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APPLICATIONS OF AMERICAN DESIGN CODES

FOR ELEVATED TEMPERATURE ENVIRONMENT

L. K. Severud

March 1980

...---------DISCLAIMER --------1 k nsored by an agency of the United States Government.

This boOk was prepa;ed as an accoul'll 0':::an spo11gencv thereof, nor any of their employees, makes anv Neither the United .... 10111s C":'vernment mes vanv leycil liobilitV nr rewonsibilitV for t~ acoJracv. warranty, express or lmphed, Of assu. f rmation apparatus. product. or process d1sclv:.!d .. ~r comple1eness. ~ uselulness of a~:fri in o privatel~ owned righu .. Aelerence herein to anv_ spec1f1c

represen1s 1hat 1U. use y,()\Jld not . ":; 1rade name. trademark, manufacturer, .Of otherWise, ~ commercial product, ~ocess, o; serv1~ V d rsement, recommendation, or favoring by th.e United not necessarily constitute Of imp:~~! e;h: views and opinions ot authors expressed herein do not

~:~::;::::' ,: 1;;~h°!:: the. Uni;ed States GovernrT'llil•l ur onv flYN"V thereof.

To be presented at th~ Intern ational Conf erence, "Engineeri n g Aspec ts o f Creep" at the University o f Sheffield , U. K. in Sept. 15-19 , 1980.

HANFORD ENGINEERING DEVELOPMENT LABORATORY Operated by Westinghouse Hanford Company, a subsid iary of Westingho11c;p Electric Corporation. under the Department of

Energy Contract No. DE-AC14- /6r f 02 1 70

COPYRIGHT LICENSE NOTICF By i<Upl1 nco ol th is 1rt1cl t . lh t Publish er 1nd t o1 1tc 1pie nl 1cknowlr d&" lht U.S Ga.trnmtnf s 111h1 lo nl11n 1 nonu cl umt . 1oy1il1·h tt hc tnst 1n i nd lo 1ny copy111hl

U1Tt•inc lh11 p1p<1.

DISTRIBUTION OF THIS GOCUME:ir IS UHL~

DISCLAIMER

This report was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency Thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof.

DISCLAIMER

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L. K. Severud, BSc, MSc, PE Manager · Of Plant Analysis, FFTF Project W.:stinghouse Hanford Co. Richland, WA . U. S. A. 1 Applications of Amer ican Design Codes For Elevated Temperature Environment

SYNOPSIS

This paper first provides a brief summary of the ASME Code rules of Case N-47. Then an overview of the typical procedure and analysis ingredients used to demonstrate Code compliance is provided. Application experience and some examples of detailed inelastic analysis and simplified-approximate methods a r e given. Recent developments and fu ture trends in design criteria and ASME Code rules a re also presented.

INTRODUCTION

1. Many of the Liquid-Metal Fast Breeder (LMFBR) componenu, upt:.i:atc at rPmpPr11tu;r;e~ and durations where metal behavior is time dependent . Although design and operation of components in the creep range is not new to aircraft and rock­et engint:.s, steam turbinP.s, etc ., past American nuclear power plant designs have all been for operation in the subcreep regime . Moreover, components that have been developed for operating in the creep regimes have usually evolved gradually over many years of testing and oper­ating experience. In other cases, they had characteristics such as short lifetime require­ments, small size, easy inspection and mainten­ance, etc. , so that they could be developed economically am! in a reasonahly short time using mainly trial and error testing. Since the LMFBR nuclears plants and their component s must oper­ate safely and r eliably for 20 to 40 years , in­service inspection is much more difficult; and as the stress analysis methods for µ1eJicting in­elo·sLlc time depend ent rP .sponse have greatly advanced in the past half-decade, more emphasis is being plar.ed upon analytical design tech­niques . 2. Assurance of structural integrity for systems subjected to sustained operation in the creep range and significant thermal transients requires well-qualified analytical techniques to predict the time-dependent structural behavior. uf ..:.omplci~ aomponPnr r.onfiii;1,1rations. Detailed inelastic analysis procedures, although time con­suming and expensive, are required to accurately predlcL structural response. Simplified methods are needed for use in design, both in the early stages for scoping and in the final stages under conditions where design margins are inherently larg9 nngning coorginated test programs are needed to support both the development dnd verificatiun cf analysis procedures. 3. The American industry and the U. S. Dep~rtment of Energy has had a long program of LMFBR development. Structural desigu cx iterin have bee n rleveloped and published through ASME committee activities. This included participa­tion from the U. S. industry, U. S. DOE owned laboratories (i .e., Oak Ridge National Labora­tory, Argonne National Laboratory, Hanford

