific heat, except for some varia­tions of y at the lowest temperatures. Consequently, the coefficient of thermal expan­sion vanishes at zero degree Kelvin. The value of the Gruneisen's constant is appro­ximately equal to two. Precise data are given by Corruccini (Properties of Materials, Chem. Engg Progress, June 1957, p. 264). Gruneisen formula does not fit very well the experimental data for amorphous or polymeric substances.

It can be seen from table 3 that the total contraction between room-temperature and 0°K varies for metals between 0.15% (titanium) up to 0.7% (lead and indium). It is much lower for pyrex glass (0.057%) and still lower for fused silica, and generally much larger for plastics (PTFE 2.1 to 2.46%, FEP 1.7%). The values obtained for epoxy resins vary widely between 0.4% (matching with aluminum or silver) up to 1.2%.

358

Table 3 THERMAL CONTRACTION BETWEEN 293 °K AND 0°K

Material [zf///]°93 in % Reference

Material [Al/l]°293 in % Reference

Epoxy resins [Al/l]°293 in % Reference

Plastics [zf///]°93 in % Reference

Cu 0.326 C1) Mg

0.503 (2)

Stycast 2850 GT

0.405 (3)

Teflon PTFE 2.1 to 2.46

(2)

Ag 0.413 C1) Ti

0.155 (2)

Stycast 2850 FT

0.425 (3)

Zn 0.683

o St. St. 304

0.304 (2)

Stanford 43

0.500 (3)

Teflon FEP 1.7

Al 0.415

o

Torr Seal

0.680 (3)

In 0.706 O

Monel 0.261 (2)

Epon 828

0.910 (3)

Pb 0.708

0) Incond 0.238 (2)

Resibond 907

1.035 (3)

Fe 0.198

o Brass 0.397 (2)

Araldite 501

1.107 (2)

Ni 0.224

0) Pyrex 0.057 (2)

Epon 815

1.22 ^3)

C1) Properties of materials at low temperature—A compendium, V.J. Johnson—Pergamon Press—1961. (2) Properties of materials at low temperatures—R.J. Corruccini, Chemical Engineering Progress, June 1957, Vol. 53, no. 6, pp. 262-267. (3) Thermal expansion of expoxies between 2 and 300 °K—Hamilton, Greene, Davidson, Review of Scientific Instruments, May 68, pp. 645-648. (4) The Journal of Teflon, Du Pont French edition, Jan.-Feb. 1968, p. 7. This paper also gives values for charged PTFE and FEP. (5) Numerous values for a number of alloys are given by Clark in Cryogenics, vol. 8, n° 5, Oct. 1968, pp. 282-289.

^

Table 4 compares the a coefficient of metals (between 0.05 for iron up to 0.25 for lead, indium and zinc at 100°K, and 0.11 up to approximately 0.3 for the same metals at 250 °K). With those measured on polystyrene, polyethylene, polypropylene and polyethylene-polypropylene mixtures (0.5 to 0.6 at 100°K, 0.6 to 1.5 at 250°K). The thermal contraction of plastics is definitely much higher than that of metals.

This difficulty may be overcome by filling these resins or plastics with low expansiv­ity powders, fibers or fabrics. Hamilton Greene and Davidson (R.S.I., May 1968, pp. 645-648) have shown that filling with fine quartz powder reduces the contraction proportionally to the percentage of charge. Campbell (Advances in Cryogenic Engi­neering, vol. 10, 1964, pp. 154-162) found in laminates that the contraction perpen­dicular to the layers may be estimated by adding the contractions of each component proportionally to its volume. In the direction of the fibers, the contraction is mainly controlled by the constituent having the higher young modulus, which can be the fiber in the case of glass fibers, or the resin in the case of nylon fibers.

Material 10 4 aa t 100 °K 1 0 4 a a t 2 5 0 ° K Reference

Material

1 0 4 a a t 100 °K 1 0 4 a a t 2 5 0 ° K Reference

Material

104ocat 100 °K 1 0 4 a a t 2 5 0 ° K Reference

O Same as (*) in (6) Zakin, Simha,

Cu 0.105 0.160

o Fe

0.050 0.111

o PE cast or machined

0.50 1.50 (6)

table 3.

Table 4 EXPANSION

Ag 0.148 0.186

o Ni

0.063 0.120 O

COEFFICIENTS

Zn 0.242 0.294 O

Poly­styrene 0.58 0.63 (6)

Al 0.122 0.218 (l)

Poly­propylene

0.52 0.80 (6)

In 0.240 0.304

o Poly­

ethylene 0.50 1.10 (6)

Pb 0.250 0.283 O

PP (91) (a) PP (68) (a) PP (52) (a) PP (36) (a) PP (13) (a) PE(9)

0.57 0.80 (6)

PE (32) 0.52 1 (6)

PE (48) 0.50 1 (6)

Hershey—Journal of Applied Polymer Science, vol

PE (64) 0.50 1 (6)

. 10 (1966),

PE (88) 0.50 1.1 (6)

, pp. 1455-1473.

It is not in the scope of this survey to give precise values of the numerous mechanical properties which can be considered (yield, rupture, impact and fatigue strength under tensile, compressive, shearing or torsional conditions), because a great number of data has been given in a copious literature.

The main materials to be considered are structural alloys and laminates. Pure metals have generally a limited mechanical strength, despite an improvement at low temperature.

As regards alloys, the general trends of evolution of the properties at low tem­perature is a moderate increase of young modulus (of the order of 10%), a more or less marked improvement of yield strength (of the order of 30 to 100%), and an important rise of rupture strength by a factor of 2 or 3. The elongation are often not reduced, even at 4°K.

The most used alloys are the different brands of stainless steels, with a Young modulus in the range of 20,000 kg/mm2. Their yield and rupture strength depends strongly on their metallurgical state. Between 300 and 4°K, they respectively vary from 20 to 70 and from 60 to 130 kg/mm2 when annealed; in the cold worked state, they vary between around 100 to 160 (yield strength) and 130 to 250 kg/mm2 (rupture

360

strength). These figures are only given as representative examples, and literature has to be consulted for accurate values. For instance, one can refer to the following papers: Nachtigall-Klima-Freche, Journal of Materials, vol. 3,2, 1968, pp. 425-443; Mechanical and physical properties of the austenitic chromium-nickel stainless

f Liquid nitrogen /Bar

2 » " 3£ars

3 " 5Bars>

4 liquid hydrogen /£ar

5 " " 3Bors 6 " * 53ar$

7 Liquid helium I3ar

8 - 3£ars

9 Transformer oil

U^ffJJ Spheres 0= 62,5

Fig. 13

steels at subzero temperatures, published by The International Nickel Co; Metal Progress, vol. 92, 2, Aug. 1967, pp. 106-115; Low temperature mechanical properties of Cu and selected Cu alloys, NBS Monograph 101, 1967.

361

Aluminium alloys are also widely used as structural materials. Their Young modulus is lower than that of steels (7,000 at 300 °K, 8,000 at 4°K). Cold work increases markedly their mechanical properties: the yield strength increases from 12 kg/mm2

in the annealed conditions to 30 kg/mm2 in highly cold worked state at room tempera­ture, and at liquid helium temperature respectively from 25 to 60 kg/mm2. The rupture strength rises from 25 to 60 at 300°K and from 30 to 80 kg/mm2 at4°K. The elongation depends much more on the metallurgical state than on temperature; it lies in the region 10 to 40%. Somes references are: the previously reported Nachtigall-Klima-Freche's paper; Develay and al., 1966 Cryogenics Engineering Conference; Revue de VAlu­minium, Feb. 1962, etc.

Welded zones may present lower mechanical properties than the assembled parts, especially at low temperatures. One can refer to the following papers: for stainless steels: Welding Research, June 1968, pp. 286 S - 288 S; for aluminium alloys: Welding Research Supplement, July 1965, pp. 317 S - 326 S and Welding J., 47,10 October 1968, pp. 462 S to 471 S.

Plastic materials are very weak materials if used without a reinforcement. Except Teflon, they often cannot even stand their internal stresses if cooled in bulky volumes. However, thin films or coatings behave quite correctly at low temperatures, especially polyethylene terephtalate films and epoxy coatings or various enamels. This is the reason why literature reports mostly data on reinforced resins: the most generally used are epoxies with a reinforcement made of glass cloth or roving {Advances in Cryogenic Engineering, vol. 11, pp. 470-477; Materials Engineering, vol. 66, 6, Nov. 1967). Nevertheless phenolic, polyester or silicone laminates can also be considered. Materials with the best low-temperature mechanical behaviour have always a high percentage of charge (60 to 80%). Adhesives, particularly epoxies, may give very good bonds between two laminates or between a laminate and a metal, or between two metals, but careful technological studies have to be made in order to get the best geometrical configuration to the joint.

IV — DIELECTRIC PROPERTIES

The study of dielectric properties of materials at low temperature is a very wide field, and most ot the literature on this subject reports values measured at rather moderate voltages.

However, the design of cryoelectrotechnical machines requires a large amount of data measured at high voltage, in view of applications in transmission lines, trans­formers, electromagnetic storage coils with fast discharge, etc.

High voltage dielectric strength has been measured mainly by Matthes in the United States, and by a cooperative action engaged in France by ALSTHOM, LCIE and C.G.E. Such tests need an expensive infrastructure ard especially trained teams.

The first results of research show that liquid hydrogen and nitrogen, particularly under a moderate pressure, present a very high dielectric strength both in breakdown and flashover tests. Liquid helium is a rather poor dielectric (fig. 13)

The parameters involved in the electrical breakdown of solids insulators are very numerous. If the possible cracks due to thermal contraction are avoided, the main phenomena to be considered are the intrinsic breakdown, the thermal breakdown and the different effects connected with corona discharges.

As regards the intrinsic breakdown, the impurities and physical defects, which are known to be very harmful at normal temperature, might well be favourable at low temperatures, because they could noticeably reduce the mean free path and the ionizing energy of the charged particles responsible for this type of breakdown.

362

At low temperatures, the dielectric constant is reduced because of the freezing of the motion of polar groups. The ionic and electronic currents become extremely small, often as well as the loss factor. Accordingly, the thermal effects are very low, but, as the thermal conductivity is also quite reduced, it is by now impossible to ascertain if a thermal breakdown is likely to happen. The] measurements of the

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363

The problem of corona discharges presents itself under an aspect different from the phenomena known at room temperature, because the gaseous inclusions are necessarily of a different nature at low temperature, and because their dielectric strength may be not much lower than that of cryogenic liquids. Under this respect, the behaviour of insulators in liquid nitrogen and hydrogen is generally better than in liquid helium. This is particularly the case for surface discharges, which may extend on very large distances in this fluid. This effect is important enough to give a break­down strength of solid insulators in liquid helium much lower than the figures obtained under vacuum at 4.2 °K on the same materials. Even in liquid hydrogen and nitrogen, the influence of discharges may explain that a stacking of thin non-porous films often presents a noticeably smaller breakdown voltage than a single film of the same total thickness.

Anyway, the general trend at low temperature is that, under the form of thin films, plastic materials generally have a better dielectric strength than at room tempera­ture. Research is going on for larger thicknesses, and a great amount of work remains to be done. The difficulty for such tests is to make a reasonable number of tests without bringing back the experimental apparatus at room temperature, so as to reduce the consumption of cryogenic fluids and to gather a sufficient number of data within a limited time.

A still larger amount of work is necessary before obtaining reasonably optimized insulating systems. The materials and structures may be quite different from those in practical use at room temperature, though paper impregnated with liquid nitrogen or hydrogen might give good results.

The hypothesis that the operating gradients on electrical insulation in liquid nitrogen or hydrogen might be double or triple of those in use in the present designs does not seem for the moment unreasonable. This could add a considerable interest in cryoelectrotechnics, besides the gain in the electrical resistivity of conductors.

CONCLUSION

The study of the behaviour of materials at low temperature offers a wide field of interest to searchers: quite new phenomena are met with, some giving rise to a widespread range of values of some properties, often extending on several orders of magnitude. The engineer in charge of the design of a cryoelectrotechnical apparatus has accordingly to face indeed new problems, and to gather a great number of data which are largely lacking in the technical or scientific literature. However, his enthu­siasm is supported by the faith in the final result: large superconducting magnets already enter the industrial phase of development, and it may be hoped that motors, transmission lines, transformers will follow within the next ten years, in order that wholly cryogenic electrical systems may be designed in the medium-term future. Such systems could then be economically competitive to conventional electrotechnics, with an improvement in the efficiencies, the power-to-volume ratio and the limit of realizable unit-ratings.

364

PARAMETERS INVOLVED IN THE ELECTRICAL BREAKDOWN OF LIQUID HELIUM

J.M. GOLDSCHVARTZ and B.S. BLAISSE Department of Physics, Delft University of Technology, Delft {The Netherlands)

Parametres impliques dans le claquage electrique de Thelium liquide

RESUME : / / existe deja des applications technologiques de la supraconductivite et Von a en vue la transmission d'energie a Vaide des supraconducteurs. Vimportance de la connaissance et de la comprehension du claquage electrique de Vhelium liquide apparait ainsi. Au cours des dernieres annees les auteurs ont intensifie leur etude et ameliore les instruments en vue d'obtenir des resultats plus surs et plus reproductibles des claquages. Les auteurs ont eu essentiellement a traiter d'un grand nombre de parametres ayant une influence sur les caracteristiques du claquage electrique de Vhelium liquide, c'est-a-dire qu'ils ont du fixer et controler Vun quelconque de ces parametres, a savoir : temperature, distance entre les electrodes, purete de Vhelium liquide, purete des electrodes, nature, forme et etat de la surface des electrodes et temps d'oxydation, destruction des electrodes et son influence, duree du claquage, temps entre ceux-ci et effets de precontrainte, controle des preclaquages, Constance de la vitesse a"elevation de la tension appliquee et nature du courant : courant continu, courant alternatif ou impulsions.

Ce rapport essaie de decrire en detail la maniere dont les auteurs ont essaye de regler les parametres indiques ci-dessus.

INTRODUCTION

The present and future applications of superconductivity, particularly the future possibilities of the transmission of power by means of superconductors, mean that the electrical breakdown of liquid helium has become a matter of great importance. In fact, liquid helium can act both as an insulator and as a means to obtain the low temperature necessary to reach the transition temperature of most of the metals to become a superconductor. Nevertheless, as far as the literature reveals, hitherto little attention was paid to this subject. Only a few groups are working on the electrical breakdown of liquid helium, namely: Mathes [1] in the United States of America, Fallou and Galand [2, 3] in France, and ourselves [4, 5, 6, 7] in the Netherlands, although Blank and Edwards [8] published a paper on this matter in 1960.

The aim of this paper is by no means to present either final results or to be an up-to-date resume. Yet, in view of the current importance attached to the knowledge of the electrical breakdown of liquid helium, this paper is meant to indicate our progress in this matter and to criticize it, considering the different parameters involved, how these parameters could be commanded and moreover the fact that the results obtained were always rather disconcerting. Those parameters are. — Purity of the liquid helium; — Gap distance between the electrodes; — Temperature; — Speed of rise of the voltage applied; — Time lag between breakdowns; — Duration of a breakdown and destruction of the electrodes; — Prestressing effects.

PURITY OF THE LIQUID HELIUM

We finished a previous paper [6] on the electrical breakdown of liquid helium showing the situation with the representation of the experiments made till that time

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(the end of 1963). Figure 1 shows the curves taken at boiling point (4.2 K) and at different temperatures below the Appoint of the breakdown strength as a function of the gap distance. The curves were plotted with the highest values obtained for each gap distance and the rise of the voltage was controlled by hand. This procedure of taking the highest values was based on the fact that we considered these values to be the nearest approximation to the values of the intrinsic breakdown strength. Obser­ving these curves we concluded that the more we improved the purity of the liquid helium, the higher the values for the breakdown voltage.

In figure 2 the curves are drawn which were obtained with the equipment which has been considerably improved upon in the last few years. Each point represents the mean value of fifty values of the electrical breakdown for each gap distance. For the purpose of comparison with the curves of figure 1, the curves corresponding to the highest values of the breakdowns are also drawn and the scale is the same in both figures. For simplifying things and for practical reasons, only one batch of liquid

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helium was used in the Dewar during a whole one day experiment, and the measure­ments of figure 2 were performed only at boiling point.

It is evident that curves of figure 2 are lying in a lower region than those of figure 1, even those corresponding to the highest values. Here must be remarked that for all experiments of figure 2 we used a Vycor glass superleak as a filter to keep the liquid helium free of macroscopic impurities, whilst a superleak was not used in the experi­ments of figure 1. Here must be also pointed out that in the case of figure 2e, we had a crack in the Vycor glass superleak. Probably, we had less pure helium. Nevertheless, the curves of figure 2e are shifted to a higher position, This fact would indicate a contradiction: the liquid helium is less pure and the values of the breakdown voltages are higher than those of the curves of figure 2a, b, c, d. Features of the curves of figure 2 are shown in the preceding table.

As a matter of fact, the control of the purity of the liquid helium is very difficult if not impossible, and one can never know exactly how clean the cell and the helium are. Before each whole day experiment the cell was chemically cleaned and then, after being mounted, it was evacuated. But in the cases of figure 1 and figure 2 there are some important differences. In the former the apparatus was pumped during one night before filling it with liquid helium; after this only a moderate vacuum of the order of 10" 1 torr was obtained. On the other hand, in figure 2, apart from the fact that all the experiments were made with the improved equipments [5] plus all the changes described here, the dismountable bottom part of the cell which contains the electrodes, figures 3 and 4, was always cleaned in an ultrasonic system. Then this part was mounted and a vacuum of the order of 10" 5 torr was maintained for about two days. Even so the curves are kept in a lower region.

THE GAP DISTANCE

The gap distance is a relevant parameter in the study of the electrical breakdown and one should be able to control it very carefully. In the case of the electrical break­down of liquid helium the control of this gap distance is not at all simple, for the gap distance must be changed and controlled from the outside of the cryostat, i.e., from the outside of the cell which is inside the liquid helium Dewar which, in turn, is inside the liquid nitrogen Dewar. Since the new test cell for the electrical breakdown of liquid helium was designed [5], we have introduced numerous changes, some slight and some important for the better performance of the apparatus as a whole. But the principle of operation remained the same.

That which makes things more complicated is that one should be able to turn the support P of the six different cathodes k by pushing the pin p of figures 3 and 4, such that each one of the cathodes can face the anode a. This rotation is commanded from the outside by means of the magnets m fixed on the mechanism M, figure 5. The anode is fixed and faces one of the slit-windows of the Dewars. Thus, one can always see the electrodes from the outside.

To change the gap distance between the electrodes, the support P of the cathodes is pushed down or up by means of the action of the transmission T and the spring v2, figures 3 and 4. This transmission T is commanded by the head H of the apparatus which is in the upper part and outside the cell, figure 5. The same transmission is also used to rotate the cathodes. The distance between the head H of the instrument and the support P, figures 3 and 5, is approximately one metre and so is also the transmission T, because the electrodes system must be deep in the liquid helium bath if one wants to make an appreciably long lasting experiment, that is, to obtain a sufficiently large number of breakdowns to give some statistical sense to the whole experiment. The average number of breakdowns we performed in each one of the last experiments, figure 2, is of the order of one thousand.

370

Fig. 3

To measure the gap distance and its changes, we now use the clock micrometer C of figure 5 instead of the normal micrometer in the former equipment. The new micrometer is a very sensitive and accurate instrument (*) : 1 division = 1 urn. It has the advantage that its bottom ruby spherical surface is always pressed by a spring on the head H of the cell which can be moved up and down. This avoids the error

(*) C.E.J. Mikrokators No: 509/4, Eskilstuna, Sweden.

371

introduced each time a contact was made between the micrometer and the head of the cell in the equiments used before. The anode a, figure 3, obviously had to be mounted on an insulation material, in this case a thermoplastic called Delrin. This

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does not permit the outgassing of the electrodes by means of an induction oven. However, it seems from the reproducibility of the curves of figure 2, that this would not be an important drawback.

With this configuration of the electrodes, and by controlling everything from the outside of the cryostat, we reached a very good constancy of the gap distance; a maximum error of 2.5 urn in all the measurements of the curves of figure 2. This

372

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means that 5 urn was the maximum change measured in the gap distance before and after each set of fifty breakdowns.

TEMPERATURE

Temperature is another parameter that we controlled in our experiments on the electrical breakdown of liquid helium, although in the last experiment (those represented

373

in figure 2) we concentrated our attention only on the measurements at boiling point (4.2 K) due to the uncertainty of our results of previous experiments. However, we can fix any desired temperature between 4.2 K and 1.3 K using a very simple and reliable manostat, as described by Walker [9], but which we improved upon consider­ably.

OTHER PARAMETERS

The electronic equipment permits the direct control of the speed of rise of the applied voltage and the time lag between breakdowns. Moreover it cuts off the break­downs and thus diminishes the destruction of the electrodes. The total electronic automatic system shown in figure 6 will be described in general. The automatic remote

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control (4) controls the voltage supply and the speed of rise of the applied voltage to the cell (1) through the electronic relay (3). The cut off time of the electronic relay diminishes the time of the cut off from 0.03 sec, cut off time of the electro-mechanical device of the power supply (2) to 20 x 10 ~6 sec [7]. The following table gives the features of this automatic remote control and the power supply.

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The number of breakdowns to be accomplished is predetermined in the unit (5) and also the time interval between them in unit (6). The time between one breakdown and the starting point of the rise of voltage for the next one can be set from 10 sec to 9 min 50 sec in steps of 10 sec. The minimum time of 10 sec was chosen because it is somewhat greater than the time necessary for the automatic resettlement of the zero of the potentiometer of unit (4).

Unit (6), which in fact is mounted together with units (5) and (4) in only one unit, delivers the "start" and "s top" signals to the digital voltmeter through unit (7). Unit (9), a binary to decimal code converter, makes the printer (10) deliver the break­down voltages.