Engineering Development Laboratory, Energy Tech­nology Engineering Center) and the U. S. Nuclear l<egulaLuLy Commisoion nPsign m~thods and base technology data have been developed pr imarily through the the conduct of the major LMFBR pro­jects and the U. S. DOE base technology programs. The first major U. s. LMFBR project was the Fast Flux Test Facility (FFTF) built six miles (lO kilometers) north of Richland, Washington . The Fast Flux Test Facility (FFTF) is a primary test facility for the U. S. LMFBR program. Although it does not have electrical power generation components, it has many LMFBR characteristics and its design and construction has provided a valu­able foundation for the U. S. LMFBR program. The second major U. S. LMFBR project is the Clinch River Breeder Reactor (CRRR) project. Detailed technical data, information and reports on pro­ject component designs, analysis methods, tests and experiences from the projects and the base technology programs are available through the U. S. DOE Exchange Program. 4 . In dcsignj,ng r.omponents for LMFBR nuclear power plants, the design analyst is faced with the task of predicting the component strengths considering effects of creep and elevated-tem­perature environments. Both time-independent and time-dependent failure modes are considered by performing stress analyses to satisfy the rules of the ASME Boiler and Pressure Vessel Code, Section III, Nuclear Power Plant Components. At thi:i 'n-iting, C:ode Cg.se N- 47 (1592) sets forth the applicable ASME rules and limits tor SecLluu III Class 1 components in elevated-temperature ~ervicc (ref . 1). The structural failure modes considered in Case N- 47 (1592) are:

1 . ductile rupture from short-term loading£, 2. creep rupture from long-term loadings, 3. creep~faLlgu~ foiluro , 4. gross distortion due to incremental

collapse and ratcheting, 5. loss of function due to excessive

deformation, 6. buckling due to short- term loadings, 7. creep buckling Jue to long-term loadings.

5. Rules and analysis methods with application experience for design to guard against ' failure

L., ,K. Se'.verud

modes 3, creep-fatigue, and 4, ratcheting are addressed by this paper.

ASME CODE RULES

6. The ASME Code rules and limits for creep-fatigue and cumulative inelastic strain set forth in Case N-47 (159 2) are of major signifi­cance to design as they of ten require detailed and costly inelastic analyses for compliance demonstrations.. Much effort has and continues to be given to develop simplified and more accurate methods, with appropriate rules and limits that maintain the conservatism needed to assure design inte~rity, plant and public safety.

7. The present Code rules are specified in the framework depicted by the flow diagram of Figure 1. Limits are placed on load-controlled stress, and strain and deformation. The strain and de­formation limits, in the non-mandatory Appendix T of Case N-47, have been used in the U. S. LMFBR projects and most design-analysis difficulties have been associated with the creep-fatigue, strain limits, and buckling requirements.

Creep-fatigue rules and procedures 7.1 The Code rules have the following ingredi­ents:

1. Stress and strain, the number of cy­cles, and time durations are the primary parame­ters tha t are used to predict and control creep­fatigue damage.

. 2. A semilinear cumulative creep and fatigue damage assessment is used, with cycle and time fractions employed for damage counters. The allowable cumulative damage limit D is a function of the cycle or t ime-fraction sum. The equation, . Eq. (5) of (ref. 1), is

f (~) + t (_.:) 2 D, j=l Nd j k=i Ta k

(1)

where

D

n

total creep-fatigue damage (see Fig. Z, which is T-1420-2 of CC 1592), number of applied cycles of loading condition j, number of design-allowable cycles of loading condition j from the fatigue curves corresponding to the maximum metal temperature during the cycles for the equivalent strain range, ' time duration load condition k,

= allowabl~ time at ~ given ~rress inten­sity (for elastic analysis) or at a given effective stress (for inelastic analysis) from load k; Td values are obtained by entering the stress-to­rupture curve at a stress value equal to the calculated stress (from load k) dlvl<le<l Uf the factor Kl= 0 .9.