374

The results hitherto found concerning the destruction of the electrodes, the prebreakdowns and the prestressing effects were already described in previous papers [6, 7].We did not yet carry out experiments on the time lag.

CONCLUSION

On the one hand, the high values of the dielectric strenght in figure 1 could be due to the high purity of the helium. They could also be due to the presence of dissolved (possibly metastably) oxygen in the liquid helium which increases the strength of liquid dielectrics, according to Swan and Lewis [10, 11]. On the other hand, in the case of figure 2, in addition to the fact that these curves were obtained with a highly improved equipment, they could also be a better approximation to the physical reality. They were obtained statistically, taking the mean of fifty measurements for each gap distance, and each of these sets of fifty breakdowns were preceded and followed by sets of fifty breakdowns at a reference gap distance of 100 urn.

This means that, for the time being, almost all parameters involved in the electrical breakdown of liquid helium can be controlled to a reasonable extent and it seems that the values for the dielectric strength of liquid helium at boiling point would be provisionally given by those curves of figure 2.

However, the situation is still uncertain and a considerable number of experiments should be carried out in order to obtain more consistent and conclusive results.

ACKNOWLEDGEMENTS

The authors wish to thank Ir. F. Landheer, Mr. A. C. Ouwerkerk and Mr. P. C. Sla-gter for their assistance in the realization of the experiments and Mr. J. R. de Haas for his useful help with the electronic devices.

REFERENCES

[1] K.N. MATHES, I.E.E.E. Transitions on Electrical Insulation, EI-2, 24 (1967). [2] B. FALLOU and J. GALAND, Rev. Generate de VElectricite, 11, 594 (1968). [31 J. GALAND, C.R. Acad. Sc. Paris, 266B, 1302 (1968). [41 B.S., BLAISSE, A.v.d. BOOGAART, Annex 1958-1 Bull. I.I.R., pp. 333-340, Com. I. [5] J. M. GOLDSCHVARTZ and B.S. BLAISSE, Cryogenics, 5, 169 (1965). [6] J .M. GOLDSCHVARTZ and B.S. BLAISSE, Brit. J. Appl. Phys., 17, 1083 (1966). [7J J .M. GOLDSCHVARTZ and B.S. BLAISSE, Appl. Sci. Res., 19, 14 (1968). [8] C. BLANK and M.H. EDWARDS, Phys. Rev., 119, 50 (1960). [9] E. J. WALKER, Rev. Sci. Instr., 30, 834 (1959).

[10] D. W. SWAN and T.J. LEWIS, Proc. Phys. Soc. (U.K.) 78, 448, (1961). [11] D.W. SWAN, Proc. Phys. Soc. (U.K.) 78, 423, (1961).

DISCUSSION

D.A. SWIFT (U.K.) — Has the influence of the cathode, both with regard to the type of material and the surface finish been considered ? Perhaps this could explain some of the results obtained at very small gaps in Mr Goldschvartz's paper.

J.M. GOLDSCHVARTZ — As far as the material of the cathode is concerned, it does have an influence since the emissivity is different but we are keeping the same material, in order both to avoid the introduction of new parameters and to be able to compare the new experiments with the previous ones. As far as the surface finish is

375

concerned, it also has an influence since at very small distances the irregularities could introduce errors in the gap distances and the local electric field is changed. Neverthe­less, we have evidence that breakdowns do not take place preferentially from the craters produced previously.

J. GALAND (France) — a) You gave us an impressive list of parameters acting on the breakdown values of liquid helium. Nevertheless, I wonder whether you have forgotten some of them.

b) First of all, what do you think of the action of the bubbles ? c) Then, don't you think that, instead of the temperature itself, it is the pressure

you have above the liquid, pressure which is lowered in order to lower the temperature, which acts as the actual parameter ?

J. M. GOLDSCHVARTZ — a) Of course, we did forget a lot of parameters. b) There is a model or " ad hoc" explanation for electrical breakdown for hydro­

carbons made by Sharbough in the U.S.A. and which could be extended to some liquified gas as N 2 , 0 2 , A, etc., but not to liquid helium. It might be extended, even­tually, to liquid helium I, but never to liquid helium II, for its heat conductivity does not permit the formation of bubbles.

c) I do not quite understand the question. We measure the temperature by measuring the vapour pressure, so that I think it is exactly the same thing. For each pressure there is a temperature and vice-versa.

R.J. MEATS (U.K.) — Is the variation of ambient pressure (i.e. helium gas pressure in the cryostat) a possible cause of scatter in your results ?

J.M. GOLDSCHVARTZ — No, I do not think so. Eventually, it would depend upon the order of magnitude of the ambient pressure. However, for the scatter shown the differences in temperature were never big enough.

376

COMPORTEMENT DIELECTRIQUE DE L'HfiLIUM LIQUIDE ET HYPERCRITIQUE SOUS TENSION ELEVEE

B. FALLOU (Mme) et J. GALAND Laboratoire Central des Industries Electriques, Fontenay-aux-Roses

J. BOBOet A. DUBOIS Laboratoires de Marcoussis — Centre de Recherches de la C. G.E., Marcoussis {France)

High voltage dielectric behaviour of liquid and hypercritical helium

SUMMARY: The authors have used, for these experiments, several vessels allowing measure­ments to be made in a temperature range from 1.5 to 20 °K. One of these vessels which can be used with voltages higher than 100 kV, is provided with large windows through which it is possible to observe breakdown phenomena.

The dielectric strength of liquid helium has been measured at 4.2 °K, in uniform AC field, for gaps ranging from .1 to 10 mm. It has also been measured in non-uniform field {point-to-plane electrodes), for similar gaps, under AC and DC conditions. The results obtained in those different cases show a wide scatter of dielectric breakdown values, and, moreover, in non-uniform field conditions, an important polarity effect and several pre-breakdown phenomena.

A study of the dielectric behaviour of superfluid helium {down to 1.5 °K) and hypercritical helium {up to 10 bars) has also been carried out. Values obtained in both cases pointed out no actual differences in behaviour between these fluids and normal boiling helium.

1. INTRODUCTION

Le developpement d'une cryoelectrotechnique utilisant des materiaux supracon-ducteurs sous tension elevee necessite une connaissance precise du comportement dielectrique des differents materiaux d'isolation. Pour coordonner les travaux de recherche effectues en France, dans le domaine de la cryoelectricite, un groupe de travail a ete cree, sous l'egide de la Delegation Generale a la Recherche Scientifique et Technique.

Les travaux relatifs a Thelium liquide, effectues dans ce cadre, ont pour but, d'une part de fournir des donnees utilisables pour des projets de materiel, a savoir les variaiions de la rigidiie dielectrique sous tension alternative en fonction des divers parametres et notamment de la distance, d'autre part d'acquerir quelques informations sur les mecanismes de disruption en utilisant des champs continus tres divergents. C'est ainsi qu'on a ete amene a etudier non seulement le comportement de Thelium a 4,2°K, mais aussi celui de l'helium superfluide et hypercritique.

2. DESCRIPTION DE L'APPAREILLAGE

Trois cryostats differents ont ete utilises. Pour les mesures aux grands ecartements, un cryostat special de diametre interieur

550 mm a du etre realise par Tun des laboratoires* (fig. 1). La conception de cet appareil est originale sur plusieurs points : Trois fenetres de grandes dimensions ( 0 = 300 mm) permettent les examens des

phenomenes selon deux directions orthogonales. La mise en froid de l'ensemble peut etre assuree par deux voies associees ou

separees a volonte : un cryogenerateur est relie par une conduite sous vide au cryostat et assure au niveau du recipient interieur de Tappareil deux zones froides, Tune a

♦Laboratoires de Marcoussis-Centre de Recherches de la C.G.E.

377

80°K au voisinage de la partie superieure, l'autre a 20°K au voisinage de la jonction d'un ecran passif avec I'enceinte interne.

Fig. 1 — Schema du cryostat 0 550 int£rieur. Construit aux laboratoires de Marcoussis.

378

De plus, une circulation de fluide peut etre introduite dans la partie inferieure de cette enceinte interne.

II en resulte une possibility de descente tres rapide en temperature, sans utiliser un refroidissemnt prealable a l'azote, et une economie de fonctionnement interessante.

Une temperature de 50 °K est atteinte avec le seul apport du cryogenerateur fonctionnant pendant une vingtaine d'heures. Apres remplissage avec de I'helium liquide, on observe alors une consommation de 1 litre de liquide par heure, ce qui autorise de longues campagnes d'essais sans rechauffement.

Pour les mesures aux plus faibles ecartements, tant a pression normale qu'en suprafluide, un second cryostat, de conception analogue, mais de diametre interieur 60 mm seulement, a ete utilise. II est muni d'une garde d'azote et de deux hublots diametralement opposes. Sa consommation, inferieure a 0,1 1/h, donne un autonomie d'environ 24 h. Une pompe, d'un debit de 25 m3/h permet de descendre a des tempe­ratures de l'ordre de 1,5 °K.

Quant aux mesures sous pression elevee, elles sont effectuees dans une enceinte de diametre interieur 550 mm et permettant des mesures jusqu'a une pression de 20 bars.

Des traversees a haute tension ont ete etudiees specialement pour chacun de ces appareils. Dans le premier, on utilise un tube d'acier inoxydable enrobe d'une resine epoxy chargee qui permet d'atteindre 100 kV en valeur efficace. Cette traversee supporte Tune des electrodes ainsi que le dispositif de reglage de l'ecartement qui s'opere par vis micrometrique. Ce reglage est, dans certains cas, controle par visee optique. La mise a zero se fait par verification electrique du contact des electrodes.

Pour le second, malgre ses dimensions reduites, une traversee en quartz metallise, de diametre 30 mm, a permis d'atteindre des tensions superieures a 50 kV. Quant au troisieme, du fait des problemes d'etancheite poses par l'utilisation de I'helium sous pression elevee, il a ete muni d'une traversee constitute d'un cable coaxial dont les extremites sont noyees dans une resine polymerisee sous vide in situ.

Pour limiter la degradation des electrodes au cours des ruptures successives et du fait de l'impossibilite de leur remplacement en cours d'experience, un double dispositif de limitation du courant d'arc a ete utilise en alternatif : limitation en intensite par une resistance serie et en temps par un trigatron. En continu, des alimen­tations a faible debit ont ete utilisees.

Les temperatures ont controlees, au niveau des electrodes, par des thermocouples or-fer (0,03 %)/chromel.

3 . COMPORTEMENT DIELECTRIQUE DE L ' H E L I U M A PRESSION ATMOSPHERIQUE

3.1 Tension disruptive en champ uniforme. — Pour la determination de ces valeurs, on a utilise un systeme d'electrodes sphere-plan, en acier inoxydable ou en cuivre. Pour les petits ecartements, jusqu'a 3 mm, l'electrode spherique a un diametre de 15 mm. Au-dela, son diametre est de 100 mm. La tension etait appliquee lineairement de facon uniforme.

Les resultats des mesures effectuees dans ces conditions a 50 Hz sont reportes figure 2. La courbe a ete tracee a partir de points obtenus en faisant la moyenne de nombreuses valeurs individuelles reparties sur plusieurs remplissages des cryostats.

Dans un champ pratiquement uniforme, on constate que la tension disruptive n'est pas une fonction lineaire de l'ecartement. Si, aux forts ecartements, la rigidite dielectrique de I'helium parait peu elevee, il n'en reste pas moins que, pour de petits intervalles, les valeurs obtenues sont comparables a celles de bon nombre d'autres dielectriques liquides, et notamment de gaz liquefies. En effet, pour un ecartement de

379

1 mm, les 22 kV de l'helium sont comparables aux 28 kV de l'azote, de l'hydrogene ou de 1'huile de transformateur.

Les observations, tant optiques qu'electriques, effectuees au cours de ces mesures, n'ont permis de mettre en evidence aucun phenomene predisruptif.

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e (mm)

Fig. 2 — Tension disruptive de Thelium liquide a 4,2 °K en fonction de l'ecartement des elec­trodes.

3.2 Rupture en champ divergent. — Les electrodes, dans ce cas, etaient constituees du meme plateau que precedemment et d'une pointe d'acier, de rayon de courbure a l'extremite voisin de 35 urn. Les mesures de tensions disruptives, effectuees dans ces conditions, ont fait apparaitre un effet de polarite extremement important, comme le montrent les courbes de la figures 3 ou les valeurs de tension reportees dans le cas d'essais en alternatif sont les valeurs de crete. Le comportement en pointe negative est comparable a celui a frequence industrielle, bien que, pour une raison encore mal connue, les valeurs de crete dans ce dernier cas soient superieures a celle obtenues dans le premier. Par contre, en pointe positive, la tension de rupture reste constante pour une grand plage d'ecartements. Cette tension correspond a des champs, a l'extremite de la pointe, voisins de 106 V. c m - 1 pour lesquels un certain nombre de phenomenes predisruptifs sont apparus.

Tout d'abord, une lueur rouge prend naissance dans la zone a champ tres intense. Selon toute vraisemblance, cette lueur correspond a une excitation des atomes d'helium.

A l'apparition de cette lueur correspond, sous tension continue, la circulation d'un courant entre les electrodes. Une observation oscillographique a montre que ce courant n'etait pas de nature impulsionnelle, mais continu et tres stable. Son intensite croit tres rapidement avec la tension pour atteindre des valeurs aussi elevees que plusieurs micro-amperes.

D'autres part, ces phenomenes optiques et electriques s'accompagnent de mouve-ments de liquide, celui-ci etant chasse par la pointe et anime d'une vitesse elevee, ainsi que d'une ebullition tres violente, l'espace voisin de la pointe s'emplissant alors d'une tres grande quantite de bulles de petites dimensions.

380

3.3. Dispersion des resultats — Les valeurs individuelles de tensions disruptives presentent des dispersions elevees, le coefficient de variation etant, en moyenne, de l'ordre de 20%. Les diverses tentatives d'explication n'ont pas encore abouti a une comprehension complete de ce phenomene.

( y — pointe negative

Q — pointe positive

j ) _ t e n s i o n alternative (valeurs de

Fig. 3 —Tensions disruptives de Thulium liquide en champ divergent.

La premiere hypothese supposait une action des particules solides en suspension dans l'helium liquide de qualite industrielle : air, eau, particules metalliques, etc... Une introduction massive d'oxygene dans le cryostat n'a permis de mettre en evidence aucune modification ni de tension disruptive, ni de dispersion. II est cependant possible que les impuretes existant au prealable dans le liquide aient une influence et que le fait d'en apporter une grande quantite n'ait que peu d'effet supplemental .

Par ailleurs, il etait possible d'imaginer une action des bulles existant en quantite relativement importante dans le liquide a sa pression de vapeur saturante.

Des mesures de decharges partielles, effectuees avec un appareil permettant de detecter des impulsions de charge apparente inferieure a 5 pC n'ont permis de mettre en evidence aucun phenomene predisruptif, ce qui pourrait indiquer que, si une decharge se produit dans une bulle, elle conduit a une rupture totale de l'helium situe entre les electrodes. Par ailleurs, lors de mesures de rigidite dielectrique de l'helium gazeux, a des temperatures peu superieures a 4,2°K, on observe une lique­faction due vraisemblablement a l'electrostriction, ce qui permettrait de supposer que, sous champ intense, il est peu probable que des bulles puissent exister dans la zone interessee.

Enfin, aucun courant de conduction n'ayant jamais pu etre mesure dans ce fluide, en champ uniforme, jusqu'a la rupture, on pouvait supposer qu'une absence d'ions etait a l'origine de la grande dispersion constatee, la disruption pouvant alors etre initiee par l'arrivee aleatoire d'un rayonnement naturel ou cosmique. Dans ce sens, une source d'iridium 192, de 15 mCi, a ete disposee a l'interieur de l'une des electrodes de fa^on a en extraire des electrons par effet Compton. Par ce procede, on a pu effectivement constater une reduction importante simultanement de la valeur moyenne

381

des tensions disruptives et de la dispersion, cette derniere etant reduite a des coeffi­cients de variation de l'ordre de 5 % comme le montrent les courbes de la figure 4. II est done evident que la dispersion est attribuable, en partie tout du moins, a l'absence d'ion germe dans la zone soumise au champ.

II convient cependant de noter que l'effet de cette source d'ionisation devient nul lorsque le liquide cesse d'etre en ebullition. La presence d'ions n'est done pas une condition suffisante pour l'amorcage d'un arc a champ constant. La presence d'une phase gazeuse semble necessaire, en admettant que Fexistence, dans le champ, d'une bulle unique peut conduire a la disruption totale et immediate de l'ensemble du dielectrique.

10 15 20 25 30

Fig. 4 — Repartition des tensions disruptives a 4,2 °K (ecartement 1 mm).

En champ divergent et en l'absence de phenomenes predisruptifs, les coefficients de variation sont du meme ordre de grandeur qu'en champ uniforme. L'apparition de ces phenomenes predisruptifs reduit considerablement la dispersion des valeurs de tensions disruptives, ce qui, a la lueur de l'hypothese emise ci-dessus pourrait etre du a la presence d'ions dans le liquide associee a la formation de bulles dont, even-tuellement, des bulles crees a la surface de la pointe par une cavitation resultant du mouvement du liquide.

4. RUPTURE DE L'HELIUM ENTRE 10 MILLIBARS ET 10 BARS

4.1 Press ions infer ieures a la pression atmospherique. — Les mesures ont ete effect uees principalement en champ uniforme, a divers ecartements, a des temperatures telles que, pour chaque pression, l'helium etait sur sa courbe de vaporisation. Tant en suprafluide (entre 10 et 40 mbar) qu'en helium normal, le comportement dielectrique

382

de l'helium ne presente pas de modification sensible par rapport a celle observee a pression atmospherique. Notons cependant une diminution progressive et reguliere des tensions disruptives lorsque la pression decroit. Le rapport des rigidites dielectri-ques a pression atmospherique et a 10 mbar est de l'ordre de 2.

Aucune modification sensible des dispersions par rapport aux mesures a pression atmospherique n'a ete observee, ce qui tend a montrer que la presence de bulles dans le liquide n'est pas la cause premiere de cette dispersion.

En champ divergent, et en helium suprafluide, l'apparition des phenomenes predisruptifs n'est pas aussi systematique. On note neanmoins le meme effet de polarite que dans l'helium normal.

4.2 Helium hypercritique. — Ces mesures sont faites pour des pressions superieures a la pression critique de l'helium (2,26 bars a 5,2 °K). Elles ont ete limitees a 10 bars du fait des difflcultes rencontrees pour assurer l'etancheite au niveau de la traversee a haute tension. La montee en pression est faite par 1'intermediate d'helium gazeux comprime a temperature ambiante. Le gaz arrive par un serpentin baignant dans l'helium liquide de maniere a ce que sa temperature soit progressivement abaissee jusqu'a l'Squilibre. La figure 5 donne les resultats de mesures pour un ecartement de 1,7 mm en champ uniforme. Elle montre que la rigidite dielectrique de l'helium hypercritique n'est multipliee que par un facteur de l'ordre de 1,5 lorsque la pression passe de 2,3 a 10 bars.

50-

40

30H

20 P (bar)

1 r

Fig. 5 — Tension disruptive de l'helium hypercritique, en fonction de la pression (ecartement de 1,7 mm).

La dispersion des mesures est extremement reduite puisque, pour des series de 5 points, le coefficient de variation ne depasse pas 3 % .

5. CONCLUSION

L'ensemble des resultats exposes ci-dessus montre que, contrairement a ce qui aurait pu etre suppose du fait de la simplicite de la structure de ce fluide et de sa basse temperature d'ebullition, la rupture dielectrique de l'helium liquide n'est pas un phenomene simple. De meme, du fait principalement des points bas dus a une dis-

383

persion elevee, son utilisation sera peut-etre difficile dans certaines applications particulieres. Pour celles-ci, la solution aux problemes d'isolation electrique pourra etre l'utilisation d'un vide pousse qui serait alors susceptible d'assurer, simultanement, l'isolation thermique des supraconducteurs.

Les recherches exposees ci-dessus ont ete effectuees sous l'egide de la Delegation Generate a la Recherche Scientifique et Technique.

DISCUSSION

R. J. MEATS (U.K.) — One of the disadvantages of liquid helium as a high voltage dielectric is the way in which the breakdown voltage does not increase proportionally with electrode gap. The last figure suggests that the change in breakdown voltage of liquid helium with pressure is not discontinuous on passing through the critical pressure. It is of practical significance to know whether under these conditions a « gap effect» still occurs, since the helium can then be regarded as a very dense gas. Have any measurements been made at high pressures with electrode gaps other than 1.7 mm?

Mm e B. FALLOU — The authors agree with Mr. Meats upon the fact that hypercritical helium may be considered as a very dense gas, and therefore might exhibit a « gap effect» which is less important than that exhibited by liquid helium. Up to now, measurements have only been carried out for a 1.7 mm gap, but further study is to be done in the near future.

D. A. SWIFT (U.K.) — Have the authors considered the influence of the cathode, both with regard to the type of material and the surface finish ? Perhaps this could explain some of the results obtained at very small gaps in Dr. Goldschvartz' paper.

jy[me B FALLOU — The influence of the material and the surface finish of the electrodes has been considered in uniform A. C. fields only, so that it is not possible to relate the results to the influence of the cathode alone.

It has been assumed that the metal effect, observed when comparing stainless steel and aluminium electrodes, is in fact the result of their different behaviour under the action of sparks, and seems therefore to be related mostly to the surface finish.

That surface finish appears to be one of the most important parameters, but its actual effect is difficult to put out clearly. As a matter of fact, with polished elec­trodes, higher mean values were obtained, but the scatter of the results was left unchanged. On the other hand, when comparing the mean value of the first ten points to that of the last ten ones, for a series of about 100 results, there is generally no significant difference between them. However, the very first spark occurring with a newly polished electrode does not lead systematically to a higher breakdown value.