3. An equivalenr strain range approach to multiaxial stress-strain effects is used. Rules for determining the equivalent strain range are intended to be applicable whether µrluLipal stra:f,ns change directions during the C) :le or not. The equivalent ct:rf in range i.s

6E . _Vz (6E 6E ) 2 + (6E 6E ) 2 (2) equiv -3 x - y y - z

Jl/2

+ (6E - 6E ) 2 + 6 (6E 2 + 6E 2 + flE 2) z x xy yz zx

2

which is Eq. (6) from Code Case N- 47 .

4. Different evaluation procedures are em­ployed, depending on whether elastic or inelastic analysis was used. The creep-fatigue calculation procedures for Code Case N-47 are presen t ed bv Campbell (ref. 2), and the background for the present creep-fatigue code rules is discussed in detail in the ASME criteria background publica·· tion (ref. 3). Recent modificat i ons to the elastic analysis rules were described by Severud (ref. 4).

Fig. 1.

OE SIGN LIMITS

llVELC SH,VK:I

r---r.--:;:.OMITS ~-======--i f • : ~ ,...-.._,

LEVEL 0 SfRVICI

I J~»~··).r) ~~;;f~ '

IN llt-MAN ATOll!Y

ltOLtYITSUMl..USl"Kl"fO IPfTIOfCCSMONSH.CIFICAJIOfol

illl!!P r) COO.fllOU.. lOou.&11 n h '--"' -h.AJo ToC 4'0"-TJlS

\..J =~~~:::~,~ 0 COioH'UflDOUA,.flh

Flow diagram for elevated temperature analysis (ref. 1). Reproduced by permission of the ASME.

- I~ o.6 w 0. 4

0.2

0:-~-:-...,...:--:1-:-~'---""' 0 0.2 0.4 0.6

l:,ij

Fig. 2. Code Case N-47 creep-fatigue damage ~uvelopc (ref. 1).

Ratcheting and cumulative inelastic strain limits 7.2 T~~ r.nrlP places the following limits on the maximum accumulated inelastic strain for patent material (see T-1310 of Case N-47):

•Strain averaged through the thickness, 17. •Strain al Ll1~ surfaces, du2 ttJ "n equivalent

linear distribution of strain through the th i.ckness, 27.

• Maximum local strain, 5%

Inelastic str ains accumulated in we ld regions are computed using parent material properties and

L. K. Severud

these calculated strains are limited to one-half the strain values permitted for the parent material.

Buckling and instability rules 7.3 The Code has both time- independent and time-depend ent rules. The limits and design (actors are placed on load- controlled and stra in­controlled buckling as shown in Table 1.

Table 1

Buckl ing limits and factors

Table T- 1521- 1 ( Case N- 4 7) •rime- 1.ude pcndent Buckling Limits

Load Strain

~ Factor

Des i gn Conditions 3. 0 1.67 Operating Conditions 3 . 0

Normal 3.0 1.67 Upset 3 .0 1.67 Emergency 2.5 1.4 Fault ed 1.5 1.1 Testing 2.25 1.67

Tabl e T- 1522- 1 Time- Dependent Load- Controlled Buckling

Factor s

Operating <.:orH11 1..luu.J

Normal Upset Emer gency FH.ultcd

1. 5 1. 5 1.5 1.25

The load or strain factor is equal to the loatl (strain) which would cause instant ins tability divided by the expected load (strain). To pro­tect against load-controlled time-dependent creep buckling instability , the design has to be demon­strated capable of withstanding during the specified lifetime a load history obtained by multiplying the specified Operating condition loads by the factors given in Table T-1522-1. The background to these values is given by Berman, Gangadharan, and Gupta (ref. 5).

CHARACTERISTICS OF LMFBR COMPONENT DESIGNS

8. The temperatur~~ f or the CRRR and recent American LMFBR systems are abou t 50°F lower th an for FFTF.

Typical design and environment conditions and their bounds for LMFBR components that are designed to satisfy the ASME Code are described in Table 2.

Table 2 -----Typical bounds on LMFBR design and environment conditions

Mater i al s of interest 304 and 316 stainless steels; Hi - Fe-Cr alloy 800 . grade 2 ~ 2 1/4 Cr-1 Mo

El'"dr.nnm.-nt,a Air, sodium , gaseous nitrog:eo, l"IUulll.Ji •• ..,.,, , .!-!~!.."II. rll\K , Ar, He

Temperatures Moa t1y 1000°:' - l 050°F or belov (but some vi t h 1200°F &nd a. !'ew 15oo•n

Design life 20 to 30 years Thenna.l. transient events About 1000 tot&l cycles; usuall y

about 20 to 50 cycles vith - 10°F/ sec and a 6'1' or ~00.,F , the rec t nf the transients usually have rates -4°r/ see .,r less a.nd 6T >300°F .

Strain ra.n!P= Usually about 0.2 to 0.3S or less ; ra.rely greater than 0 . '%

Frequeocy ot strains High t o lov, Crom 1 cps to 1 cycle per f tve y.:ars

!rradiation neutron i rradiation nuence b less t han 1020 neutrona/cm2 ( E >O.l MeV} over the component. lifeti!lle.

3

9. Pressure stresses are usually low, but some systems have sizable average stresses on a section due to gross thermal expansion (i . e., piping systems) and mechanical loads (i.e., nozzle loads induced on ves sels ) . The critical stress or strain locations include those where geometrical discontinuities exi st, l oca t i ons where thermal environment discontinuit i es exi st (i.e., a half- full tank), and locations of t h i ck­ened wall sections (i.e . , tube- to-tubesheet connections in heat exchangers) . The stress or strain field often is characterized as a linear variation (bending type) through the section . Of course, nonlinear variations of stress and strain through the section fr equently occur also due to local structural discontinuities and non-1 i.near thermal gradients.

'IYPICAL PROCEDURES AND ANALYSIS I NGREDIENTS

10. The designs and stress analyses of LMFBR elevated- temperature components are initiated using common elastic analysis methods, and often finite-element computer techniques are employed. Due to the high cost and large· number of man­hours and computer time required f or de tailed inelastic analyses, extra e ff orts are made to allow demonstration of compliance with t he Code design rules using only the less informativ e elaStlt.: a .. .ilyoii;: finning:;;. These efforts i nvolve (1) optimizing component configuratiohs Lu t~Juco discontinuity and peak stress es, (2) removing weld locations from high stress regions, (3) selecting or changing material t ype for improved streneth properties, (4) improving t hermal­hydraulic analyses of sys tem characteris tics t o remove undue conservatisms in magnitudes of pre­dicted thermal shocks and operating temperatures, and (5) using simplified and usually conservative inelastic response approximate methods (ref's . 6, 7, 8). Nonetheless, some components with envelope restrictions and/or with the more severe operating conditions have to be evaluated using detai led inelastic analyses. 11 . Elastic analysis methods have advancetl tremendously in the last decade, largely due to computer solutions utiliz i ng finite-element formulations . Inelastic methods have also advanced in the last ten years tn the extent thaL elaatic~plastic-creep incr emental-load solutions to one- , two- , and s ome three­dimensional structures are now procedurally possible. Reasonable interim guidelines for inelastic analysis material and constitutive equations have been developed Pugh (ref. 9). However, inelastic anal ysis still has short­comings and analy tical limi t a tions. The mat e r ial characterization equations and the required con:Jtitut i ve equations are still in an early stage of development. Although the computer large-core requirements, h i gh computing cost, vast a111ounto of rnmputer printout data, and overall lengthy stress analysis schedules dre being reduced, they still impose a significant deterrent to inelastic analyses. 12. Fnr FFrf components subjected to signifi­cant creep and cyclic load con<llL l u11:i, in <;1 l:>">t ic structural analyses were necessary in order to predict the time- and cycle-dependent stress ~nd strain data necessa r y fo r the creep-fa tigue and ratchet strain analysls. 13. Normal design analysis practice was to combine the less severe transients (such as normal startup and shutdown) with the two most severe transients and analyze the r esulting