J. M. GOLDSCHVARTZ (The Netherlands) — I am interested in two or three points of the curve of figure 2. Could you tell me please, the error with which those points were drawn ? I mean the maximum error, the standard deviation, and how many breakdowns were done for each.

Mm e B. FALLOU — A very large number of measurements of breakdown voltage of liquid helium have been carried out for gap lengths between .1 and 2 mm. For example, the distribution of individual values may be summarized by the following typical statistical figures.

384

Gap Number of Average length individual value mm values kV

.1 50 4.5

.2 100 8.55

.5 115 17.4 1 20 22 2 20 33

Standard Coefficient Range deviation of variation

kV kV %

2 - 6.3 1.07 23.8 4.8-12.5 1.23 14.4 9.2-23.2 2.3 13.2

15.8-27.2 3 13.6 26 -42.2 4.45 13.5

385

SOME PROBLEMS OF HEAT EXCHANGE ON SURFACES OF CYLINDRICAL BODIES

V.G. PRON'KO, L.B. BULANOVA, V.G. BARANOV and L.S. AKSELROD Moscow {U.S.R.R.)

Quelques problemes de l'echange de chaleur a la surface de corps cylindriques

RESUME : On donne les resultats de recherche sur un canal annulaire refroidipar azote liquide. On montre Vinfluence de la longueur du cylindre, de Cespace libre du canal annulaire et du

degre de sous-refroidissement de Vazote liquide sur la temperature de Vechantillon avec differents debits de chaleur.

On etudie le commencement de Vebullition superficielle.

INTRODUCTION

It is essential for designing power and electronic devices to estimate the intensity of the heat exchange on the surfaces of cooled elements. A number of recently published papers deal with the investigation of nucleate boiling of liquid helium and the determination of critical heat fluxes relevant to the solution of problems connected with the cool-down and stability of superconducting devices [1,2]. In addition to studies of heat transfer in the nucleate boiling range, investigations of the heat transfer in the film boiling range and organization of forced circulation heat exchange in cooling ducts are of great interest for designing and exploiting superconducting devices [3]. This paper reports the results of investigations of the influence of some conditions on the heat exchange. The fact that the experimental tests were carried out using only liquid nitrogen is certainly partly limiting the possibility of direct use of these results for designing of superconducting devices. The results given will neverthe­less be useful when the problems of cooldown of aggregate power elements from the surrounding temperature to that of working conditions, are solved and also when optimal geometrical and regime parameters of cooling ducts are chosen.

T H E EFFECT OF SAMPLE SURFACE MATERIAL ON THE SAMPLE COOLDOWN RATE.

During the first stage of the cooldown the temperature of the solid (Tw) is much higher than the temperature of the liquid and it is physically impossible for the liquid to come close to the hot wall. The heat exchange in these conditions occurs in the film boiling mode. The highest temperature at which the liquid can exist is called "the temperature of the liquid superheat limit (Tsl)". When the wall temperature drops below Tsl the conditions appear for wetting of the surface and this occurs at some or other temperature depending on hydrodynamic circumstances of the process. Recently published data [4, 5] show that the lower temperature limit of film boiling can vary widely depending on the thermophysical properties of the surface.

The heat transfer coefficients and the temperatures of film-nucleate boiling tran­sition were obtained with the non-stationary method while treating the cooling curves of hollow cylinders, from 4 to 85 mm o.d., 300 mm length and of 0.5-5 mm wall thickness. The cylinders were fabricated of different materials. After immersing the samples in liquid nitrogen the temperature decreased from 300 °K to 80 °K. The preliminary analysis of the known data and experimental results have shown that the cooling of all samples went in the regular (quasi-steady) regime.

387

The sample cooldown time depends considerably on the temperature of transition to nucleate boiling which is a more intensive heat transfer process than stable film boiling. The wall temperature at which the temperature curve has sharp break shows the lower limit of the stable film boiling. This limit is somewhat diffused, and for exact determination of the temperature of the "film boiling crisis" the temperature distribution curves were drawn for 257 tests. Figure 2 shows the difference of boiling crisis temperature for steel and copper cylinders. The crisis temperature dependence on surface material is even more obvious when one considers cooldown of stainless steel cylinders coated with a thin layer of teflon. One can conclude from experimental results that for materials with low thermal conductivity the boiling crisis occurs at higher temperatures of the cooled surface.

Visual observations and cinema- and photographing of film boiling process made by different authors [6, 7] indicate the wave character of the vapour film movement, the vapour film thickness being of the same order as wave amplitudes. Not considering the wave origin we shall only mention that owing to periodical changes of film thickness surface-non-uniform character of heat transfer intensity takes place. Owing to rather high frequency of film oscillations the surface response reaction appears, that is the temperature fluctuations along the surface. Surface temperature at each point oscillates with the film oscillations frequency.

Using the solution of the temperature wave diffusion problem for stationary-periodical oscillations of surrounding medium temperature one can describe for film boiling the increase of surface temperature oscillation amplitudes by the number

\JX-c- p - c o

where a heat transfer coefficient; X, c, p thermal conductivity, specific heat capacity and density of material respectively; co = 2 7iv where v—film oscillations frequency

T ^ m a x - T « ; m i n = / ( B i * )

Assuming for non-stationary heat regime (cooldown of solid) the maximum local surface temperature being equal to the mass averaged wall temperature TWmax = Tw we have:

AT = T ^ - T , , , = f a — W Wmin J I

\jX-c- p*co

For materials with lower heat conductivity the temperature oscillation amplitude would be higher, which agrees with the experimental results.

After the minimum local surface temperature values reach some critical value the film boiling crisis occurs. The experimental data analysis for large volume boiling shows that the crisis temperature is near to the temperature of superheat liquid limit. Assuming that the boilingcrisis surface temperature for the process considered remains constant T* = Tsi one can regard Tcr to be a function, of Bi*. Not considering the possible changes of steam film oscillation frequency and heat transfer coefficients, i.e. assuming that co = const, and a = const., one can write Tcr = / (1 /yXc .p) .

The ratio of surface temperature oscillation amplitudes for surface materials with different properties can be presented as follows:

ATi = / ( / i - c -p) 2

AT2 V ( ^ c - P ) i

388

Teen cy€LncL€ZS coated with cpn-7jl^

"90 100 ^ 110 120 130, 160 170 180 100 rp'K F^Ql dLfixtfuilon of the Jnifi, temptzatuze (ret) /or lo/ne —'— matet/Zoes (n -tpanUty of tedis)

1 r- V o M A* £4 0>*i • 0,6 0,7 , qs 0,3 / 1,0 rMM JjfrJL teducins cooCdov/n ttmk vs thickness of ° MM

various coailnp

389

The experimental data on crisis temperature for steel samples with and without the teflon coating brings us to the following results:

T ; rr - T * _ 180 -107

T ; r - T * ~ 1 1 5 - 1 0 7 ~

where T'cr and T"r—mass-averaged wall temperatures of crisis for steel samples and coated samples. The theoretical value of this ratio for the same materials is

AT;; = KX-c-p)st= /7,900-0.068-9.5 = n

AT;r \l(X'cp)ft V 1,400-0.15-0.22

One notices the rather good agreement which supports the correctness of initial assumptions. We should mention that the effect of teflon coating on heat transfer to liquid depends on the thickness of teflon layer. The crisis temperature moves towards higher values and cooling time decreases as the thickness of teflon coat increases. With further increase of teflon coat thickness the cooldown time begins to grow because of increase in thermal resistance of coat layer (fig. 1). The optimum thickness of coat Teflon-734 lies between 80-120 ^. In case of different coating materials the qualitative character of coat thickness influence on cooling rate is preserved (fig. 3).

The tests carried out confirmed the possibility of the essential shortening of cool-down time of cryogenic devices as a result of coating the cooled surface with materials possessing certain thermophysical properties. One can estimate the usefulness of any material as a coat with the help of the given relations.

THE INVESTIGATION OF HEAT TRANSFER FROM THE INNER WALL OF AN ANNULAR DUCT

This work was aimed at investigating the heat transfer and the beginning of the surface boiling of subcooled liquid nitrogen on the surface of the heater imitating the cylindrical object.

The research was carried out with subcooled liquid nitrogen flowing through annular ducts of 8.1, 10.1 and 12.1 mm inner diameter (d0, 100, 150 and 200 mm length (/), .8-3 mm radial clearance (S) and do/d* ratios ranging from 1.16 to 1.7, where do is the outer diameter of the annular duct.

The flow rate of liquid nitrogen was varied in different tests (at pressures from 1 to 10 atm) from 0.4 to 1.7 m/sec, subcooling at the entry was 3 ° to 20°. The range of heat flux densities from 6.103 to 6.104 kcal/m2.h. The experimental section is shown on figure 4. Thin-walled stainless steel tubes carrying large electric current served as heater elements. Such tube (fig. 4 position 1) served as an inner wall of an annular duct, the outer wall being the calibrated glass tube (fig. 4 position 2). Under the heat discharging surface six differential copper-constantan thermocouples were mounted. The electromotive force produced by the thermocouple corresponds to the temperature difference between the inner surface of the heating element and the liquid at the entry to the test section. There were also thermocouples for measuring the absolute values of temperatures at the entry and the exit of the annular duct.

The curves shown in figure 5 illustrate the effect of initial flow instability on heat transfer in annular canal when heat is discharged from the inner wall of the canal. The tests were made on 6 test sections. The increase of the heat transfer coefficient in

390

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the first part of the test section can be estimated by the coefficient

oc

c/LfS heat transfer coefficient in the first part of test section; oc heat transfer coefficient in the parts relatively far from the circular canal entry.

The test results are treated and presented in a table of values of Et for varying rela­tive length of canals and Reynolds numbers.

Table 1 THE VALUES OF EJ FOR CIRCULAR CANALS

\ Ude

Re \ 10 12 15 17 20 25

10 000 12 500 15 000 17 500 20 000

1.34 1.255 1.19 1.163 1.153

1.255 1.18 1.14 1.13 1.12

1.21 1.16 1.13 1.12 1.115

1.16 1.126 1.1 1.095 1.09

1.11 1.09 1.08 1.076 1.07

1.065 1.052 1.048 1.046 1.044

1 1 1 1 1

The equivalent diameter de is defined as follows: de = d0 — dt = 28. The values of the heat transfer coefficient obtained in short canals were recounted so as to get the values of coefficients for ratio l\de > 25 using data from table 1. Diagram points for each test section lay with some scatter on a straight line with slope 0.8. In the tests conducted the ratio d0/dt varied insignificantly (from 1.16 to 1.7) and observed stratification of test data was small. Therefore no correction was made for relative width of annular space, though this parameter had its effect. The test points settle themselves with ± 7 per cent scatter near the straight line corresponding to the equation:

Nu = 0.0208 Re 0 : 8 Pr° ' - 4 E ,

If heat flux densities are small, the heater surface temperature increases with the increase of the distance from the inlet of the canal. When the surface boiling appears, the temperature of the inner wall becomes constant along the length of the canal. This temperature exceeds by several degrees the temperature of liquid saturation. The surface boiling region moves towards the inlet of the canal with the increase of heat flux density.

The results of tests conducted under constant flow rates and inlet temperature conditions are shown on fig. 6. When heat flux density increases then at first the heat transfer coefficient decreases.The temperature decrease is more noticeable in sections nearer to the inlet of the canal. This effect can be explained by disturbance of hydro-dynamic and the thermal regimes of flow in the section where initial instability of flow takes place because of the development of surface boiling. The heat transfer coefficient grows essentially with further increase of heat flux density which is connected with developing of surface boiling. Nevertheless the surface boiling (formation and movement of vapour bubbles) is unnoticeable in visual observation up to the point marked " K " on the curve (fig. 6). In the intensive surface boiling area the tendency of decreasing curve slopes with growth of liquid nitrogen flow rate is noticed. It is

392

connected with the growing influence of forced convection on the development of surface boiling process.

The results given allow estimation of heat transfer intensity and temperature distribution along the annular cooling duct.

REFERENCES

[1] Proceedings of the Second International Cryogenic Engineering Conference, Brighton. U.K. (7-10 May 1968).

[2] Proceedings of the XVth All-Union Low Temperature Conference. Tbilisi (1968). [3] V.G. FASTOVSKY, U.V. PETROVSKY, A.E. RAVINSKY, Cryogenic Technics, "Energuia",

Moscow (1967).

393

COOLDOWN OF SYSTEMS ELEMENTS IN THE FORCED MOTION CONDITIONS OF THE COOLING AGENT

V.G. PRON'KO and G.M. LEONOVA Moscow (U.S.S.R.)

Refroidissement d'elements de systemes dans les conditions du deplacement force de l'agent de refroidissement

RESUME : Dans ce rapport on donne les resultats de recherches sur Vintensite d'echange de chaleur dans des tubes de faible diametre lors du refroidissement par azote liquide ou par helium gazeux froid.

On determine les modes d'ecoulement a deux phases par des methodes visuelles {cinema et photographie rapides). On propose des equations permettant de determiner Vintensite d'echange de chaleur dans les conditions de deplacement force de Vagent de refroidissement dans des tubes de faible diametre.

INTRODUCTION

Increase in size and weight of cooled elements of superconducting power devices necessitates careful analysis and special organisation of heat exchange while cooling them down from the temperature of surrounding medium to that of working condi­tions. Since the liquefaction of helium is rather expensive and heat of vaporisation is small in comparison with enthalpy, the preferable regimes of cooldown (taking into account the character of change of heat capacities of the materials with temperature) are:

1. Coodown with liquid nitrogen to the temperature 80-90 °K; 2. Cooldown with gaseous helium received from liquefier or refrigerator.

Cooldown is as a rule realized under conditions of forced circulation of cryogen. The paper concerns these cooldown regimes applied to widely used elements of power systems, that is to cylindrical ducts of small diameters.

T H E STRUCTURAL CHARACTERISTICS OF TWO-PHASE FLOW IN DUCTS OF SMALL DIAMETERS

The investigation of the first cooldown regime was carried out by forcing liquid nitrogen through the test section from the metal dewar into a glass one, where the separation of gas-liquid mixture took place: while liquid gathered at vessel bottom, the gas went to the gas meter through a special pipebend in the lid. The test sections were cryogenic tubes of small diameter with high-vacuum isolation. The tube was equipped with thermocouples for measuring temperature at several points along the tube. The thermocouple readings were recorded with electronic one-point potentio-metric recorders using a recorder chart motion rate of 9600 millimetres per hour. The test apparatus was also provided with several pressure points.

The tests were made in the following ranges: mass rate from 11 to 300kg/m2. s, inlet pressure from 1.1 to 2.2 bars, pressure drop from 0.1 to 1 bar, using three tubes of 4.2 m in length each and 1.7 x 0.3 mm, 2.7 x 0.2 mm and 3.9 x 0.25 mm diameters, made of nickel or German silver.

As long as the wall temperature of the duct is higher than the nucleate boiling limit (Leidenfrost point) it is impossible for liquid to exist near the wall and super­heated vapor separates it from the wall (the film boiling mode). After the wall tem-

395

perature reaches the nucleate boiling limit heat transfer intensifies because of transition to nucleate boiling. Film boiling takes 90 per cent of cooldown time. The investigations have shown that the appearance of liquid phase in the flow not only does not increase the heat exchange intensity but even diminishes it.

The authors made investigations of the flow structure using high-speed cinema­tography on the visible parts of the test ducts [1]. It was found that in a wide range of flow conditions (0.03 < x < l ) the reserved disperse-circular mode of flow in the tube (drops of dispersed liquid in vapor flow) persisted with slight variations of drop dia­meter. Heat transfers from the wall to vapor and from vapor to liquid drops. Direct measurements of flow temperatures as well as calculations made on the basis of the two-stage model of heat exchange show the presence of substantial superheating of vapor caused by poor development of interphase heat exchange surface.

DETERMINATION OF COOLDOWN TIME OF DUCTS

During experimental investigation of the process of cooling down the canal with gas the gaseous helium was used. The warm helium gas was served from a vessel through the reduction gear to a coil pipe submerged in liquid nitrogen, then to the test section and through the throttle valve to the gas meter. The tests cover the following range of parameters: inlet pressure from 5 to 10 bars, mass rate from 40 to 260 kg/m2. s, Reynolds number from 1.104 to 3.104.

The analysis of large amount of test data on cooldown of ducts with liquid and gas has allowed to establish the general laws of the cooldown process and to suggest a practical method for calculating the cooldown time based on these laws.

First it is necessary to note the fact that in most cases for different sections of tubing one can assume the constancy with time of the rate of change of the wall temperature. The comparative analysis of results obtained in [4, 5] and of test data shows that these cases correspond to the region where

^ > 20 (1) G9C9

where d duct diameter; a heat transfer coefficient; G^ mass flow rate of gas; Cg heat capacity of gas; Z coordinate.

The heat removal in this region is determined not so much by the heat transfer coefficient as by the heat capacity of the cooled duct and the possibility for a flow to carry the heat out of the considered section. In these conditions the change in the duct wall temperature T^ to within ±10 per cent can be described by the following equation:

T^-T^ V nd8pw'C„Z.

where 5, pw, Cw—wall thickness, density and specific heat respectively, TWo—initial wall temperature, Ten—entering gas temperature.

The temperature of the parts far off the duct entry remains constant for a long time. This result can be explained by the flow temperature in these parts being equal to the initial wall temperature.

396

The analysis of equations describing the heat exchange shows that the degree of ineffectiveness of the heat exchange surface is also a function of a complex parameter:

nd§pw-CwZ

One must keep in mind the fact that the analysis of the heat exchange equations was made assuming constancy of mass flow rate and heat capacity values as well as of heat transfer coefficient; the duct wall was assumed to be thin enough for its heat conductivity along the axis to be negligible. In order to make the obtained results more exact the treatment of experimental data was performed. The test results to within 20 percent can be approximated by the following two equations: In a region

0 < G g ° g T - < 0.8 T"~Tin = 1 nddpwCwZ

_G*^L > 0#8 ndbpwCwZ

T - T x w xin

T - T 1 Wo xm

= e x P r - l . l f - ^ ^ E o.sY] L V ndSp.C.Z ) \

(3)

In case of appreciable changes of parameters the calculations with equations (3) are valid for periods of time during which the parameter values may be considered constant.

The equation (2) can be used for calculating the time needed to reach the tempe­rature of transition from film to nucleate boiling while cooling of ducts with liquid nitrogen (Tw = 132°K). Mass flow rate of gas in these conditions is determined by means of isothermic flow formula using a time and position average for the gas temperature.

REFERENCES

[1] G.M. LEONOVA, V.G. PRON'KO, J.G. VINOKUR, The forms of vapor liquid flow in con­ditions of film boiling. Teploenergetica, 10, Moscow (1968).

[2] A.A. GUKHMAN, L.C. AKSELROD, V.G. PRON'KO, A.B. BULANOV, D.A. KAZENIN, G. M. LEONOVA, Some results of investigation of heat exchange between liquid and largely superheated wall. Teplophysica vysokikh temperatur, 6, 4 (1968).

[3] M.F. LAVERTY, W.H. ROSENOW, Paper A.S.M.E. (1965), NWA/HT-26. [41 J. BAUCHILLOUX, J.P. HUFFENUS, R.G.T. LEMAITRE, 54 (June 1966). [5] J. W.H. Cm, Adv. Cryogenic Engg, 10 (1965).

DISCUSSION

R.G. SCURLOCK (U.K.) — These results are in close agreement with the cooldown experiments at Southampton University. May I help to answer Dr. Norris' question about Tse. Figure 3 of our paper, p. 198 (Cooldown of Long Ducts) clearly shows the occurrence of Tse between 100 and 110 °K below which the cooling rate at each station increases very rapidly.

May I now ask a brief question ? Did you observe pressure oscillations during cooldown with the pressure meters included with your test apparatus ? If so, how large were they ?

V. G. PRON'KO — No pressure oscillations were observed.

397

MEASUREMENTS OF NUCLEATE BOILING HEAT TRANFER TO LIQUID HELIUM FROM A SIMULATED

SUPERCONDUCTOR ARRAY WITH COOLING CHANNELS

S. KUGLER and I.C. CROSSLEY The British Oxygen Company Limited, Research and Development, London {United Kingdom)

Mesures du transfert de chaleur dun appareil supraconducteur simule avec rainures de refroidissement a l'helium liquide en ebullition par nucleation

RESUME : Le fonctionnement satisfaisant des appareils supraconducteurs stabilises depend du transfert efficace de la chaleur.

Le CERN de Geneve a conclu un contrat avec la societe ou travaillent les auteurs pour la determination, a trois temperatures du bain d'helium liquide, des differences de temperature et des flux de chaleur correspondants, ainsi que du point de transition de Vebullition par nucleation a Vebullition en film, dans un modele de la bobine supraconductrice proposee pour la chambre a bulles BEBC.

Un modele a ete construit representant des sections de quatre spires d'un enroulement a galettes doubles. Chaque supraconducteur a ete simule par une plaque en cuivre avec element chauffant en acier inoxydable. Les differences de temperature entre chaque plaque et le bain d'helium liquide ont ete mesurees avec plusieurs thermocouples. Les signaux ont ete amplifies par un microvoltmetre amplifiant et etaient soit lus directement soit enregistres au moyen d'un galvanometre a miroir reflechissant des rayons ultraviolets. Le courant pour Velement chauffant a ete fourni par une source de courant continu tres stable et pouvait etre enregistre.

La precision de tous les instruments etait superieure a ±0 ,01°K et toutes les mesures etaient reproductibles.

Nous avons trouve que dans la configuration proposee le point de transition de Vebullition par nucleation a Vebullition en film se produit pour un flux de chaleur de 0,4 watt/cm2 environ, et que la valeur de ce flux limite augmente avec Vabaissement de la temperature du bain. Lors de la transition la difference entre une plaque en cuivre et le bain etait de Vordre de 0,4 °K. / / y avait une hysteresis considerable, la transition inverse se produisant a la reduction du flux de chaleur a une valeur de 0,25 watt/cm2 environ.

Veffet des bulles d'helium produites au dessous de la section experimentale a ete etudie, ainsi que diverses configurations des rainures de refroidissement.

1. INTRODUCTION

The CERN organisation in Geneva is at present engaged in the design and con­struction of a large bubble chamber, incorporating a superconducting split solenoid magnet of some 4.7 m diameter. This magnet will be of horizontal pancake con­struction using flat strip conductors separated by load bearing cooling strips.

The magnet is intended to be fully stable and consequently it is necessary to know the maximum heat flux which can safely be dissipated in such a magnet by nucleate boiling of helium.

Measurements on the heat transfer to boiling liquid helium, in particular in narrow vertical channels, have been made [2, 3, 4, 5, 6] but these have usually been on a small scale and on simple geometries. It is therefore difficult to predict accurately the performance of an actual large magnet configuration. Furthermore in an actual magnet the problem is complicated by the possibility that considerable volumes of gaseous helium generated in lower windings may bubble through a section requiring cooling.

To determine safe operating limits for the CERN magnet and to find out the relationship between the conductor temperature rise, and the heat flux dissipation from the conductor, a model simulating a section of the magnet winding, was built and tested in liquid helium. The effect of several parameters on the heat transfer

399

performance of the conductor was investigated, including different materials and geometries for the cooling plates which separate the individual turns of the winding in any one pancake coil, and variations in bath temperature, etc.

2. EXPERIMENTAL RIG

The heat transfer test rig was designed to simulate heat transfer and boiling of helium from the superconducting magnet windings. Figure 1 is an exploded view of the test rig showing that a total of four short lengths of winding were simulated, these being positioned in relation to each other exactly as in the magnet, and provided with cooling plates and vertical spacers which in all basic dimensions duplicated the first alternative of the proposed magnet designs.

• THERMOCOUPLE POSITIONS (BURIED IN 'CONDUCTOR: BUT SHOWN ON COOLING PLATE FOR CLARITY)

STAINLESS STEEL-

TYPE M COOLING PLATE

•Ol M M COPPER SHIM

INSULATION^ J ^ '.5 M M

3-5 M M

SIMULATED CONDUCTOR

.0625 MM HEATER

8 MM

.5 M M INSULATION

. H FOUR CT\COOUNG [jy Assrs

FIG. IC TYPICAL ARRANGEMENT SQUARE STUDS

FIG. ID TYPICAL ARRANGEMENT

VERTICAL STUDS

FIG. IE TYPICAL ARRANGEMENT

DIAMOND STUDS DIAGRAMMATIC

END VIEW OF ASSY

Fig. 1 — Views of model.

Although consisting of four conductor elements the rig was conceived as two separate models capable of being tested individually, each representing a vertical section of the magnet. Each model consisted of two conductor elements one above the other together with a well shielded helium vaporising hearer below the bottom conductor to simulate the effect of vapour rising from still lower windings in the

400

magnet. The two halves of the rig were identical except for the choice of cooling plate. Plates were made either of plastic (Type I) or metal (Type M), viz: 1. Square stud plate of glass fibre reinforced epoxy resin — Plate I; 2. Square stud plate of copper — Plate M l ; 3. Vertical channel plate of copper — Plate M2; 4. Diamond stud plate of copper — Plate M3; 5. Round stud plate of copper — Plate M4.

The plates were dimensioned 10.0 cm x 8.8 cm and the arrangements of studs are shown on figures 1C, ID and IE for Type I, Ml , M2 and M3. The round studs of Type M4 were 0.175 cm high, 0.81 cm diameter, with 42 equally spaced over the plate, and two solid vertical strips 0.62 cm wide x 8.8 cm long running one along each edge of the plate.

The copper plates were all of O.F.H.C. copper and incorporated a stainless steel backing. Each simulated conductor consisted of a strip of high conductivity copper, of resistance ratio 340 instrumented with eight gold + 0.03 per cent atomic iron vs chromel thermocouples, as developed by Berman [7 & 8]. A heating pack, consisting of a copper shim, a thin mylar insulator, a stainless steel strip heater and a backing of thick mylar insulation was clamped tightly against the copper rear face. Silicone grease was used to ensure good contact between the copper shim and the simulated conductor.

In order to obtain a practical result the surface finish of the copper was left in the "as manufactured" condition. No precautions were taken to clean or polish the surface, which was left exposed to the laboratory atmosphere for a number of days.

3. INSTRUMENTATION

The thermocouples were arranged to read the temperature difference between the copper and the helium bath. The hot junctions were made and attached to the copper using indium buried in holes sunk in the heated face of the simulated conductor—their approximate positions are shown on figure 1. Low 'noise' thermocouple switches were used to select the thermocouple to be read—the output was read on a sensitive micro-voltmeter and could be recorded on a U/V galvanometer. It should be noted that the cooling plates could be changed without in anyway interfering with the thermocouples.

Considerable testing of the instrumentation was carried out at a bath temperature of 4.2 °K to prove the equipment and eliminate extraneous noise. Ultimately an accuracy of measurement, and repeatability of better than 0.03 °K was achieved.

The heaters were energised using superconducting magnet power packs, with very stable D. C. output. Voltage taps were taken to the heater terminals in the helium.

4. CRYOGENICS

The experiment was carried out in a standard stainless steel helium cryostat with nitrogen shielding. The model was suspended from the brass top plate by 2 stainless steel rods—radiation shields were provided. Carbon resistors ,were used for helium level indication.

The bath temperature was controlled manually via the cryostat pressure by means of a pneumatically actuated valve, operating on the boil-off gas as it left the cryostat, and before it passed through an air heater to the balloon storage system.

401

5. TEST PROCEDURE

Test as proposed by CERN [1] were carried out on cooling plates Type I at a bath temperature of 4.4°K, and on Ml at bath temperatures of 4.4, 4.5, and 4.6°K. Test on plates, Type M2, M3, and M4 were only performed at 4.5°K.

0 5

4 45 4 5 4 55 BATH TEMPERATURE *K

Fig. 2 — Effect of bath temperature.

Test were of two forms:

(a) Measurements of temperature differences at various heater power levels, both increasing and decreasing, the readings being taken on the microvoltmeter.

(b) Measurements of temperature differences with heater power increasing or decreasing steadily, with additional helium vapour generated in a lower pack passing through the test section at various steady rates. For these tests both thermocouple output and heater current were recorded simultaneously by means of the U/Vgalvano-meter.

Originally, all thermocouples were read; in later tests only selected, representative instruments were read, as the temperature variations between individual thermocouples were found to be very small.

0-5

2,

i O S * -

0 DIAMOND STUDS O SQUARE STUDS X VERTICAL CHANNELS

BREAK AWAY

RECOVERY

LOWER HEATER POWER INPUT-WATTS

Fig. 3 — Effect of vapour blanketing.

402

6. RESULTS

6.1. Effect of Measurement Position. — On each conductor there were slight temperature variations between the thermocouples, which reproduced accurately during repeat tests, and which in no case exceeded 20 % of the average temperature difference between conductor and helium bath.

These variations did not in any obvious way correlate with the height of the thermocouple in the bath, or position on the simulated conductor. Although there was some slight evidence that thermocouples at the top of the plates gave somewhat higher temperatures than the others. However, when the cooling plates were changed the temperature pattern changed also.

It was concluded that the measurements represented small but real temperature variations between different parts of the conductor, caused by the cooling plate geometry.

On the graphs individual thermocouple measurements have not been shown, but merely the highest and lowest thermocouple readings at any given power.

6.2. Effect of Cooling Plate Material. — In the original series of tests at 4.4 °K geometrically identical cooling plates made in an insulating material (Type I) and metal (Type Ml) were tested. The insulating plates proved to be inferior both in the upper and lower position, as illustrated in table I.

Table 1

Break-away flux watts/cm2

Type Ml

Upper Lower

0.434 0.458

Type I

Upper Lower

0.375 0.342

It should be noted that in this paper the heat fluxes have all been referred to the effective heat transfer area in contact with liquid helium, and not to the total geome­trical surface area of the simulated conductor.

In view of these results no further tests were performed on Type I plates.

6.3. Effect of Vertical Position.—The boiling curves of all the geometries tested were influenced by whether the conductor under test occupied the upper or lower position in the rig. Figures 4, 5, 6 and 7 show for each geometry the boiling curves for the upper and lower conductors, and the various critical flux values are given in table 2.

Table 2

Break-away flux Recovery flux watts/cm2 watts/cm2

Upper Lower Upper Lower

376 .412 .254 .242 449 .379 .301 .339 470 .468 .337 .314 506 .517 .347 .363

Type M l Type M 2 Type M 3 Type M 4

403

O S

0 - 2

0

U z QC

LL Q

Q! 2 LU P- mm —

&z^~~

r\

- 1 .

I i

1 i

— i !

HEAT j

**~1^~~

w

| s r = - = ^

FLUX

—-,-4!_ --*-

K 1

9 1

.*? >"

- -- i ""*"

WATTS/CM2.

m**~

LOW

ER

. J ' i

V

Fig. 4 — Heat flux square studs.

2 -0

1 - O

0 -5

0 -2

0

LU U z LU ff

u_ LL Q a: 5 UJ H-

- - —

HEAT

1 3

. , - ,

FLUX

1 1 1

t 1

1 !-.J. WATTS / O

| ■ 4 T 1 1 i

. - • ' ' -

+*

3 k

Fig. 5 — Heat flux vertical channels.

o-s

0-3

0-2

o LU U z LU a. LU LL LL Q Q. 5 t-

HEAT FLUX

«— *' - 1 -c

3 t = 1 • 1

1 1

_i_^«

r

WATTS/CN

r s ^ ^ ^

_ m_^g0^*

i*.

-~-~~'

" " i

l

1 si 14

-«<^i

1

■

Fig. 6 — Heat flux diamond studs.

404

A study of the figures in table 2 indicates that no straightforward predictions of the effect of vertical position are possible; the cooling plate geometry appears to be the main parameter in deciding whether the highest critical fluxes are obtained in the upper or lower position of the rig.

0 2 0 25 0 3 0 35 0-4

Fig. 7 — Heat flux round studs.

0 5 5

6.4. Effect of Cooling Plate Geometry.—This was investigated at a bath temperature of 4.5°K for Type Ml , M2, M3, and M4 cooling plates. The results, for both upper and lower packs are plotted in figures 4, 5, 6 and 7.

In each case it is found that the curve of temperature difference vs heat flux is reversible in the nucleate boiling range. At break-away flux there is the expected increase in temperature difference. On decreasing the heat flux there is considerable hysteresis, with an extensive region of film boiling, until recovery occurs.

It is interesting to tabulate the values of break-away and recovery flux, which show that the round stud pattern, Type M4 has the best performance.

Two points should be borne in mind: (a) The effective heat transfer area in contact with the liquid helium varies as

follows: Type Ml — 58.25 cm2

Type M2 — 42.5 cm2

Type M3 — 62.5 Type M4 — 56 cm

(b) The actual power dissipated from a conductor model at transition is of consi­derable significance. On this basis the diamond and round stud pattern are substan­tially better than the other geometries.

Table 3

Type Ml Type M2 Type M3 TyjMj M4

Break

Upper

21.9 19.1 29.4 28.4

-away power watts

Lower

24.0 16.1 29.2 29.0

Recovery power watts

Upper

14.8 12.8 21.0 19.4

Lower

14.1 14.4 19.6 20.4

405

6.5. Effect of Bath Temperature.—This was investigated on Type Ml cooling plates and the effect on break-a way and recovery flux for both upper and lower packs is plotted on figure 2. It is seen that all values increase as the bath temperature is lowered. This is in general agreement with Kutateladze's correlation [9] for the variation of critical flux with pressure.

6.6. Effect of Added Helium Vapour.—This was investigated at 4.5 °K for Type Ml , M2, M3 and M4 cooling plates, and the results are plotted in figure 3.

It had been expected that the helium vapour would reduce transition fluxes. In fact, the results show that in general the effect of helium bubbles rising from a lower source is to reduce the break-away flux while leaving the recovery flux substan­tially unaffected. At sufficiently high bubble input hysteresis is largely eliminated, and break-away and recovery fluxes coincide.

It has thus been shown that the region of hysteresis is reduced in extent by distur­bances to the flow and gross disturbances eliminate it altogether.

For the purpose of designing fully stabilised superconducting magnets of the type considered in this paper, it would therefore not be safe to design for heat fluxes higher than those defined by the recovery condition which, however, is unaffected by the excess helium vapour.

6.7. Temperature Rises.—Analysis of all the results obtained indicates that the temperature rise experienced by the conductor during nucleate boiling are generally in the range 0.1 °K to 0.5 °K with a maximum of 0.8 °K for the Type M3 cooling plate.

For film boiling the temperature differences were within the range 1.8°K to 5°K in the region between break-away and break-back.

7. COMPARISON WITH PREVIOUS WORK

Work previously reported has been on plain channels and is thus strictly speaking only comparable with the Type M2 cooling plates. Using the equation [4] for channels heated from one face the critical quality at break-away for the Type M2 plate was:

(a) Upper pack q = 0.191 (b) Lower pack q = 0.156

Wilson [4] states that a critical quality of 0.3 gives a reasonable indication of the onset of break-away in narrow channels. This is, however, true only for a limited range of conditions, when the channel width is small compared with the height, and the critical quality drops with increasing channel width.

Extrapolating Wilson's experimental points gives a value of q = 0.183 for conditions equivalent to ours.

This compares well with the values obtained despite the different bath temperature and assymetry of the model.

ACKNOWLEDGEMENTS

The work described here was carried out by BOC under contract to CERN, and the authors wish to thank the CERN organisation, as well as BOC for permission to publish this paper. Thanks are also due to our colleagues at BOC, Morden, Mr. M.N. Wilson of R.H.E.L., and to Dr. F. Wittgenstein, Mr. A. Herve, and Dr. E. U. Haebel of CERN for considerable helpful advice.

406

REFERENCES

[1] Contract Specification for Checking of the Heat Transfer in Cooling Channels of the BEBC Magnet, (D.Ph. II/BEBC/FW/mg) (22nd February, 1968).

[2] S.G. SYDORIAK and T.R. ROBERTS, "Study of Boiling in Short Narrow Channels and its Application to Design of Magnets Cooled by Liquid Hydrogen and Nitrogen", / . Appl. Phys., U.S.A. (February 1957).

[3] S. G. SYDORIAK, and T. R. ROBERTS, " Critical Nucleate Boiling of Liquid Helium in a Simu­lated Wire Wound Magnet", Annex 1966-5 Bull. LI. R. pp. 115-123, Commission I, Boulder.

[4] M.N. WILSON, "Heat Transfer to Boiling Liquid Helium in Narrow Vertical Channels", Annex 1966-5 Bull. I.I.R., pp. 109-114, Com. I, Boulder.

[5] J. C. BOISSON, J. J. THIBAULT, J. ROUSSEL and E. FADDI, " Boiling Heat Transfer and Peak Nucleate Boiling Flux in Liquid Helium", Adv. Cryog. Engg, 13 (August. 1967).

[6] S. LEHONGRE, J.C. BOISSON, C. JOHANNES and A. DE LA HARPE, "Critical Nucleate Boiling of Liquid Helium in Narrow Tubes and Annuli", Proc. I.C.E.C.2. (May 1968).

[7] R. BERMAN, andD.J. HUNTLEY, "Dilute Gold-Iron Alloys as Thermocouple Material for Low Temperature Heat Conductivity Measurements", Cryogenics (June 1963).

[8] R. BERMAN, J.C.F. BROCK and D.J. HUNTLEY, "Properties of Gold +0.03% (at) Iron Thermoelements between 1 and 300 °K and Behaviour in a Magnetic Fiel", Cryogenics (Aug. 1964).

[9] S.S. KUTATELADZE, "Fundamentals of Heat Transfer", E. Arnold (1963).

REMARQUE

E. CARBONELL (France) — Nous avons entrepris exactement la meme etude aussi pour le compte du CERN au Centre d'Etudes Cryogeniques de L'Air Liquide (Sassenage, France).

Nous avons observe les memes resultats generaux sauf en ce qui concerne les valeurs de flux de chaleur critique qui sont environ 40%plus petites que votre "break away", mais cependant egales a vos "recovery values". De plus, nous n'avons pas constate l'effet d'hysteresis montre dans les figures 4 a 6.

Pour tenter de comprendre les differences entre vos resultats et les notres il faut noter que la difference de temperature en fonction du flux de chaleur a ete etudiee a flux de chaleur croissant puis decroissant, conformement au programme represents sur la figure 1. Dans ces conditions nous n'avons pas observe d'hysteresis.

r ->—■ ■—> 1 j > — - i

|—> 1 ' 9 1

Temps.

Fig. 1 407

La relation entre le flux de chaleur et I'ecart de temperature est representee en figure 2.

Ebullition

vecov/ev

EbuLLi tion en PiLr

b-Yeck kck 6oaci

OScilla6/'onS

FLuoc ale choLeu-r

Fig. 2

Vous n'avez pas detecte d'oscillations dans la zone de transition (fig. 2) car vous avez utilise un microvoltmetre. Nous avons mis en evidence le phenomene classique des oscillations grace a notre chaine de mesures (amplificateur galvanometrique SEFRAM) specialement adaptee a leur frequence (quelques hertz).

408

SOME FUNDAMENTAL PROBLEMS WITH SUPERCONDUCTING ENERGY STORAGES

F. IRIE and K. YAMAFUJI Electronic Engineering Department, Kyushu University, Fukuoka {Japan)

Quelques problemes fondamentaux du stockage de l'energie supraconductrice

RESUME : L'energie dissipee dans une bobine de stockage pendant un cycle de charge et decharge d'energie est traitee thioriquement d'une maniere phenomenologique, en mime temps que son rapport avec les pertes des machines qui y sont liees. Pour le calcul de la perte dans la bobine on adopte une forme generate proposee pas les deux auteurs pour la force d'ancrage.

Si la « perte relative » est definie par Venergie de perte divisee par Venergie emmagasinee, la perte relative par cycle de charge et decharge presente un maximum en fonction du champ maximal emmagasine. On demontre aussi que la perte relative est inversement proportionnelle au rayon de Vespace interieur de la bobine. Quant aux pertes des machines, on analyse les pertes subies pendant un chargement a tension constante et une decharge a puissance constante. Les pertes dues a la resistance de Vinduit et des fils conducteurs, qui constituent la plus grande partie des machines, sont donnees en fonction du rendement nominal des machines.

On conclut que pour les bobines tres superieures a 150 MWH la perte relative de la bobine divisee par le rendement du refrigerant est negligeable par rapport aux autres pertes, mais ce n'est pas le cas pour les bobines plus petites.

An energy storage device is one of the promising applications of superconducting coils which have come in the stage of industrial application recently. A superconducting coil has already been used as an element of energy storage in a pulsed system [1], and now there appears some possibility of its use as an energy storage device in electrical power systems. In such a device it must be noted that some losses arise in the courses of the charging and discharging, while no loss arises when energy is kept storing. Materials used in such coils are nonideal type-II superconductors with high magnetic critical fields, such as Nb-Zr, Nb-Ti, and Nb3Sn. The magnetization of such material shows large hysteresis which is a part of the losses in a supercon­ducting coil. Such a loss is dominant when the current is not changed very rapidly. If the current is varied very rapidly, a flux-flow loss must be added to the above kind of loss. In a superconducting coil, however, the current is normally varied slowly, and hence it seems enough to regard the hysteresis loss as the main loss in a superconducting coil.

In this paper, the hysteresis loss in a superconducting coil is discussed quantita­tively, together with the losses in room-temperature devices connected to the coil. The problem of flux jumps which lead to a quenching of superstate is also very important. These jumps are closely related to the heat produced in the coil due to hysteresis loss [2], but this is not considered in this paper.

LOSSES IN A SUPERCONDUCTING COIL

Fundamental equations which describe the flux motion were introduced by the authors [3] for a quantitative treatment of the losses in a superconducting coil [4]. First, the force balance equation is given by

6B2/8 ndr+tacBY/8 n + A,pvB/8 n = 0 , (1)

Second, the equation of continuity of flux density is given by

8B/6f+k6(rvB)/r8r = 0 . (2)

411

In these equations, B denotes the absolute value of the flux density, r the radial variable in a cylindrical coordinate system, v the absolute value of the velocity of fluxoids, P the viscosity coefficient, and X = Sign(- <)B2/<)r) represents the direction of flux motion. The first term of equation (1) represents the Lorentz force density to drive fluxoids in a superconducting coil. The second term is the so called "pinning force density " which is characterized by two parameters, a and y. The third term is the viscous drag force density for flux-flow. The parameters, oc, p, and y can be regarded as substantial constants under a given temperature. In principle, any real number can be chosen as the value of y, but observed magnetization curves in any nonideal type-II superconductor seems described well by such a value of y as being in the range of 0 < y < 1. In addition, most of the former models for the pinning force density are included in the present expression as special cases.