L. K. Severud

combination. To account for the total strain histor y, an appropriately increased number of combined transients was used in the creep-f atigue and r a t cheting analyses. 14 . In order to perform the component elevated­temperature fatigue (often referred to as a creep- fatigue interaction) and cumulative strain analyses , the incremental and cumulative stresses and strains (six stress and six strain elastic, plastic, and creep components) must be computed for each load step throughout thermal and mechan­ical load duty specified as the basis for component design.

DETAILED I NELASTIC ANALYSES

15. Mos t designs are based on elastic analysis, screening r ules and simplified inelastic evalua­tions. Also , about fo r t y maj or FFTF component detailed inelastic analyses have been accompli sh­ed, ,primarily for final design confirmation. Some of these inelastic analyses and findings were report ed by Bigelow (ref. 10) in 1975 . References to a number of the papers on these analyses are given (ref's 11- 18). Since 1975, papers on additional FFTF analyses used to demonstrate elevated temperature Code compliance were published (ref . 19- 27) . The number of t hese inelast i c ana l ys es per t ype of component a r c given in Table 3 .

Table 3

Ma j or de tailed inelas tic analys es for FFTF

Type Component

Pi pelines vith approx. elbov elements Pipe components ( elbovs, flue d he ad anchors 7 e t c. ) Heat exchanger Pump Valves Tanks Reactor internals Miscellaneous

Number of Ana lyses

8

9 4 2 4 4 5 4

Costs usually were . $30 ,000 t o $75, 000 with s ome cos t s up to $100,000 for a total cost of about $2,000,000. The time rPrptired to carry out an inelastic analysis usually was in the 3 to 6 months range. Clearly, inelastic analys is is too costly for use in preliminary design and should be used mostly for final design conf i r ­mation and technology development. 16 . To provide an overview of inelastic cycl e and ti~e-dependent ana lysis, some of the analy­sis ingredients and stress- strain response for the FFTF Intermediate Heat Exchanger ( IHX) primary closure seal (re t . 12) will be discussed . Then additional examples that show a variety of stress-$train reoponse are provided. 17 . The IHX is a counte rflow shell and tube type exchanger that transfers heat from the radioactive primary sodium flowing on the shell oide to the secondary sodium flowing inside the tubes . The s truc tural material used in Lite nm is type 304 stainless steel. 18 . The details of t he support regi on are

hnwn in Fig. 3, where the C- shaped cross section of the toroidal shell segment, us ed as ~losurc seal member, can be s een . 19 . The seal , in addition to res i sting the sodium pressure, must be capable of accommoda­ting differenLla l Lhermal movem<m t <> h<>. t ween t he

4

adjacent massive members. Loadings consider ed in evaluating design adequacy were associated with the Ul and U2 temperature transients shown schematically with the pressure histogram in Fi g. 4 . 20. Finite-element methods were us ed in the analysis of the seal ; the f in i te- element idealization is shown in Fig . 5 .

V CENTIEA LINE OF IHX HAN GI NG SU' l'O" T IHS1

SHEAR KEY FO RC ED Jlll NCi CSKJ I

SHEA R ILOCK I

I I

I IUNOl E SU'f'OIU

FLANGE IY FI

I

I I

I I

I

Fig. 3. Primary cl osure sea l (C- s eal) support struc tuce.