If we regard each layer of the coil as a thin hollow cylinder, the static distribution of flux density when an external magnetic field is applied parallel with the cylinder axis can be obtained by taking a limit of v -► 0 in equation (1). The result for the outer side of the kth layer is given by

^B0VY = M B i ? ) 2 ^ - ( 2 - Y ) a ( r - r<{>)/2 , (3)

where r<£> denotes the radius of the outer surface of the layer, Xok and B<£> (1) denote the values of X and B at r<£>, respectively, and the subscript " 0 " and the super­script " / " represent the value of the corresponding quantity in the outer side and the surfaces of the layer, respectively. A similar equation is derived for the inner side of the layer, but with a replacement of the subscript "o " by "/". Typical examples of the distribution of the flux density in the kth layer are shown in figure 1, where the value of X is — 1 when the fluxoids are penetrated into the layer with increasing B<£> and +1 when the fluxoids are expelled with decreasing B££\

Since the magnetic field is not varied very rapidly in a storage coil, the pinning loss is more dominant than the flux-flow loss, as mentioned before. The pinning loss arises when the flux inside the sample is redistributed according to the variation of the applied field. Since the density of the pinning loss power is given by vaBy/87i the loss energy in the kth layer can be obtained by

»-.-?]>(! '};}

+ ik

CW

J'S? rvBYdr, , (4)

where / denotes the length of the coil, r\p and r(0f denote the innermost points

of the redistribution made by the flux invasion from inner outer surfaces of the layer, respectively, as shown in figure 1.

The variation rate of the hysteresis loss of the whole coil with a variation of current is derived by changing B into the current through the coil. Its normalized form is shown as

6Wp/8i = [(2-Y)/(5-y)] U , (0 -D/ 2 ( i ) / ( 2R 0 -D) ] i 1 " ' (5) where

M

/ i ( 0 = [ (5-Y) /2M 5 -^ £ [X\{\ kA"< {}<£> i-X£> 4e>)2 -

-W(k-l)*-HtiiU-X%i%)2l, (6) M

/2(i) = [(5-y)/2M5-'] £ [_Wk4-"<(2klM-l)(Wi-WiV)2-k=l

- W (k -1)4"* {2(fe -1 ) /M - 1 } (KV i-W <?*')2] • (7)

412

( a )

Ha

( b) ( c ) ( d )

Fig. 1 — Flux distribution in a layer of a coil for various states of magnetization.

In these equations, R0, D, and M are outer radius, thickness of winding and total number of layers of the given coil, respectively. Small letters wp and i are the norma­lized quantities of Wp = £Wpik and current I with V 0 H^/8TC and Ifl, respectively, where V0 denotes the volume of the superconductor in the coil, Ha denotes the field at which the penetration of flux into the layer is just completed and is defined by

H f l E E [ ( 2 - y ) a d / 4 ] 1 / ( 2 - ^

with d denoting the thickness of the layer, and la is the specific value of current which produces Hfl,

When 0 ^ i ^ 1 in the initial magnetization process, it can be put WV = 1, Xofk = - 1 and i,(jj} = ioV = 0 for every layer. Then equations (6) and (7) can be calculated by approximating the summation by integrations to lead to

7.(0 = i2 for 0 < i < 1. (8)

The function f2 (i) is not important because D is to be made much smaller than R0 in a desirable coil, as shown later. When i exceeds 1, fluxoids penetrate completely in some layers as shown in figure lc. In this case/! (/) is given by

/ i ( 0 = [ l + ( 5 - Y ) x2^{x-(x2^-l)1K2-^}2dx-]f for i ^ 1. (9)

When the value of / is decreased from the maximum value im, then fi (/) takes several forms according to the value of im. Here we shall show, for example, the form of fx (/) in case of im ^ /pM, where ipM denotes a limiting value of /, beyond which the flux distribution in the outermost layer comes to have a single valley in the second cycle of the magnetization process:

/ i (0 = C2' _ r>(5-7) / (2-Y) ( 1 - 2 - l / ( 2 - y K 2

I: + ( 5 - y ) l / ( 2 - Y ) x 2 - Y { x - ( x 2 - y - l ) 1 / ( 2 - Y > } 2 d x ] r - 3

for i ^ / pM (10)

The value of wp can be obtained by the integiation of equation (5) by the use of these forms of/x (/), and numerically calculated curves are shown in figure 2 for y = 0, 0.5 and 1, respectively. It is to be noted in this figure that the curves are different between the cases for magnetization and demagnetization.

413

LOSSES IN AN ENERGY STORAGE COIL

To discuss the efficiency of an energy storage coil, it seems convenient to define the "relative loss" of the coil, £L, by a ratio of loss energy during charging and discharging processes to the stored energy at the final state of the charging process.

2 10

wp

101

o 10

-1 10

io"1 io° io1 i 102

Fig. 2 — The accumulated amount of normalized loss energy density generated during the increase of normalized current from 0 to /, or during the decrease from 102 to i.

An alternative definition of the relative loss is given by the ratio of t he loss density to the stored energy density multiplied by the volume ratio of the superconducting material to the inner space of the coil, when the volume of the superconductor is small compared with that of the whole space occupied by the coil. Calculated values of such a relative loss are shown in figure 3 as a function of Hm/Ha, where Hm is the maximum field in the coil which is attained at the final state of charging process It is to be emphasized that each curve has a maximum, and hence the relative loss is remarkably small for large values of Hm. This fact is very convenient because the value of Hm should be taken as large as possible to obtain a large capacity for energy storing.

An approximate formula of ^L for this region is shown as

£L = Ax(10Hfl/HJr, (11)

where x is defined by x = 2Mde/R0 with de denoting the effective thickness of a single layer of the coil. The quantities A and T are functions of y and change with the history of magnetization of the coil as

A = 0.25/(2.2 - y ) and r = 1.3-0.45y

II 1

1 III

L-r

L

L L r

r f u p £

L—J i

■ ■ i u i | 1—i i i 11 HI i i i 111 I M

—* A

/ \ // ~ic''' j

jy , - *' "N J

/>£^^^" -y frr ^^ 1

Y f / \ -I // / — r = i X- J ! iff r= 0.5 x-1

.7 !! — rz ° M • i ! — > initial U

y j J magnet izat ion ij

y / / ^"^ second J / ij m«agnctiz4tion J

/ U 1 ' ■ ' > '' ' ■ ■ ■ ■ ■ i ll ■ '1

414

for an initial charge-discharge cycle,

A = 0 . 1 5 / ( 2 - Y ) and r = 1 .3-0 .50y

for a second cycle and later ones.

Now it may be instructive to show typical examples for the orders of magnitude of these quantities. The values of Hw and Hfl for standard commercial superconducting wires or ribbons are about 100 kOe and 3.5 kOe, respectively, and Mde for this case is about 5 mm. It is to be noted that Mde is a net total thickness of superconductor, not including the portions of normal metal. In addition, the value of y for such materials is 0 .5~ 1. Then the value of £L becomes 0.03/Ro~0.06/Ro; that is ,(5—10) x 10 ~5 for R0 = 6 m. But it is to be noticed that the important quantity is not £L

but is a relative loss, £L, converted into the value at room temperature. The value of £L f ° r t n e above case is given by 5~10% if it is taken into account that the effi­ciency is about 1/1000 for a large refrigerator.

o 1 0

A fC

-1 1 0

10° 101 102 H m / H a 103

Fig. 3 — Relative loss in a coil of 100 layers for the second or later cycle of magnetization and demagnetization.

LOSSES IN A SYSTEM OF ENERGY STORAGE

To consider the total loss seen form the power line connected to the storage de­vice, many kinds of losses other than the abovementioned loss must be taken into account. Although the amount of these losses depends on the kinds of input and out­put devices, here a motor-generator-type AC-DC converter is chosen as an input as well as an output device for example, where an AC machine is connected to a power line. The armature resistance loss may then predominate over other losses in a DC machine. When the coil is charged by a DC generator with a constant voltage V, the current i increases following the usual equation:

L(dildt) + Ri + e = V , (12)

where L is the inductance of the coil, R is the resistance of armature and lead wire, and e is the additional voltage induced in the coil by the invasion of the flux into the superconductor. But e in this equation can be neglected because e is clearly written as the form of/( / ) di/dt, and also because the induction voltage from the flux in the inner space of the coil must be much larger than that from the flux in the super-

415

T 1 1 I I I I I I 1 1 1 I I I I II 1 1 1 I I | I U

conductor. The relative loss for charging, £RC, which is the above loss divided by the stored energy, under the assumption of t < L/R is given by

£RC = 2*f/3T, (13)

where T ( = L/R) is the time constant, and tc is the charging time. The power discharged form the charged coil can be controlled by the adjustment

of field current. The value of current / in a discharging process to make the discharging power Po constant is given by a solution of the following equation:

Li(di/dO + R*"2 + Po = 0 , (14)

if the additional e.m.f. from the flux invasion into the superconductor is neglected. The solution of this equation is given by

i2 = £ e x p ( - 2 t / T ) - ( P 0 / R ) [ l - e x p ( - 2 * / t ) ]

c i 2 - 2 P 0 l / L , (15)

where im is the maximum current stored in the coil. This solution shows that the condition of constant output power can be satisfied only for finite period of time ta, which is approximately given by L/m

2/2P0. The field current /> which holds the conditions of constant output power becomes

iF oc Po/oi ^ P0/co \il - 2 P 0 */L]*. (16)

Such a value of /F can be retained automatically by the application of a feedback system.

The relative loss of this kind for discharging period, C,na, is given by

CRd = £ R / 2 P o - (17)

Since T is given by x = 2P0fa//OT2R, the relative loss for one cycle of charging and

discharging, £R , can be written as

CR S CRC + CM = 0* R/Po) 0c/3 td + i ) . (18) Here im and P0 are to be chosen as the rated current and rated power of the DC machine, respectively. The part of the factor z' R/Po that comes from the armature resistance may be known from the efficiency of the DC machine. For a large DC machine this value is about 3%. But the total value of the factor may be somewhat larger than this value, considering that the loss of lead wire which is generated in the dewar must be divided by the efficiency of the refrigerator. As remaining losses, there are field loss, iron loss, mechanical loss of DC machines, and also the losses of AC machines, though these losses may not exceed ^R .

Thus the resultant relative loss from machines and lead wires may amount roughly to 5~10%. This value does not vary much with the variation of the rated power, while the value of ^ is inversely proportional to the radius of the coil for a given stored field. For the coils larger than the aforementioned one (R0 = 6 m), ^ Pre~ dominates. Thus it may be instructive to know what the capacity of the storage coil having this critical value of the radius is. In principle any magnitude of capacity can be obtained for a given ££ • But the capacity is determined uniquely, if an optimum economical design to minimize the volume of superconductor in a toroidal coil for a given storing energy is desired. For a toroidal coil with the centre-line radius of p cm and winding radius of R0 cm, the total stored energy is given by

Ws = ( n / 2 ) B ^ p 3 ( l - x ) 2 [ l - ( l - x 2 ) * ] x K r 7 ( J ) , (19)

416

where BM is the maximum field in the coil and x is defined by x = R0/p. The volume of superconductor Q in this coil is given by

Q = 4n2RMde. (20)

An economically optimum value of x can be obtained as x = 0.2 by the mini­mization of equation (20) under the condition of constant Ws and BM in equation (19). Then the capacity of the coil of R0 = 6 m, whose relative loss was shown to be 5~10%, is obtained as

Ws = 150 M W H .

Thus it can be concluded that the loss of the storage coil must be taken into account for the storing capacity less than about 150 MWH, while the coil loss is covered by other losses for larger coils.

ACKNOWLEDGEMENTS

The authors are indebted to N. Sakamoto for numerical calculations and helpful discussions.

REFERENCES

[1] E.J. LUCAS and Z. J.J. STEKLY, AVCO EVERETT Research Lab. AMP 246. [2] K. YAMAFUJI, M. TAKEO, J. CHIKABA, N. YANO and F. IRIE, / . Phys. Soc. Japan, 26 in

press; J. CHIKABA, F. IRIE and K. YAMAFUJI, Physics Letters, 27 A (1968) 407. [3] F. IRIE and K. YAMAFUJI, / . Phys. Soc. Japan, 23 (1967) 255. [4] F. IRIE and K. YAMAFUJI, Physics Letters, 24 A (1967) 30; F. IRIE and K. YAMAFUJI, Proc.

1st Int. Cryog. Eng. Conf. (1967) p. 177; F. IRIE, K. YAMAFUJI and N. SAKAMOTO, J.Inst. Elec. Eng. Japan, to be published.

417

ETUDE ET REALISATION D'UN ENROULEMENT SUPRACONDUCTEUR DESTINE AU STOCKAGE

DE L'ENERGIE

B. GIRARD et M. SAUZADE Institut a"Electronique Fondamentale, Faculte des Sciences, Orsay {France)

Study and construction of an energy storing superonducting coil

SUMMARY: The paper describes a superconducting coil used to store energy and restore it in about one millisecond. Specially studied is the problem of current interruption.

The coil was made from Nb- Ti stranded cable stabilized with pure aluminium. Its maximum energy content is approximately equal to 20 kilojoules and it is closely coupled with an aluminium ribbon coil which takes the place of the secondary coil of a non-iron transformer. The super­conducting winding has ten times as many turns as the secondary winding. In this way it is possible to discharge the stored energy at a lower impedance {less than \Q) or at a higher impedance {about \0Q) by connecting the load to the secondary or the superconducting circuit.

The current is rapidly discharged by opening the circuit via a vacuum switch using on artificial zero-current injection system. This switch can break 1 000 A, without failure, with overvoltages of 4 or 5,000 V. It is capable of withstanding even higher voltages. The switching time is only a few micro-seconds.

The stabilized superconducting coil withstood many charge rates of a few tenths of seconds duration and discharge rates of 3 or 4 milli-seconds. Transfer energy rating is about 65%. At the moment, a part of the coil energy is being lost through induced currents being dissipated through the metal walls of the cryostat.

1 — INTRODUCTION

Depuis F apparition des supraconducteurs stabilises, on peut envisager de stocker de facon sure de fortes energies dans des enroulements supraconducteurs. Grace aussi a raccroissement de leur stabilite, on peut les employer dans des systemes fonctionnant en impulsion sans craindre de les voir transiter a Fetat normal.

Nous decrivons ici un ensemble utilisant un enroulement supraconducteur capable de stocker une energie maximale de 20 kJ et de la restituer dans une charge exterieure en un temps de Fordre d'une milliseconde. Un autre enroulement est en construction qui pourra emmagasiner une energie trois fois plus grande.

Grace a sa disposition en galettes et a son secondaire couple, notre bobinage peut s'adapter a des charges d'impedances di verses. En particulier, il permet de decharger Fenergie sous un courant plus eleve que le courant de charge.

Ce nouveau mode de stockage de Fenergie pose un probleme au moment de la liberation; il s'agit d'interrompre rapidement un courant continu dans un circuit fortement selfique. Notre solution a ete d'utiliser un interrupteur electromecanique sous-vide equipe d'un systeme d'extinction d'arc dont les principaux avantages sont: le faible encombrement, le temps de coupure bref et une puissance de coupure elevee.

2 — PRINCIPE DE L'EXPERIENCE

La figure 1 montre le schema general du montage. Une alimentation continue, basse tension, reglable, d'une puissance maximale de 5 kW charge 1'enroulement supraconducteur. Lorsque l'intensite desiree est atteinte, on libere Fenergie stockee en ouvrant Finterrupteur Sj . On peut proceder de deux manieres differentes.

419

1. — Decharge directe de la self de stockage

La charge resistive est placee aux bornes de Si ; Elle se trouve done introduite dans le circuit au moment ou s'ouvre l'interrupteur et I'energie vient s'y dissiper. La source d'alimentation reste dans le circuit mais du fait de son impedance faible I'energie consommee est negligeable.

Charge (procede n°l) i — v w s / — I

A L i

Alim. de courant

de protect ion

Charge (procede n°2)

-^^C-ryostat

Fig. 1

2. Decharge par V intermediate d'un enroulement secondaire couple La charge est placee aux bornes d'un enroulement secondaire fortement couple a

l'enroulement supraconducteur. Durant la croissance lente du courant primaire la tension induite au secondaire est negligeable. Au moment de l'ouverture de Si, le courant primaire decroit tres rapidement en induisant un courant au secondaire. La duree de cette periode de transfert est determinee par la valeur des elements RC de protection places aux bornes de Si. Ce circuit de protection est destine a require les surtensions dues a la self de fuite entre les enroulements primaire et secondaire. La duree de la periode de transfert est inferieure a 1 ms. Dans une seconde periode, I'energie transferee au secondaire se dissipe dans la charge. L'allure des courants primaire et secondaire est visible sur la figure 2.

Dans cette seconde methode, le rendement de recuperation de I'energie stockee est plus faible que dans la decharge directe car, du fait de 1'imperfection du couplage

100A. 240 A.

2mS. Courant p r i m a i r e

2mS. Courant secondaire

Fig. 2

entre les enroulements, une partie de I'energie ne peut pas etre transferee au secondaire et vient se dissiper dans la resistance de protection. Dans nos experiences, ces pertes s'elevent a environ 10% de I'energie stockee.

420

L'utilisation de Tun ou l'autre des modes de decharge et la possibility d'associer en serie ou en parallele les 4 galettes de Tenroulement primaire conduit a des selfs de stockage variant de quelques mH a 620 mH. Le dispositif s'adapte done, suivant le temps de decharge desire, a des impedances variant de 100Q a une fraction d'ohm.

3 — CONSTRUCTION DU BOBINAGE

Le bobinage supraconducteur est compose de quatre enroulements entre lesquels sont imbriquees les 3 galettes qui forment Tenroulement secondaire (voir photo). Chaque enroulement primaire a 620 spires de cable de Nb-Ti stabilise. Le cable est un toron de 7 fils de 25/100 de mm de diametre gaines d'une couche de 0,025 mm de cuivre. La stabilisation de ce toron est assure par une enveloppe en aluminium pur d'un diametre de 1,82 mm, Tisolement est obtenu par un ruban de mylar de 25 \i. En

421

echantillon court, le courant critique est de 320 A a 50 kG et la transition a l'etat nor­mal est parfaitement reversible. Le cable est done tres bien stabilise. Les conditions de refroidissement a l'interieur du bobinage etant plus mauvaises, l'enroulement complet n'est plus stable mais il supporte neanmoins des cycles de croissance et de decroissance rapide du courant jusqu'au voisinage de ses caracteristiques critiques sans perdre ses proprietes supraconductrices. Pour l'ensemble du bobinage alimente en serie le point de transition se situe a 250 A, soit une energie stockee de 20 kJ. On a fait de nombreux essais a 16 kJ, soit un courant de 225 A.

L'enroulement secondaire est en conducteur ordinaire; n'etant traverse par un courant que pendant le temps bref de la decharge d'energie, son echauffement est negligeable. II se compose de 3 galettes de 50 spires de ruban d'aluminium pur d'une largeur de 1 cm. Les galettes sont impregnees dans une resine epoxy et assurent la rigidite mecanique du bobinage primaire.

L'ensemble du bobinage a une longueur de 22 cm et un diametre exterieur de 25 cm il est supporte par un mandrin en plexiglass. Les isolements entre couches et entre gallettes sont assures par des feuilles de mylar. La rigidite dielectrique de l'helium gazeux etant tres faible, il faut apporter beaucoup de soins a l'isolement des conduc-teurs (descentes de courant, contacts) qui ne baignent pas dans l'helium liquide.

4 — FONCTIONNEMENT DE L'INTERRUPTEUR SOUS-VIDE EN COURANT CONTINU

L'interrupteur choisi (Jennings RP 900 K) est prevu pour un courant de 400 A a 50 Hz. II a une tension d'isolement de 70 kV. Un tel interrupteur seul n'est pas capable de couper du courant continu car lorsque les electrodes s'ecartent, une tension faible (20 V environ) sufflt a entretenir un arc qui maintient l'interrupteur virtuellement ferme. En annulant artificiellement le courant pendant un temps tres court, on arrive a eteindre Tare et l'interrupteur est alors reellement ouvert. Le circuit d'extinction d'arc (figure 3) comprend une capacite convenablement chargee que Ton decharge

Thyr is tor Impulsion de commande <>_

kmw-

-AAAAA-

Sll

tL -A /WW Circuit d 'extinction d 'a rc

Fig. 3

dans l'arc en fermant l'interrupteur S2 des que les electrodes de Si sont a leur ecarte-ment maximum. Le temps de fermeture de S2 etant sensiblement egal au temps d'ouver-ture de Si (10 ms), les deux interrupteurs sont actionnes simultanement. Nous avons mesure la duree de l'impulsion du courant d'extinction d'arc; elle est de 10 JLXS dans nos experiences. Son energie est faible : 1 Joule pour couper 500 A. D'autre part, le

422

condensateur sert egalement de protection en limitant la surtension apres l'ouverture de Si.

En cas de defaillance de l'extinction d'arc, il faut eviter que toute l'energie stockee ne vienne se dissiper entre les contacts de 1'interrupteur. Dans ce cas, un relais tempo­rise referme Si au bout de 100 ms.

5 — RESULTATS

Les premieres experiences furent limitees par l'apparition d'arcs entre les conduc-teurs de courant dans l'helium gazeux des que la surtension depassait 5 kV. Par la suite, nous avons pu faire croitre l'energie stockee jusqu'a 16 kJ. La constante de temps de liberation, definie comme le temps au bout duquel 80% de l'energie recupe-rable a ete liberee, a varie de 9 ms a 1,5 ms suivant la charge. Le rendement du transfert varie parallelement de 70% pour une constante de temps de 9 ms a 50 % pour 1,5 ms. Ces pertes importantes sont dues au couplage entre le bobinage et les parois en acier inoxydable du cryostat. Elles seront presque totalement supprimees lorsque nous utiliserons un cryostat a parois en matiere plastique. II ne restera alors que les faibles pertes (une centaine de Joules) dans 1'interrupteur pendant les 10 ms que dure l'arc.

L'interrupteur a fonctionne sans defaillance pendant toute la duree de nos expe­riences. II a effectue plusieurs coupures de courant a 900 A bien que son intensite nominale soit de 400 A. Dans les conditions les plus dures, il a supporte une surtension de 23 kV apres rupture d'un courant de 450 A. La vitesse d'etablissement de la tension apres rupture etait alors d'environ 100 kV/ms. L'energie necessaire a l'extinction d'arc est toujours tres faible vis-a-vis de la puissance a couper : la tension initiale de la capacite de 8 uF est de 600 V pour couper un courant de 900 A.

6 — CONCLUSION

Par suite du couplage important entre le bobinage et les parois metalliques du cryostat, une partie de l'energie se dissipe a chaque decharge par les courants induits. Nous n'avons done pas pu faire un grand nombre de cycles successifs a fort courant a cause de la consommation excessive d'helium liquide. II est neamoins prouve qu'avec des supraconducteurs stabilises on peut liberer l'energie stockee en impulsion breve dont la puissance instantanee atteint 10 MW.

Avec le nouveau bobinage en construction, place dans un cryostat en matiere plastique, on espere atteindre des puissances 3 fois superieures.

Le probleme de la coupure de forts courants dans un circuit selfique est egalement resolu par l'emploi d'un interrupteur electromecanique sous-vide dont les performances ultimes sont, sans doute, tres superieures aux conditions de nos experiences.

Nous avons utilise notre dispositif pour alimenter un tube a decharge lumineuse du genre de ceux utilises pour pomper les lasers a rubis. A performances egales, notre ensemble, y compris le cryostat et les interrupteurs, est d'un encombrement plus faible que la batterie de condensateurs equivalente.

Cette etude a ete menee a bien grace a l'appui de la D. R. M. E.

423

STOCKAGE D'ENERGIE DANS UN ENROULEMENT SUPRACONDUCTEUR

M. FERRIER Direction des Etudes et Recherches, Electricite de France, Clamart {France)

Energy storage in a superconducting winding

SUMMARY: This study examines the problems of the storage of energy on the scale of the fluctuations of the national or regional load, i.e. of the order of 1,000 MWh. Various configu­rations are reviewed and the short solenoid of circular cross-section is picked out as being economically the most favourable. The field produced and the distribution of the mechanical constraints are discussed for this configuration. An examination of the various losses {at the surface of the cryostat, by mechanical or magnetic hysteresis) shows that they are acceptable. The circulation of the cryogenic fluid is discussed in detail and it is shown that it can be maintained by natural convection; it is also shown that the ratio of stabilizing material to superconducting material can be reduced to 10:1. Finally the modes of connecting the storage device to the grid are examined and there is a discussion of the optimum size of the installation as a function of the financial benefits that can be attributed to the different services provided by the device.

Si c'est sous forme electrique que l'energie est le plus souvent transportee, au contraire, son accumulation sous cette forme n'avait pas ete envisagee avant ces dernieres annees. L'energie pouvait etre stockee sous differentes formes [1] : electro-statique dans des condensateurs, chimique dans des batteries, cinetique dans des machines tournantes ou potentielle dans des centrales de pompage par exemple. Si on exclut l'accumulation sous forme d'energie potentielle, tous les procedes prece­dents sont trop couteux pour etre employes au stockage d'une energie importante. Quant aux stations de pompage, l'energie qu'elles peuvent accumuler est limitee par l'existence de sites convenables.

Aussi est-il interessant d'examiner les possibilites de realisations d'accumulateurs d'energie sous forme de champ magnetique, rendues possibles par 1 'absence de resis-tivite dans les supraconducteurs. II n'existe actuellement aucune realisation a grande echelle d'un tel dispositif, cependant des bobinages supraconducteurs ou a metaux refroidis sont deja utilises pour stocker puis liberer des energies de l'ordre de 106

a 108 J [2-5]. On se propose ici, au contraire, d'examiner les problemes poses par des bobinages emmagasinant une energie nettement plus grande, par exemple 101 3 J (soit 2 800 MWh). En effet seule l'accumulation d'energies de l'ordre de 1012 J est interessante pour la regulation de la production d'electricite a l'echelle nationale. Les problemes poses sont de nature tres differente, d'une part a cause de la taille de 1'installation et de la necessite defaire intervenir des grandeurs economiques dans 1'optimisation, et d'autre part, a cause d'un fonctionnement essentiellement different.

CONFIGURATIONS ENVISAGEABLES POUR L'ACCUMULATEUR

Nous procedons en deux etapes pour etudier la configuration de l'accumulateur: la premiere, qui| fait I'objet de ce paragraphe, degage les grands traits des differentes solutions possibles afin de restreindre le champ d'investigation. C'est ainsi que nous rechercherons d'abord quels types de structures peuvent etre envisages, en se guidant au moyen du critere approximatif que constitue la possibilite d'accumuler le maximum d'energie pour un volume donne de supraconducteur.

Designons par Bm l'induction maximale de fonctionnement; il lui correspond une valeur Jc de la densite critique de courant dans le materiau supraconducteur et Ton choisira la densite maximale de courant en fonctionnement normal Jw a une

425

valeur juste inferieure; distinguons encore la densite moyenne de courant J dans l'espace. L'energie stockee est W = JJJ (B2/2 | i 0 ) dv et le volume de supraconducteur utilise est V = (1/JC) JJJ |J| dv. Considerons une solution proposee et appliquons lui une homothetie geometrique de rapport x; pour conserver les memes conditions physiques de fonctionnement (en particulier l'induction qui, comme on le verra, determine aussi les contraintes mecaniques) il faut considerer une densite moyenne de courant J dans le rapport x (puisque chaque fil reste parcouru par le meme courant). On voit que l'energie stockee est multiplied par x3 et le volume de supraconducteur par x2. Aussi, afln d'eliminer le facteur taille de Installation on recherchera la structure qui correspond au maximum de W2/V3.

Sachant a l'avance [4, 6] que les problemes de resistance mecanique de l'accu-mulateur sont considerables, que l'enceinte cryogenique n'est pas d'un cout negli-geable, il faut eliminer les solutions de forme trop compliquee, et done, ne retenir que les repartitions de courant a symetrie de revolution, les spires etant soit dans des plans paralleles, soit dans des plans meridiens.

Lorsque toutes les spires sont placees dans des plans meridiens, elles constituent une sorte de bobinage referme sur lui-meme. Par application du theoreme d'Ampere, on montre facilement que le champ magnetique est nul a l'exterieur de l'instal-lation ce qui est un des avantages fondamentaux de cette classe de solutions. Ces solutions seront definies par le trace des spires dans un plan meridien, en remarquant que, si les solutions les plus simples ne comportent qu'une spire dans chaque plan, il est cependant avantageux d'en placer plusieurs de formes differentes. C'est ainsi que pour une section rectangulaire a 1 seule couche de spires on obtient un parametre de performance W2/V3 = 1,1.1020 J2 /m9 (Toutes les evaluations sont conduites avec les valeurs de Bm = 7 T et Jw = 3.109 A/m2), alors que si, au voisinage de I'axe les spires rectangulaires sont regulierement reparties, on obtient W2/V3 = 1,97.1020

J2/m9. Nous avons conduit une etude detaillee de la forme de spire, a une seule couche, qui permet d'obtenir le maximum de W2/V3 et Ton a trouve que la section meridienne se compose (fig. 1) d'une partie rectiligne au voisinage de I'axe et d'une courbe d'equation [7]

y= ± * log uib * ! ; ■ : n f i ^

du (avec A = 0,85) VA2 - (log ujb)2

On obtient alors W2/V3 = 1,65.1020, valeur qui pourrait etre augmentee dans les memes proportions que pour la spire rectangulaire, par l'emploi de plusieurs spires imbriquees.

Au contraire lorsque toutes les spires sont placees dans des plans paralleles, le champ magnetique n'est pas confine a l'interieur de la structure, et les solutions sont definies par la donnee de la repartition de courant traversant l'un des plans meridiens. L'imperatif d'obtenir une inductance propre elevee (forte energie stockee) conduit a placer les spires tres proches les unes des autres; en revanche la contrainte de ne pas depasser l'induction de fonctionnement Bm limite cette possibilite. L'optimum est atteint [8], lorsqu'on ne prend en compte que le volume de supraconducteur, pour un bobinage en forme de tore dont le grand rayon R vaut 4,14 fois le petit rayon r (fig. 2). A l'interieur de la section meridienne la densite de courant doit etre a repartition de revolution autour de la fibre moyenne du tore et decroitre de facon inversement proportionnelle a la distance y a la fibre moyenne, sauf au voisinage de cettet fibre ou elle doit etre constante et done egale a la limite technologique.

ELEMENTS OU COUT Nous examinerons trois postes : le^cout du materiau supraconducteur, le cout du

cryostat et celui des elements destines a assurer le maintien de la structure face aux

426

--*-102m

[20,7m |

65m

Fig. 1 — Schema de Taccumulateur a une seule couche dans des plans m6ridiens.

Paroi a la remperahjre de I'azore liquide

Paroi a ia remperahjre de I'helium liquide

Element- comporranr des tils supraconducheurs erunegaine de circulah'on d'helium

Fig. 2 — Representation sch6matique de l'accumulateur a spires dans des plans paralleles. (Les 616ments constitutifs du support micanique ne sont pas repr6sentes).

efforts mecaniques. Le calcul du volume de supraconducteur et de la surface de cryostat est facile pour chaque structure.

Pour les efforts mecaniques on retiendra les conclusions suivantes [8] : — Les contraintes dependent beaucoup du niveau de l'induction et atteignent des

valeurs considerables lorsque l'induction depasse 10 T; elles ne dependent pas de la taille de l'accumulateur et done pas non plus de l'energie a stocker ;

— Pour la configuration a une seule couche de spires optimales dans le plan meridien (fig. 1), les efforts correspondent, dans toute la partie incurvee de la spire, a une

427

traction simple dans le sens de la spire, efforts qu'il est facile de contenir. Dans la partie rectiligne de la spire il se superpose a cet effort de traction une poussee axiale constante qui peut etre retenue par une frete ; Pour la configuration optimale a spires dans des plans paralleles, les contraintes sont maximales sur la fibre centrale; la composante la plus defavorable y correspond a une traction dans le sens perpendiculaire aux plans meridiens et sa valeur est

K Ho

log 8R 0,5

D'autre part les contraintes macroscopiques ont des valeurs propres de signes opposes, ce qui entraine la possibility de flambage.

Les elements de cout sont donnes au tableau 1, pour un accumulateur de 1013 J et par rapport a la solution ci-dessus.

Tableau 1 RECAPITULATION DES ELEMENTS DE COUT DES DIFFERENTS ACCUMULATEURS POUR UNE ENERGIE

DE 1013 J, AVEC J,„ = 3.109 A/m2, Bm = 7 T ET UNE CONTRAINTE MAXIMALE DANS LES ELEMENTS SUPPORTS DE 1 0 9 P a

Designation de la

solution Schema

Volume .. . de supra-geometnques , , conducteur

Grandeurs Surface de cryostat

Volume de support Cout

Solution optimale M^0 ,v 68 m

8,5 m 42,3 m3 2,3 .104m2 1,39.104 m3 1

Spire meri-dienne rec-tangulaire optimale

R

*2b*

|

R = 55,7 m a = 74,2 m 96,3 m3 1,55 .105 m2 irrealiste b = 37,1 m

Spire rectan-gulaire a densite repartie

R = 70,2 m A = 101,8 m 82 m3 7,58.104 m2 irrealiste r = 21,9 m

Spire meri-dienne opti­male

A = 130 m A R = 1 0 2 m 85m 3 8,5. 104 m2 1,47. 104 m3 1,90

r = 20,7 m

ENVIRONNEMENT THERMIQUE ET STABILISATION

Les pertes parietales procedent de deux phenomenes physiques difficiles a dissocier: rayonnement et conduction; de nombreuses solutions sont proposees par les fabricants

428

de cryostats ou ces processus sont imbriques; il parait raisonnable d'adopter la valeur de 0,05 W/m2 a 4°K. Le poids considerable de Taccumulateur oblige a prevoir aussi les pertes par le systeme de supension. II semble judicieux de soutenir le bobinage et son support au moyen de tiges en acier inoxydable et d'assez grande longueur; on peut alors reduire les pertes a 400 W a 4°K.

Les pertes par hysteresis magnetique sont plus importantes : en appelant d le diametre des elements supraconducteurs, la perte volumique d'energie est approxi-mativement BmJcd/2 pour chaque cycle. Pour Tensemble de Taccumulateur les pertes par hysteresis magnetique represente done la fraction (Jm/Jc) d/2r de l'energie stockee; ainsi pour des systemes de conducteurs de 5 mm d'epaisseur chaque cycle ne dissipe a 4°K que la fraction 10" 4 de l'energie stockee, ce qui represente une perte de 3,4% de l'energie maximale stockee.

Les pertes par hysteresis mecanique du materiau support sont moins faciles a determiner. Designons par Y le module d'Young du support qui est soumis a une pression p. Avec les resultats du paragraphe precedent on trouve une energie stockee de 36 n2 R r2 (B^/jx2,), soit la fraction 9B2

1/u0Y de l'energie electromagnetique. En prenant pour Y, 2.1010Nm~ 2 on trouve done que l'energie mecanique ne represente que 8% de l'energie electromagnetique. Si Ton veut, par exemple, ne perdre par ce processus que 1 % de l'energie stockee, il faut que le materiau support ne dissipe a chaque cycle que 0,17% de l'energie mecanique qu'il absorbe; on ne dispose pas, actuellement, de donnees sur l'hysteresis a tres basse temperature, mais les chiffres precedents paraissent raisonnables car on peut penser que l'hysteresis diminue a ces temperatures.

Les evaluations precedentes montrent que, pour un accumulateur de 1013 Joules, on peut admettre comme limite superieure de la puissance perdue a la temperature de l'helium liquide, 104 W. Aucune installation actuellement realisee dans le monde n'atteint cette puissance. Neanmoins, la realisation de tels refrigerateurs est d'ores et deja tout a fait envisageable et des estimations de cout montrent qu'il etait tres justifie de ne pas faire intervenir le prix du refrigerateur, qui est negligeable par rapport au cout de l'installation.

Enfin pour la stabilisation, il parait exclu de prevoir un regime perturbe dans 1'ensemble du bobinage; les puissances mises en jeu sont trop considerables et la solution doit etre recherchee dans une surete absolue de resorption de petites zone normales. On montre alors que la section de supraconducteur Ss est liee a la section dejstabilisant St par la formule :

S[ _aPAT S* ~ 2 J 2 p

ou a est le coefficient de transfert (a ^ 5000 W/m°K), p est la resistivite du stabilisant (nous avons choisi p = 8.10"J 1 Qm). A T est l'ecart de temperature entre le metal et l'helium. P est un facteur de forme (|3 = 2\/n pour une structure cylindrique). Pour un courant de 105 A par fil on trouve que Sf/S4 = 40, mais Sr/Ss = 11 seulement pour un courant de 2.103 A.

PROBLEMES D'UTILISATION

Nous avons exclu l'insertion de Taccumulateur dans un reseau en placant Taccu­mulateur en serie avec les installations de consommation, cause d'un surdimension-nement considerable des installations de conversion du courant.

429

Le systeme envisage (fig. 3) prevoit deux onduleurs pour alimenter la distribution, Tun en serie avec raccumulateur, I'autre en parallele. Le premier est utilise sous une tension variable E2 et le second sous une tension fixe E.

Distribution

Production

Fig. 3 — Utilisation proposed pour raccumulateur.

Fig. 4 — Comparaison entre le cout de raccumulateur et l'Sconomie sur les installations de production. ;h

430

On montre alors que, en fonction de la puissance consommee P, et des constantes / (inductance propre de l'accumulateur) E (tension de la ligne) et I (courant dans la ligne) on determine E2 par la formule :

1 ' '~ -EI)dx ( E 2 - E ) 2 / ( P - E I ) J>-'

oil l'origine des temps est choisie a un moment o u P = El. Cette solution ne conduit a aucun surdimensionnement pour l'onduleur en parallele avec l'accumulateur, et a un facteur 2[1 + (E/ V /co AP)] pour l'autre en designant par AP l'amplitude des variations de puissance consommee et par co leur pulsation.

A ce stade on dispose de tous les elements pour le choix du courant de fonction-nement. Le courant dans chaque brin doit etre assez faible pour ne pas trop augmenter les pertes magnetiques et favoriser la stabilisation (par exemple 2 000 A); le courant d'alimentation ne doit etre ni trop grand pour ne pas trop surdimensionner l'onduleur, ni trop faible, ce qui conduirait a des tensions de fonctionnement trop grandes et des champs electriques importants dans le bobinage. Enfin il ne faut pas trop fractionner chaque conducteur car cela ne conduit a aucun gain de stabilite et risque d'augmenter les couts. Aussi la valeur retenue est de 1,4.105 A pour 1'alimentation; chaque conduc­teur etant ensuite divise dans l'accumulateur en 70 brins de 2 000 A.

CONCLUSION

Reste a juger de la rentabilite globale de ce dispositif. On dispose de tous les elements de cout, mais avec une grande incertitude sur les prix des differents compo-sants. En regard, les elements de valorisation de 1'installation sont de deux sortes : D'abord, l'existence de l'accumulateur permet de reduire la puissance installee a cause de la pointe journaliere de consommation; cet effet est analyse a la figure 4, qui montre que le gain est d'autant plus grand que les fluctuations de consommation AP sont plus grandes. Ainsi, dans le cadre des hypotheses faites pour les evaluations de couts, le stockage semble devoir etre nettement rentable, dans le cas du reseau francais, pour une fluctuation de puissance de 20 000 MW. En outre il faut tenir compte de l'usage qui peut etre fait d'un accumulateur pour soulager momentanement un reseau a la suite de la perte accidentelle d'une centrale. II semble qu'il suffise alors d'un accumulateur de taille beaucoup plus reduite (1010 J environ), mais l'estimation du service rendu est alors beaucoup plus delicate.

REFERENCES

[1] R.C. CARRUTHERS, The storage and transfer of energy. High Magnetic Fields (1962). [2] V. A R P , Cryogenic Coil for megajoule energy storage—R.375—Proceedings of the int.

symp. on magnet techn. (1965). [3] L. DONADIEU, D.J. ROSE, Conception and design of large volume superconducting

solenoids—High magnetic fields (1962). [4] Z. J. J. STEKLY, Magnetic energy storage using superconducting coils—AVCO-EVERETT

report no. AMP. 102. [5] J. SOLE, Stockage d'6nergie. Possibilites des supraconducteurs en vue des decharges de

grande puissance. Rapport CEA—R—3243 (1967). [6] B. LEON, Calcul des contraintes m6caniques dans les bobinages de revolution destines a

la production de champs magnetiques intenses—R G.E.T. 73 n° 12, p. 632 (1964). [7] M. FERRIER, Comparaison des performances de divers accumulateurs — Note interne

E .D.F . — Direction des Etudes et Recherches — HM.041-53 (1968). [8] M. FERRIER, Problemes poses par le stockage d'&iergie 61ectrique dans un bobinage

supraconducteur. Note interne E. D. F. — Direction des Etudes et Recherches — HM.041-26 (1968).

431

DISCUSSION

C. DAM MANN (France) — La multiplicity des parametres qui conditionnent un tel systeme, rend difficile la recherche d'un optimum reel. Cependant, dans le cas des energies elevees tout particulierementjepense que l'optimum doit etre recherche par la determination de la densite de courant effective du bobinage qui rend minimal le cout total du systeme (bobinage et cryogenie). En effet, pour les grands bobinages de stockage d'energie, le cout optimal n'est pas directement lie a la minimisation du volume de supraconducteur.

M. FERRIER — La premiere ebauche de calcul ne prend en effet en compte que le cout du materiau supraconducteur. En revanche, l'optimisation complete, correspondant aux courbes de cout fournies, prend en compte le stabilisant, le cryostat, le support mecanique... On remarque alors que l'optimum est beaucoup plus plat en fonction du champ magnetique.

P.H. ASHMOLE (U.K.) — Has Mr. Ferrier any actual, rather than relative cost figures for his superconducting energy storage proposals?

M. FERRIER — The purpose of this work has been to study the technical feasibility of such a storage device, and economical aspects only concerned the choice of the best configuration. In fact, the large incertitude concerning the prices of fashioned materials involved in the design cannot be reduced at present. Never­theless one can say that the cost of superconducting energy storage would be greater than that of pumping units and, optimistically speaking, comparable to the gains in power equipment within a few years.

C. FURTADO (U. K.) — How are you going to feed the stored energy into the 50 Hz grid ? What is the cost of this ancillary equipment in comparison with the cost of the coil ?

M. FERRIER — The connection between the coil and the network largely depends on the part assigned to the storage (peak-lopping, reserve in case of fault...) and on the position of the coil in the network (at central position, or at the end of a line). The solution presented in this paper concerns a coil at the end of a DC line, near a centre of distribution. A variable direct current flows through the coil and the decrease of this current corresponds to a voltage applied to the 2n d converter that means transmission of power from DC circuits to AC circuit.

So far no studies have been done at E. D. F. concerning the cost of converters, but I think it is reasonable to assume that their cost is negligible compared to that of the coil.

432

SUPERCONDUCTIVE ENERGY STORAGE WITH A FLUX PUMP*

D.E. SPIEL Atomics International, Canoga Park, California

U. S. Naval Postgraduate School, Monterey, California

and

R.W. BOOM Atomics International, University of Wisconsin, Madison, Wisconsin (U.S.A.)

Stockage d'energie avec une pompe a flux

RESUME : On a eprouve une bobine de stockage d'energie de 8 000 amperes 3 000 joules char gee par une pompe a flux a redresseur pleine-onde. Vappareil a 1 rn.de haut et a un diametre de 33 cm et un poids de 60 kg. Toutes les parties sont concues pour un courant de 16 000 A au moins. On discute la limite de 8 000 A par rapport au comportement electrique et thermique des elements de commutation de la pompe. On donne les details de la construction de la pompe et de la bobine et Von examine Vefficacite de pompage et de stockage.