JOso<'F f-1s6 HOURs-j

·-7 CREEP ~COOL. f-156 HOURS-j

~ HU U2 D WN

HEAT UP UPS ET D HU Ul 1

HEAT UP UPSET j CO

j I ~ I ONE CYCLE I I ... :,;: 1-t-1------------l1 (~

TI ME-

Fig . 4. Load hi s tory of primary C- s eal for finite- elemen t ana l ysis .

r"""' ""' o• '"'

I OUTSIO( SUlllFACE

Uf"ll f OiltM Sl-l (LL • TMICKNUS OF

~~~~'~F2:~V

Fig. 5 . Fin i te-element idealiza tion of C- seal.

.. _ .. , • .+011 11 0 ..il.c,L

' O t S•l.AC [Ml~ T 1 1 • Vllll fl CA l

OIS .. LAC(lrjl l NT

•,·oo

L. K. Seyerud

21. for most of the elements, the circumferen­tial strain range was the largest and was the major contributor to the computed fatigue damage. However, strains in the meridional direction were not small and were shown to be influenced significantly by yield stress and thickness changes. The stress-strain results for the meridional direction at the inside surface of element 1 (Fig . 5) are shown in Fig. 6 . In this case, the loading histogram consisted of a U2 cycle followed by three combined U2 and Ul cycles (Fig. 6) and, finally, three Ul cycles. The results in Fig. 6 indicate some of the com­plexities of stress-strain responses that can be encountered in practical applications. 22. The stresses and strains for the circum­ferential direction, although not included in this description, must be included in any Code evaluation. The solution gives the principal stresses and principal strains, and, the elastic, plastic, and creep components are calculated for each point in time. 23. Some additional examples of stress-strain response have been selected from the FFTF design to indicate the range of responses observed. However, it should be remembered that the results depicted are usually for one direction only, and rDsvlrs for other directions are needed to define thP. states of stress or at Stralu. 24. The first additional example is thP. stress­strain response of a structure (Wu, ref. 14) installed in the upper plenum of the FFTF reactor vessel to preven t cur e effluent from breaking the surface of the so<lium pool . This structure is required to withstand 120 cycles of normal reactor startup and shutdown, 705 cycles of thermal transients of one magnitude (x transients), and 20 cycles of a second type of thermal transient (y transient) . Thli! y transient produces more stringent loading conditions than the x transient. The temperature gradient through the thickness of the plate is very small during normal startup and shutdown, and the strain range for the associated cycle is very small. Thus, two repetitions of cyclic behavior due to a y transient followed immediately by a less severe x transiPnt were selected for Pxamiqation. The stress-strain results for one of the principal directium, (y) of an Pl P.ment in the plate are shown in Fig. 7. 25. One FFTF pipeline inelastic analysis by Huang (ref. 22) , shown in Figures 8 and 9, resulted in the stress-strain response of Figures 10 and 11. Anot her FFTF pipeljne analysis by Pan (ref. 16) revealed t ypical time­dependent stress relaxation shown in Figure 12. 26. Examples of complex geometry and large finite element models l hat havP been used on a few very expensive inelastic analyses are shown in Figures 13 and 14 .

SIMPLIFIED AND APPROXIMATE METHODS

27. Many simplified and approximate methods were used in tne FFTF design . ~arh has its own limitations and advantages. In general, limitations of approximate methods are due to simplifications in materials or structure models. Thus, the analysi.c; output is approxi­mate and often incomplete and sometimes grossly conservative results occur. Justification for final reliance on the simplified modeling can Le difficulr rn develop. However, the advantages of the approximate methods are that they dre l ow in cost, quick and easy to apply, the methods