INTRODUCTION

In general, a superconductive energy storage unit includes a storage magnet, a persistent switch to trap the energy in the magnet, a power supply to charge the magnet, a scheme to remove the stored energy, and a cryogenic system to cool the superconductors. Compared to ordinary resistive inductors there are several major advantages for the superconductive case: (1) energy can be stored indefinitely without loss; (2) charging can be slower, permitting use of relatively simple and inexpensive power supplies; and (3) weight and size efficiencies are significantly improved since current densities are several orders of magnitude higher. The major disadvantage, of course, is the cryogenics involved. Taking these factors into account one is led to the conclusion that for sufficiently large energies ( > 100 kJ) superconductive units are economically reasonable.

This paper is a report on two of the technical areas mentioned above, a flux pump-power supply and 10 kJ storage magnet designed for especially high currents. Empha­sis is placed on overall efficiency and possible extrapolation to higher energies. Since this work is still in progress there are mentioned below two sets of specifications, one for performance to date and the other for original design specifications.

DESIGN

The requirements of the unit were to store up to 10 kJ in such a manner that discharging into a 5 uh-10 mQ load would deliver a 20 kA peak current in a damped oscillatory mode at 5 kHz once per 5 minutes. A very important additional constraint was that this device should be a model study for MJ systems; this meant that simpler techniques appropriate to 104 J systems, but not appropriate in a MJ system, were avoided. The three major design areas covered here are: (1) flux pump, (2) magnet, and (3) assembly and interconnections.

* Work supported by NASA Contract NAS 8-11907.

433

The flux pump is a modified fullwave rectifier type with saturable transformer similar to that introduced by Watt in 1964 [1]. A simplified circuit is shown in figure 1. Current, I, is built up in the storage magnet, L, by periodically reversing the flux in the pumping core C as switches Sx and S2 in the secondary loops are alternately turned on and off. In the normal state, Sx and S2 are sufficiently resistive so that they function as diodes [2]. Core C is heavily overdriven from saturation to reverse satura­tion by each reversal of Vp so that 2§s is added to the secondary each half cycle, where (|)s is the saturation flux of core C. In this way, load current accumulates in L via the one turn secondaries at the constant rate Al/cycle = 4(|)S/L, for L much larger than other secondary circuit inductances. The complete saturation of core C is impor­tant in order to reduce the losses involved in switching.

Current Limiter

—UULU/1

Fig. 1 — Flux pump circuit. Energy storage magnet L; stored current I; saturable trans­former core C with primary P and two turn center-tapped secondary S; primary volt­age ± Vp; and cryotron elements Si and S2 switched resistive by cores C and C"

The sequence for a pumping cycle can be described as follows, with reference to the circuit in figure 1: 1. Initially assume a steady state with a current I flowing in the upper secondary

branch and no current in the lower branch; St is superconducting and S2 is normal. Core C is saturated; the primary current is set by the current limiter; and the primary voltage can be taken as positive, + Vp.

2. The first step is to turn off the drive to core C" which causes S2 to become super­conducting. Following that, the C core drive is turned on causing Sl to become normal, which transfers the stored current I to the lower branch without changing the direction of I through L.

3. Next, pumping takes place by reversing the primary voltage to - V^. A secondary voltage, Vs = d(j)/df = L dl/dt = L I 0 / T = Vp/N, appears across the storage magnet and increases the stored current by Al = 2§JL during a time period At = L AI N/Vp. N is the number of primary turns and T is the total pumping time to store the final current I 0 . T = n 2 At, where n is the number of complete pumping cycles.

434

4. After time At, core C saturates, Vs drops to zero, and I + Al circulates persistently in the lower branch.

5. At this point, Sx is made superconducting and S2 normal which transfers the stored current I + Al back to the upper branch. Core C remains in far saturation inde­pendent of the direction of secondary currents due to sufficiently excess primary current.

6. As in step 3, this second pumping step is accomplished by reversing the primary voltage, this time to + Vp, which applies the voltage Vs = Vp/N across the storage magnet. The current increases to I + 2 AI in the storage magnet and upper secondary branch. Thus the system returns to the initial conditions, one cycle has been completed, and 2AI has been added in storage.

The above pumping description, for the purpose of clarity, ignores energy and current losses. However, such loss considerations are of primary importance for the design selection of the various pump elements [3]. The most significant loss is unavoid­able and inherent to this pump. During pumping the resistive branch carries a back current i = 2Vs/r, where r is the normal resistance of either switch, Si of S 2 . This results in a total energy dissipation equal to 4V2

sT/r. Storage efficiency, g, considering this loss alone, is:

energy stored 1 1 /A. 8 = = = (1)

energy stored + energy lost , n L/r A A i &J bJ 1 + 8 — 1 + 4 — T 10

Equation (1) shows that for efficient operation total pumping times must be long compared to the time constant L/r; or, in an equivalent statement, that the back current in the cryotrons must be small compared to the final stored current. Perhaps a more useful expression is the energy lost in pumping E 0 and I 0 into inductor L:

8 V 4 i AE = E * = E ( 2 )

r l 0 Io

Note that neither repetition rate, core size C, nor storage inductance L affects the fractional energy loss. Only the permanently chosen cryotron resistance r and the operational secondary voltage Vs are significant with respect to this loss mechanism.

Another loss can occur when current is transferred from one secondary turn to the other unless precautions are taken. If transfer is caused by simply turning off one switch then opening the other, the loss of stored energy is 5E/cycle « (l+m) I2 , where / is the self inductance of one secondary turn plus the leads and m is the mutual inductance between the two secondary circuits. Assuming no interaction between secondaries and primary during switching, this energy is supplied from two sources. First, the energy inductively stored in (l+m) is dissipated in the initiating switch during switching. This same amount of energy is withdrawn from L to charge the equivalent inductors in the other secondary branch in which the current increases from zero to I. This repeats twice per cycle and results in a loss of stored current 81 « — (l+m) I/L and the energy loss, 8E, mentioned above. For the ideal case, in which Al added » 81 lost, it is seen that the total energy lost by this process is n(l+m) IQ/3, where n is the number of cycles. Reduction of / and m are, therefore, important and one sees at once the reasons for saturating core C before switching, since then / and m are smaller air core inductances. Also, in this context, the secondary inductance is reduced to 1/10 by dividing each secondary turn into 10 parallel turns. Buchhold has reduced this loss by a method in which the cryotrons are switched normal

435

only at low current levels [4]. In the present case, transfer of current from one secondary turn to the other can be affected by reducing core C excitation with both S, and S2 superconducting. It can be shown [5] that to affect such a loss-less transfer requires

SWITCH COIL

PRIMARY WINDING

SECONDARY

SWITCH ' ; ELEMENTS

SWITCH

CONNECTORS

SWITCH

TRANSFORMER

CONNECTOR

SHIELD

STORAGE MAGNET

436

Fig. 2 — Assembled flux pump and energy storage magnet.

a change in primary current, 5lp = ( /+w)I /M, where M is the mutual inductance between the primary and the secondary turn. In this way, by programming the primary, the cryotrons can be switched normal at low current levels.

A third loss results from hysteresis in the pumping core and the switching cores. The loss for the pump core is independent of its size (or number of cycles) while the losses for the switch cores are cyclic. A reduction of these losses therefore requires small switching cores and fewer cycles or, equivalently, a large pumping core.

The pumping core is a square orthonol toroid 14.5 cm2 by 45.8 cm mean circum­ference and weighs about 5 kg. The hysteretic loss per cycle for this material is about 13 mJ per kg. The primary winding is 3,000 turns of 0.019 cm Ti-Nb clad with copper to 0.037 cm and nylon insulated. The critical current at 30 kgauss is 70 A although only 10 A is required.

The center-tapped secondary is two turns of copper stabilized Ti-Nb strips. In order to reduce the inductance and to divide the current into several increments, each secondary turn is 10 parallel loops evenly spaced around the toroid. The Ti-Nb conductors are 1.25 cm x 0.0127 cm-thick bonded to 0.08 cm-thick copper. Each superconductor, which is required to carry up to 1,600 A, was sized to carry 4 times that amount; sufficient copper was used to completely stabilize the conductor at 5 K without any possibility of higher temperatures [6]. The inductance of each half secondary is 10" 8 H. Some features of the assembly are shown in figure 2.

The cryotron switches Sx and S2 were required to carry 16,000 A each. The normal state resistance of each switch is 10" 4 Q, a conservative choice to achieve 80 percent efficiency for a 5 min pumping time. An alloy of Nb-1 w/o Zr was used because the 4.2 K normal resistivity was reasonably high (1.3 |iQ-cm) and the critical field low enough (7 kgauss) for ease in switching. The switches were necessarily unstabilized. Tests with strips of this material showed that the current density did not vary linearily with either width or thickness [7]. For example, the current capacity of a 0.005 cm-thick strip tends to become constant for widths greater than 2 cm; and doubling the thickness to 0.01 cm increases the critical current only about 10 percent. It was found, however, that by cutting gaps along the length of the strip a significant increase in current could be achieved. For example, cutting out three 1/2 cm-wide strips spaced 1/2 cm apart on a 3-1/2 cm-wide, 0.005 cm-thick strip increased its current capacity from 1000 to 1800 A.

Based upon this experience, each element of the two pump switches consists of a configuration of five 1/2 cm-wide, 0.005 cm-thick strips separated by 1/2 cm-wide gaps for a total width of 4-1/2 cm. The strips were 19 cm long, insulated on one side by adhesive mylar, and folded twice to make a bifilar package measuring about 4-1/2x4-1/2 cm. Each element was closely fitted into one of twenty 1/8 cm-wide, symmetrically spaced gaps cut into a magnesil toroidal core with dimensions the same as the orthonol core. Only one such element is sketched in figure 1 although twenty in parallel comprise switch S t or S2 (see figure 2). Subdividing S1 and S2 into 20 parallel units reduces the inductance and losses (see eq. 1 and 2), provides the possibility for current sharing, and distributes the many joints around a cylinder for mechanical convenience.

The 80 |iH storage magnet is 27 turns of 2.5 cm-diameter cable wound in a solenoid 31 cm-OD by 11 cm-ID by 19 cm long. The cable assembly, shown in figure 3, consists of 7 subcables twisted together with internal spacing of 0.09 cm to provide surface cooling. Each subcable is composed of a copper clad 0.165 cm-dia Ti-35 w/o Nb wire soldered to six 0.2 cm-dia copper wires. The copper cladding and shunt wires are sufficient to insure stable operation to 20,000 A in a 25 kgauss field [6]. The cable is insulated with 1/3 cm mylar rope wound one turn per cm.

The assembly, shown in figure 2, is mounted on a central rod with the magnet separated from the pump by a shield which reduces the maximum field at the switches

437

to 150 gauss. The major problems of assembly were to arrange all elements so as to minimize lead inductance and to make the many hundreds of superconducting joints. Weld joints were used throughout. One 0.1 cm-diameter spot weld of 50 watt-seconds carries 300-700 A critical current in low fields for the various combinations of strip and wire used. At least three times the needed welds were made at each joint. The pump to magnet connectors were welded joints soldered to very heavy copper cylinders which serve as shunts. In figure 2 the two output connectors above the pump are shown shorted together; another cable to strip connector is directly below the magnet.

RESULTS

Five separate runs were made during which improvements and adjustments increased the system capability to the 8,000 A level. All runs were monitored with an eight channel oscillograph which recorded primary voltage and current, secondary voltage, stored current (central field), S t and S2 switch drive, and an isolated dummy S1 element. The dummy element was mounted beside one of the S t elements in one of the core slots and, with 10 A transport current, was used to monitor switch resist­ances and stray pickups. This extra element was particularly useful in the analysis of recorded performance; it also helped in setting the switch core drive levels and assisted in determining time and rate of switching and recovering.

SUBCABLE 7-0.080- in . DIAMETER 0FHC Cu-CENTER HAS Ti - Nb0.065-in. DIAM CORE

OFHCCu0.035-in. DIAM SPIRAL WRAP FOR HELIUM VENTILATION

TOTAL CABLE CONSISTS OF 7 SUBCABLES SPIRAL WRAPPED ABOUT 0NE-AN0THER

Fig. 3 — Magnet conductor

Early in the testing a curious sort of mutual switch interference was observed. When the transformer was unsaturated it acted to shield one switch from the other. At the moment the transformer core saturated, however, it ceased to shield and the stray field from the excited switch core abruptly reached the unexcited switch and drove those switch elements normal. As seen in figure 2, the three cores are mounted close together with the two switch cores separated by the pumping transformer core. The result was a degraded critical current capacity even though the stray field was only a few gauss. It was subsequently observed that a 1.5 kgauss bias field on the affected switch eliminated the degradation. In the final pump, however, it was elimi­nated by a 0.3 cm-thick magnetic shield of interleaved sheets of copper and con-netic material placed between the transformer core and each switch core.

438

During the second run a half layer short in the primary winding developed. After repair, an arbitrary maximum of 10 volts on the primary, not 100 V as planned, was used in order to complete the tests before disassembly for a primary rewind. The efficiency, according to equation 1, for a 10 volt primary and 8,000 A stored, is 97 percent. If 100 volts were used to pump to 16,000 A then the efficiency would be 87 percent and the "live" time to pump, T, would be 38 sec. The highest primary voltage used, however, was 20 volts which, for the 8,000 A achieved to date, indicates a peak achieved secondary power of 50 W at about 94 percent efficiency. The design power is 500 W.

These preliminary runs show that the 8,000 A limit was determined exclusively by the unstabilized switch elements. All transitions occurred between pumping steps when the load current was being transferred from one to the other secondary circuit, steps 2 and 5 above. The short sample capacity of the switch elements was 36,000 A as determined empirically in similar mounting configurations. Thus, the 8,000 A limit is not at all understood. These elements were, of necessity, poorly cooled; each was well insulated and tightly confined in a 1/8 cm switch gap. Gradual heating, however, could not have occurred since switching for these runs was at a low rate in order to study end point conditions.

The loss theory was directly confirmed for the / I 2 term involved in the switching step. For example, at 6,000 A, the oscillograph trace shows a secondary pulse approxi­mately 0.2 sec long and 0.2 mV high for an energy loss of 1/4 joules while /I2 is calculated to be 0.36 joules. The experimental data for this particular circuit indicate that rates up to one half cycle per second should be achievable, although definitive rate tests have not been studied.

In conclusion, it appears that high currents can be stored and pumped into a sizeable storage magnet by saturable core flux pumps. High efficiencies, 99 + percent, should be expected for such large units provided enough time (low drive) is used for charging.

ACKNOWLEDGEMENT

The authors wish to thank S.L. Wipf, Atomics International, and K.R. Mac-Kenzie, UCLA Physics Department, for many valuable discussions.

REFERENCES

[1] D.A. WATT, "A Superconducting Transformer and Charging Circuit for a High Field Magnet", A.E.R.E.—M 1397, Harwell (1964).

[2] J.L. OLSEN, "Superconducting Rectifier and Amplifier", Rev. Sci. Instr., 29, 537 (1958). [3] S.L. WIPF, "The Efficiency of Flux Pumps", Proc. Int. Conf. Magnet Technology,

Stanford 615 (1965) and Proc. Int. Cryogenic Engr. Conf., Kyoto, 137 (1967). [4] T.A. BUCHHOLD, "Superconductive Power Supply", Cryogenics, 4, 212 (1964). [5] D.E. SPIEL and R.W. BOOM, "Superconductive Energy Storage System, Fourth Tech.

Rpt., Contr. NAS 8-11907, 63 (1968). [6] C.N. WHETSTONE and R.W. BOOM, "Superconductor Stability in Liquid Helium",

Advances in Cryogenic Engineering, Plenum Press, 13, 68 (1968). [7] E .D. HOAG, "Critical Current Degradation in Wide Strip Superconductors", / . Appl.

Phys., 36, No. 3, 1183 (1965).

D I S C U S S I O N

S. H . M I N N I C H (U.S .A . ) — What was the means of transferring the energy to the load?

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R. BOOM — A series resistor of approximately 2Q was to have been inserted at the rate of about 104 to 105 ohms/sec in series with the magnet. That would divert the stored current into the load with only a few hundred joules lost.

S. H. MINNICH — The load transfer switch would be just about as big as the coil ?

R. BOOM — Yes.

S. H. MINNICH — One can then observe that the volume of the total system is 3-4 times that of the coil and that the theoretical energy density, B2/87C, must be degraded by this amount. It also seems that flux pumps tend to be as large as magnets.

R. BOOM — I agree in general with that trend, but point out that the size of this flux pump was determined mostly by the 16,000 A current level which sets a physical minimum for sizes of conductors and associated elements.

M. YAMAMOTO (Japan) — There are two kinds of flux pumps, the rectifier type and the moving field type. Both types can get heavy current, but it is very difficult to get high voltage. I think that flux pumps are not suitable for energy storage: I should like to know the lecturer's views on this.

R. BOOM — The flux pump is not particularly suitable for high voltage appli­cations since millivolt outputs are typical. The pumps are most useful when low repetition rate discharges are required. Also, flux pumps only convert energy, generally from a low current level to a high current level. As converters, flux pumps do not particularly change the logic for or against superconductive energy storage. Appro­priate power levels, pump voltages and currents, must be used.

T. BUCHHOLD (West Germany) — The flux pump built by GE works with 7.5 c/s instead of 0.5 c/s. Therefore with 7.5 c/s a pump can be built much smaller. The efficiency obtained so far is about 97.5%. It seems possible to obtain in the future 99%.

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RAPPORT GENERAL

PERSPECTIVES D'AVENIR DE LA CRYOELECTROTECHNIQUE

J. FABRE Direction des Etudes et Recherches, Electricite de France, Paris {France)

Future trends in cryoelectrotechnics

SUMMARY: The author reviews possible applications of very low temperatures to the com­ponents of large electrical generation and transmission networks in terms of probable evolutions in the next twenty years.

Taking as an example the French network where generation is in the provinces and a major part of the electrical consumption is concentrated in the Paris region, it will be seen that one of the main applications may be that ofd.c. or a.c. cryo-connections in relation to the improvements that may be made in controlling a.c. losses and in the technology of cryogenic substructure.

However, other interesting applications may be envisaged for generating stations where the use of superconducting inductors may enlarge the range of unitary power available at the moment.

Owing to their essentially large magnitude, projects on electrical power storage in super­conducting coils are more hazardous.

Finally, machines employing transition processes between the normal and superconducting states need considerable improvements from the material point of view in order that they may achieve future industrial potential.

A comparative study of the evolutions in technology and in living matter shows that a choice is necessary to avoid dissipating research efforts towards projects doomed to failure.

Parler d'avenir a propos d'une nouvelle technique est chose doublement dangereuse. Car a se montrer trop prudent, on risque de paraitre retrograde et retardataire et en cas de reussite d'etre cite plus tard a cote de ceux qui ne croyaient pas au chemin de fer, ou a l'electricite, ou a l'automobile. A se montrer trop optimiste, on peut passer pour un illumine ou un auteur de science-fiction en mal d'editeur.

A vrai dire le dernier risque est beaucoup moins grave que le premier et c'est celui que je vais choisir a condition toutefois que mon auditoire veuille bien ne pas prendre tres au serieux mes propos etant donne le caractere encore tres embryonnaire de la cryoelectrotechnique.

La caracteristique commune a toutes les machines cryoelectriques est evidemment la mise en ceuvre de tres basses temperatures. Comme dans toute nouvelle technique mettant en oeuvre des temperatures extremes, l'ingenieur doit s'efforcer de reduire les echanges de chaleur parietaux en valeur relative, ce qui peut etre obtenu soit par le choix d'une forme aussi ramassee que possible, soit par I'adoption d'une echelle convenable, ces deux moyens pouvant etre employes simultanement.

Paradoxalement, il semble — d'apres le volume des travaux qui leur sont consacres au cours de ce symposium — que les organes sur lesquels 1'efTort principal est porte actuellement soient ceux pour lesquels le facteur de forme est le plus mauvais possible, a savoir les cables de transmission d'energie. Aussi n'echappent-ils pas aux exigences de l'effet d'echelle et les projets presented au coursdece symposiummontrent nettement qu'un certain gigantisme est necessaire : dans les conditions economiques actuelles une cryoliaison ne serait comparable aux cables classiques a haute tension qu'a partir d'une puissance transitee de plusieurs GW, 3 a 6 selon le type de cryoliaison choisi.

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Ce caractere place d'emblee le developpement possible des cryoliaisons dans des perspectives tres lointaines, disons pour fixer les idees a echeance d'une vingtaine d'annees. Que seront les grands reseaux de transport d'energie a ce moment la? Quelles seront les possibilites et peut-etre les necessites d'insertion des cryoliaisons dans les reseaux? C'est ce que nous allons examiner sur l'exemple francais, etant sous-entendu que les conditions seront quelque peu analogues sur les autres reseaux nationaux de l'Europe occidentale.

La region parisienne consomme actuellement 3,3 GW et consommera vers la fin du siecle plus de 40 GW; ce sera de loin le plus grand centre de consommation en France. Les moyens de production seront concentres sur les grands fleuves (Rhin, Rhone, Basse Seine, etc.) et sur les cotes de la Manche principalement, pour une raison de disponibilite d'eau de refroidissement. Le reseau francais sera done carac-terise par le rejet en peripherie des moyens de production et une importante concen­tration de consommation dans la region parisienne.

Le reseau destine a relier la consommation a la production peut etre envisage de la maniere suivante, en supposant que seuls les moyens classiques actuels ou leur extrapolation a tension plus elevee soient disponibles vers la fin du siecle. Pour satisfaire les besoins de la region parisienne il faut faire converger 10 a 15 lignes aeriennes a 380 ou 750 kV sur une grande boucle de diametre 100 a 120 km qui evite le transit national a travers la region parisienne elle-meme. A partir de cette boucle, une seconde boucle plus rapprochee du centre est alimentee par des lignes a tres haute tension qui sont encore interconnectees. Mais a l'interieur de cette seconde boucle, des lignes multiples a 220 kV et a moyenne tension alimentent le reseau de distribution lui-meme au moyen de structures arborescentes, non maillees, qui ont l'avantage de reduire la puissance de court circuit. La securite de fonctionnement en cas de defaut est assuree par la mise en paralleles de plusieurs lignes a 220 kV (jusqu'a 4). Cela pose des problemes d'acquisition des emplacements de postes de transformation et des couloirs d'energie, problemes dont l 'E.D.F. se preoccupe des a present.

C'est a ce niveau que les cryoliaisons pourraient apporter des solutions seduisantes par leur compacite, une seule cryoliaison pouvant remplacer plusieurs lignes a 220 kV. Cependant, du fait de l'incapacite des cryoliaisons a supporter des surcharges, une structure arborescente ne saurait etre envisagee mais on peut concevoir une troisieme boucle en cryoliaison situee dans la zone suburbaine, boucle alimentee par plusieurs troncons rayonnants, de sorte que la mise hors circuit d'une section quelconque de cet ensemble n'affecte pas une zone de consommation etendue. On pourrait envisager egalement 1'insertion d'une cryoliaison a grande capacite de transport entre le centre de production de la Basse Seine et la region parisienne, le long de la Vallee de la Seine en aval de Paris, region qui vers la fin du siecle accusera un developpement industriel a haute densite.

II convient de souligner ici 1'importance et la specificite du probleme des protections en matiere de cryoliaisons et tout specialement des cryoliaisons supraconductrices a courant alternatif qui sont plus sensibles a 1'apparition de defauts sur le reseau du fait de l'absence des organes de commutation alternatif-continu et de lissage du courant. Cette sensibilite des cryocables amenera sans doute a repenser 1'ensemble des doctrines de protection des reseaux dans le sens d'une plus grande rapidite d'action en cas de defaut. C'est d'ailleurs dans ce sens qu'evoluent les doctrines des a present — et tout a fait independamment des considerations cryoelectrotechniques — puisque Ton est passe en France dans les 5 dernieres annees de 0,3 s ( l e r stade) et 0,7 s (2e stade) a des rapidites d'action de 0,15 s ( l e r stade) et 0,3 s (2e stade). Bien sur, Ton est encore tres loin de la milliseconde requise mais il faut signaler certaines etudes qui visent a declencher les protections non pas a partir des grandeurs integrees de courant et de tension (sinusoidales) mais a partir de leurs valeurs instantanees

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traitees par des calculateurs analogiques et numeriques, ceci en vue d'un accroissement important de la rapidite de decision et d'action.

Au regard des considerations qui precedent, les estimations economiques que Ton peut faire actuellement nous semblent etre de moindre importance et avoir surtout une valeur de comparaison entre solutions voisines. Un accord semble se faire a present sur le fait que les cables simplement refroidis (appeles aussi hyperconducteurs ou cryoresistants) ne semblent pas interessants compares aux cables supraconducteurs proprement dits, ni en courant continu, ni en courant alternatif. Par contre, Tissue de la competition entre cables supraconducteurs en courant continu et cables supra­conducteurs en courant alternatif semble plus incertaine et les avis sont partages. L'etude des pertes en alternatif par deplacement des vortex revet a cet egard une particuliere importance et il est bon de souligner que nous ne devons pas craindre un retour vers des etudes assez fondamentales si nous voulons maitriser ce phenomene encore peu connu. Ce sujet d'etude est relativement meprise par les physiciens du solide specialistes de la supraconduction, parce que portant sur des phenomenes submacroscopiques (tout comme naguere la turbulence a ete longtemps delaissee apres les resultats eclatants obtenus par la cinetique des gaz). L'enjeu en ce qui concerne les applications cryoelectrotechniques est suffisamment important pour qu'un effort soit fait sur le plan fondamental.

En effet, les phenomenes de dissipation d'energie aux tres basses temperatures ont un caractere tres contraignant — vous le savez — du fait de la faiblesse du rende-ment des machines cryogeniques : 1/300 a 4,2 °K actuellement pour une installation de quelque importance (1 kW a basse temperature par exemple). Ce rendement est tres inferieur au rendement theorique de Carnot et laisse un champ de recherches particulierement interessant si Ton songe au prix capitalise des pertes permanentes sur le reseau H.T. (actuellement 5,1 F/W en France). Nous pensons que des recherches vers de nouveaux procedes cryogeniques tels que celui du professeur DAUNT (Cf. Com­munication au Congres de Madrid, aout 1967) pourraient etre particulierement fecondes. Dans cet ordre d'idees mais sur un plan plus immediatement technologique des recherches interessantes sont deja entreprises et devront etre approfondies dans le sens de l'optimisation de la temperature de fonctionnement des cryoliaisons etant donne le gain en rendement des que Ton se con ten te de temperatures de 8 ou 12°K (regime hypercritique de l'helium).

Nous avons parle jusqu'a present de reseaux de transport. Beaucoup de conside­rations developpees ont une portee generate et s'appliquent aux autres secteurs de l'industrie de l'energie electrique. Cependant je ne voudrais pas quitter le domaine du transport de l'energie electrique sans mentionner une certaine analogie avec celui des transports, j'entends ceux des passagers. En effet, tout comme les tensions sont echelonnees suivant les distances a parcour ir, les moyens de transports sont appropries suivant les distances a couvrir, et Ton a remarque que le rapport de la vitesse du moyen de transport a sa distance optimale de fonctionnement est sensiblement constant depuis la marche a pied jusqu'au vol supersonique en passant par l'auto-mobile ou le chemin de fer. Ici encore, il semble que la cry oelectro technique puisse apporter des solutions seduisantes de penetration dans les zones a urbanisation dense, c'est tout au moins ce que laisse entrevoir un des rapports presentes a ce Congres concernant les vehicules a levitation magnetique a grande vitesse par bobines supraconductrices.

Mais revenons aux autres organes cryoelectriques que Ton peut imaginer dans l'industrie de l'energie electrique. Depuis quelques annees les problemes de pointe de consommation — et plus specialement de pointe quotidienne — deviennent preoccupants etant donne d'une part que les sites permettant 1'accumulation hydrau-lique sont limites a terme, d'autre part que les centrales nucleaires sont interessantes en marche permanente du fait de la lourdeur des investissements par rapport aux

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charges proportionnelles a la production. Dans ces conditions on a songe a Faccu-mulation d'energie electrique sous forme d'energie magnetique au moyen d'un bobinage supraconducteur. II faut remarquer d'emblee qu'un nouvel effet d'echelle apparait ici, du au fait que l'energie magnetique volumique est emprisonnee par des nappes de courants superficiels. Ce nouvel effet d'echelle constitue une cause supple-mentaire de gigantisme qui ne rend concevable un tel projet qu'a Fechelle d'un grand reseau national.

Outre le handicap du gigantisme congenital, ce projet a contre lui — il faut bien l'avouer — le fait que les problemes de pointe peuvent se resoudre naturellement au cours des prochaines decennies par les progres en matiere de production nucleaire aussi paradoxal que cela puisse paraitre. Si le prix de l'energie electrique d'origine nucleaire decroit fortement de nombreuses centrales thermiques classiques deviendront obsoletes et pourront passer en reserve pour la pointe. Devant cette possibility, on hesite a promouvoir des turbines a gaz qui sont pourtant de faible puissance unitaire et, a fortiori, l'accumulation magnetique avec tout ce que ce projet comporterait d'incertitudes s'imposerait tres difricilement.

Les machines tournantes peuvent beneficier des progres de la cryoelectricite. En effet revolution constante des grandes machines tournantes est dans Faugmentation des puissances unitaires. Cette evolution dans les perspectives classiques trouvera un terme certain que permettrait de franchir la realisation de grands inducteurs supraconducteurs a courant permanent. Les prototypes realises actuellement (gene­ratrices et moteurs homopolaires ou autres) sont encore tres modestes et laissent de nombreux problemes non resolus tels que ceux des balais, des efforts electro-magnetiques sur les enroulements, de la faiblesse des tensions, etc... Mais ils ont pour eux une certaine simplicity et surtout le sursis a la limitation des puissances unitaires, avantages qui permettent d'augurer une evolution favorable de ces machines a echeance lointaine.

Les divers projets dont nous avons parle jusqu'ici ne font intervenir que les qualites de conducteurs parfaits des supraconducteurs, les phenomenes de transition entre etat normal et etat supraconducteur n'intervenant que comme une gene et une limitation et le phenomene d'expulsion du champ magnetique n'etant pas utilise en tant que tel. On a songe depuis longtemps a utiliser ces phenomenes dans des machines telles que les pompes a flux et les redresseurs cryoelectriques, organes qui pourraient completer harmonieusement des reseaux de transport cryoelectriques. Mais ces projets se heurtent actuellement a de grosses difficultes technologiques inherentes aux pertes Joule dans les conducteurs quand ils sont a Fetat normal. Le probleme fondamental est un probleme de materiaux ou d'agencement de materiaux (effet JOSEPHSON, couches minces, etc. .) a grand nombre de merite (produit de resistivite a Fetat normal par le carre du courant critique) et tant que ces problemes ne sont pas resolus, on ne peut augurer de Favenir de cette classe de machines.

Nous avons examine, d'ailleurs trop sommairement, diverses possibilites d'appli-cations cryoelectrotechniques. Nous voudrions conclure par quelques considerations de portee tres generate concernant Involution des especes technologiques.

Dans Fordre biologique, les nouvelles especes naissent par apparition brusque d'un nouvel outil ou d'un nouvel organe. La mutation se produit de preference sur une espece de petite taille. Aussitot apres, il apparait une explosion de mutations secondaires sur des especes voisines ou le nouvel outil est employe a tort et a travers, souvent de facon grotesque, donnant ainsi de nombreux phyllums avortes. Seul se developpent quelques phyllums pour lesquels Fensemble de Findividu est oriente vers Futilisation exclusive du nouvel outil. Alors, la nouvelle espece grandit et s'eteint, generalement par gigantisme.

II en va de meme de Involution des especes technologiques. La mutation introduite

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par la cryoelectricite va certainement dans le sens de 1'evolution generate de Telectro-technique puisqu'elle permet une plus grande compacite de nombreux appareils. Mais il faut se metier des applications intempestives et choisir seulement quelques applications ou la supraconductivite soit utilisee pleinement et d'une maniere simple. II convient egalement d'eviter les projets necessairement gigantesques. A ce titre, il semble que des applications telles que celles des electro-aimants supraconducteurs constituent un relais tres utile dans la longue marche qui reste encore avant des applications vraiment industrielles.

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INSTITUT INTERNATIONAL DU FROID INTERNATIONAL INSTITUTE OF REFRIGERATION

177, boulevard Malesherbes, 75-Paris-17e

PUBLICATIONS Comptes rendus des reunions des Commissions Internationales

Proceedings of the meetings of International Commissions

SCIENCE ET TECHNIQUE DU FROID REFRIGERATION SCIENCE AND TECHNOLOGY

— 1960-3, Marseille (France). 1960. « Evaporateurs. Machines a absorption et a ejec-teurs de vapeur. Preemballage des aliments. Qualite des produits congeles. Refrige­ration des fleurs et bulbes. Entrepots frigorifiques. Portes isothermes et rideaux d'air ». — « Evaporators. Absorption machines and steam jet systems. Prepackaging of foodstuffs. Quality of frozen foods. Cooling of flowers and bulbs. Cold stores. Insulated doors andair-curtains » . 555 p. (45 F - $ 9.00).

— 1961-1, Wageningen (Pays-Bas - The Netherlands). 1961. « Entreposage frigo-rifique des fruits et legumes ». — « Refrigerated storage of fruits and vegetables ». 375 p. (35 F - $ 7.00).

— 1961-3, Londres-London (Royaume-Uni - United Kingdom). 1961.« Enseignement du froid au niveau moyen ». — « Training in refrigeration at medium level» — Cambridge (Royaume-Uni - United Kingdom). 1961. « Isolation thermique — Instrumentation — Transmission de chaleur — Compresseurs — Applications industrielles du froid — Transport maritime et materiel» — « Instrumentation — Heat transfer —Refrigerating compressors — Some industrial applications of refrigeration — Marine transport and equipment». 555 p. (40 F - $ 8.00).

— 1961-4, Belgrade (Yougoslavie - Yugoslavia). 1961. « Progres recents en cryobio-logie ». — « Advancement in cryobiology ». 102 p. (10 F - $ 2.00).

— 1962-1, Washington, D.C. (U.S.A.). 1962. « Conductivity thermique. Refroidisse-ment thermo-electrique. Compresseurs. Conditionnement d'air. Aliments conge­les. Organisation de la recherche ». — « Heat conductivity. Thermoelectric cooling. Compressors. Air conditioning. Frozen foods. Organisation of research ». 685 o. (50 F -$ 10.00).

— 1962-2, Santiago de Compostela (Espagne - Spain). 1962. « Entrepots et transports frigorifiques ». — « Cold stores andrefrigeratedtransport». 21 Op. (20 F - $ 4.00).

— 1962-3, « Groupe de Travail». — « Working Party ». — « Physiologie, qualite et transport des bananes ». — « Physiology, quality and transportation of bananas ». 120p. (15 F -13 .00 ) .

— 1964-1, Dublin (Irlande - Ireland). 1964. « Entrepots frigorifiques » — « Refriger­ated warehouses ». 262 p. (25 F - $ 5.00).

— 1964-2, Turin (Italie - Italy). 1964. « Conductivity et diffusivite thermique. Trans-fert de chaleur et de masse. Echangeurs. Fluides frigorigenes». — « Thermal conductivity and diffusivity. Heat and mass transfer. Heat exchangers. Refrigerants ». 190p.(20F-$4.50).

— 1964-3, Abidjan (Cote-dTvoire - Ivory Coast). 1964. « Les applications du froid en pays tropical». — « Refrigeration applications in tropical countries». 455 p. (40 F - $ 8.00).

— 1965-1, Karlsruhe (Allemagne - Germany). 1965. « Viande preemballee. (Eufs. Volaille ». — « Prepacked meat. Eggs. Poultry ». 262 p. (25 F - $ 5.00).

— 1965-3, Cracovie - Cracow (Pologne - Poland). 1965. «Vehicules frigorifiques : construction, isolation, moyens de refroidissement». — « Refrigerated vehicles: construction, insulation and refrigerating equipment». 210 p. (20 F - $ 4.00).

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— 1965-4, Tchecoslovaquie - Czechoslovakia. 1965. « Materiel frigorifique dans le conditionnement d'air et l'industrie chimique; le froid en medecine et biologie». — « Refrigerating equipment in air conditioning and the chemical industry; refrige­ration in medicine and biology ». 700 p. (50 F - $ 10.00).

— 1965-5, Suede -Sweden. 1965. « Installations frigorifiques marines ». — « Marine refrigeration ». 380 p. (35 F - $ 7.00).

— 1966-1, Bologne (Italie - Italy). 1966. « Conservation des fruits et legumes (notam-ment en atmosphere controlee). Entrepots frigorifiques ». — « Storage of fruit and vegetables (including in controlled atmosphere). Cold stores ». 670 p. (50 F - $ 10.00).

— 1966-2, Trondheim (Norvege - Norway). 1966. « Etudes sur le transfert de chaleur dans le domaine du froid (isolants; echangeurs de chaleur)». — « Studies on heat transfer in refrigeration (insulants; heat exchangers)». 320 p. (50 F - $ 10.00).

— 1966-3, Delft (Pays-Bas - The Netherlands). 1966. « Concentration et purification par congelation ». — « Concentration and purification by freezing » . 200 p. (25 F -$5.00).

— 1966-4, Londres- London (Royaume-Uni - United Kingdom). 1966. « Vehicules frigorifiques : entrees de chaleur dans les vehicules isoles; economie des moyens de refroidissement». — « Refrigerated vehicles: evaluating the heat loss from insulated vehicles; economics of refrigerating systems ». 270 p. (30 F - $ 6.00).

— 1966-6, Athenes - Athens (Grece - Greece). 1966. « Le froid dans les pays mediter­raneans (applications agricoles et alimentaires)». — « Refrigeration in Mediterra­nean countries (applications to foodstuffs)». 390 p. (40 F - $ 8.00).

— 1968-1, Avignon (France), 1968. «Chambres froides a atmosphere controlee — Entreposage des produits surgeles» — «Controlled atmosphere cold rooms — Storage of quick-frozen products » . 144 p. (25 F - $ 5.00)

CRYOGENIE PURE ET APPLIQUEE PURE AND APPLIED CRYOGENICS

— Vol. 2, Eindhoven (Pays-Bas - The Netherlands). 1960. « Methodes de refrigera­tion, de liquefaction et de separation aux temperatures basses et tres basses. Pro-prietes thermodynamiques des fiuides, des melanges et des solides aux basses temperatures. Applications des basses temperatures ( p. ex. masers, supraconduc-teurs, champs magnetiques tres intenses, chambres a bulles)». — «Refriger­ation, liquefaction and separation processes at low and very low temperatures. Thermodynamical properties of fluids, mixtures and solids at low temperatures. Application of low temperatures (e.g. masers, superconductors, high magnetic fields, bubble-chambers) ». 298 p. (30 F - $ 6.00).

— Vol. 3, Londres - London (Royaume-Uni - United Kingdom). 1961. « Methodes de production a grande echelle de gaz liquefies, conservation, transport — Metho­des pour la production et la mesure des temperatures inferieures a 1°K». — « Methods for the large scale production of liquefied gases, their storage and trans­port — Methods for the production and measurement of temperatures below 1 °K ». 196p.(25F-$5.00).

— Vol. 4, Grenoble (France). 1965. « Transfert de chaleur au-dessous de 100°K et ses applications techniques ». — « Heat flow below 100°K and its technological applications ». 365 p. (40 F - $ 8.00).

— Vol. 5, Grenoble (France). 1965. « Hydrogene liquide; proprietes; production et applications». — « Liquid hydrogen; properties; production and applications». 395 p. (50 F - $ 10.00).

— Vol. 6, Boulder, Colo. (U.S.A.). 1966. «Technologie de Thelium liquide». — « Liquid helium technology ». 555 p. (60 F - $ 12.00).

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Congres Internationaux — International Congresses

— Comptes rendus du 9e Congres, Paris, septembre 1955. Publication en fascicules separes des travaux des differentes Commissions. — Proceedings of the 9th Congress, Paris, September 1955. Separate booklets of the work of each Commission. 1500 p. (de 20 a 30 F — from $ 4.50 to $ 7.50).

— Comptes rendus du 10e Congres, Copenhague, aout 1959. Publication en 3 volumes des travaux des differentes Commissions. — Proceedings of the 10th Congress, Copen­hagen, August 1959. Three volumes of the work of the Commissions. 1680 p. (3 vol.: 275 F - $ 55.00).

— Comptes rendus du l l e Congres, Munich, aout-sept. 1963. Publication en 3 vol. des travaux des differentes Commissions. — Proceedings of the 11th Congress, Munich, Aug.-Sept. 1963. Three volumes of the work of the Commissions. 2000 p. (3 vol. : 260 F - $ 52.00).

— Comptes rendus du 12e Congres, Madrid, 1967. Publication en 4 volumes des tra­vaux des differentes Commissions. — Proceedings of the 12th Congress, Madrid, 1967. Four volumes of the work of the Commissions. 3700 p. (1 vol. : 90 F - $ 1800 ; 4 vol. : 280 F - $ 56.00).

Publications Diverses — Miscellaneous Publications

Recommandations pour un Code International d'Essais des Machines Frigorifiques l r e ed. — Recommendations for an International Code for Refrigerating Machines 1st ed. 1957. 25 p. (3 F - $ 0.75).

Catalogue de films et bandes fixes relatifs au froid. — Catalogue of films and strips on refrigeration. Part I and II. 160 p. (7 F - $ 1.50).

Conditions Recommandees pour le Transport Terrestre des Denrees Perissables. — Recommended Conditions for Land Transport of Perishable Foodstuffs. Commis­sions 4 et 7.1963. 11 p., tabl. (5 F - $ 1.00).

Recommandations pour la Preparation et la Distribution des Aliments Congeles. —Recommendations for the Processing and Handling of Frozen Foods. 1964. 123 p. (15 F-$3 .00) .

Regies pour Machines Frigorifiques. 1960. Traduction en francais de la 5e edition de « Kaltemaschinen Regeln ». 152 p., diag. (35 F - $ 7.00).

Dictionnaire International du Froid (anglais, francais, allemand, russe, espagnol, italien). 1600 expressions, dont 400 avec definition. 300 pages environ 18 x 27 cm, sous couverture cartonnee. — International Dictionary of Refrigeration {English, French, German, Russian, Spanish, Italian). 1,600 expressions, 400 of which are defined, 300 pages 18 x 27 cm, under stiff folders (60 F - $ 12.00).

Les Techniques Frigorifiques dans les Pays Chauds en Voie de Developpement. — Refrigeration Techniques in Developing Countries. 1964. 116p . (12F- $2.50).

Guide Bibliographique du Froid. — Bibliographic Guide to Refrigeration 1953-1960 10000 ref. (100 F - $ 20) 1961-1964. 6000 ref. (100 F - $ 20) 1965-1968. 7000 ref. (150 F - $ 33).

Guide Pratique de l'Entreposage Frigorifique. — Practical Guide to Refrigerated Storage. 1965. 240 p. (40 F - $ 8.00).

Conditions Recommandees pour l'Entreposage Frigorifique des Produits Perissables. — Recommended Conditions for Cold Storage of Perishable Produce. 1967. 100 p. (12 F -$2 .50 ) .

Essais d'lsolants Thermiques. — Thermal Insulating Measurements. 1968. 33 p. ( 8 F - $ 1 . 6 0 ) .

Projet de code de pratiques pour le poisson surgele. — Draft code of prat ice for frozen fish. 1969. 75 p. (7 F - $ 1.80).

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(B 868) Imprime' en Belgique par Ceuterick s.a. Brusselse straat 153 3000-Louvain Adm.-dir. L. Pitsi Bertemse baan 25 3008-Veltem-Beisem