Fig. 6.

Fig . 7.

5

" " 20

·•

· 12 ~~-~---'---L---'---'---.L...._..J

MUU0t0NAL STlllA/ N l "'I SIO( SU"fil1Ct C' I

Typical time-dependent stress-strain response curve for a C-seal.

z

~

l' l lllST CYCU OF ! flllANSIENt

l' l llST CYCLE 01' , Till.ANSl(Nf

SECOND CYC LE O~ • tlllANSl£NT

8 , ...... ,~, ...... .._... ....... ..,_~.-.. -,., ~ c~" ST lll AIN COMl"O"I ENJ !.,.Jin I

.· ,, (NO 0' 1"'1 C'f'ClE OF• T,_ANSIENT

e SECONO CYCLE OF y TlllANSIENr

Typical time-dependent str~~~-strain response curves for a structure inside the reac tor vessel.

Fi g . 8. PHL-CLS inelastic analysis finite element model.

L. K. Seve'rud

TIM,llAJUIH U MPE IATURI

121S

1200

lSO

70

Q,4 • Hc:11.1r Ill Ctup Hold o90. I

34.JS 71 104.JOS 156 .157 200 6-tl,9

" Uiil lo rm Hoo rup

IS6-Ho11r Cr• •P Hold

UI Tu1n1ionl ..• Uniform .._tup

156 -Hour Cr oop Hold

9S 14 7

65700· Hout Cr ••P Hold

UllA Tro n1i •nl ... Uniform Hool u p

o . • N<.,mbora oro lho loacl H• p1

b . Con1tanl prenwr o 91 p11(0 .0 27MPo ) h \11od fo r on t l r• lood in9

Fig. 9. PHL-CLS inelastic analysis load histograms .

"

: . / I ""Mb ... "'" ••

load 1lop1

176.7

21 ,1

1l7.9

l4,S

u ;;

- 34.S 1 .!

· U.9 ~

-10.J.4 ~

-137.9

_,, '------~--~---'---'---~--~-172.4 O S 10 IS 'lO 25 JO

S11m of llo1t ic PIHl i c ond Croop Stro int , I0-4 in /i n

Fig . 10 . Meridional stress vs total strain at the inside sur face of el ement 13 of the PHL-CLS .

,. 111.4

'" 137.f -:; .

" I<».• .. . 10 . ... ~ ><.•

.. "'-? :! .. ~ ., J 1-10 H .9

.~ a-u Nvmbw• I ,., .. .. J -KJ:I.• load ll • p • v

•o •ll 7. t

-u_L,-----'.'----'0 ____

4._ __ _,,:----:-:,

2-112.•

St.1m o f llolfic , Ploo i c a nd Creep S t. .. i ;u , 10· 4 i r11 in

Fig. 11. Cirr11mferential s tress vs total strain at the inside surface of element 13 of the PHL- CLS.

Fig . 12 .

17

10

· 10

· 17

\ r--_

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AB

OlSll m

1000 2000

,..- OllTSI D£

CD

1000 uoo

"- tN~IO(

2000

TlMElhl

FROM A· B IS T£MPERATUR£ CHANGES:

lllXl'F - 3""' - 1l!XJOF FROM C-0 IS TIMPERATURE CHANGES:

uoo" - 1o<> - 1l!Xl'F

6

3000 lhl

-

MP1

69

·69

Ty~1~~1 Ctrooc relrtx~ tioq in piping.

Fig. 13. Reduced three dimensional model 28- inch hot leg isolation valve.

Fig. 14.

11lUT ....

Uf-10£10 fM.(0 THlAMA.l OUOlMATION

Ther mal deformation from inelastic analysb uf FFTF <rnrlium va l.ve body .

L. K. Severud

usually are easy to follow and understand and they usually provide conservative results and guidance. Accordingly, simplified and approxi­mate methods were used on FFTF for three types of application. First, in preliminary design of components, they were used for; (a) screening

·and sensitivity studies, (b) estimating design margins, (c) assessing risks for fabrication releases and, (d) for guidance to design changes. Secondly, in final design stress analyses, they were used for screening to identify most critical areas for detailed inelastic analysis and in some cases were used for final Code assessments pro­viding justifications for conservatism of method arid/or application could be developed. Finally, simplified methods were used for high temperature design education, as they of ten gave insight as to how, and to what extent, various parameters control inelastic analysis results. Some o~ the simplified methods used on the FFTF are as follows:

Temperature-time limits 27.1 Many components in an LMFBR, such as fill and drain piping, operate only short times in the creep range. Recognizing that creep is stress, temperature, and time related; rules for simplified creep-fatigue and ratchet evaluation given in fig\lrl! 15 and 16 m1ri;i dP.velooed. Sub­sequently, they were made part of ASME Code Case N-47, paragraph T-1325 Test No. 4.

ALTERNATE RULES FOR CREEP-FATIGUE AND STRAIN LIMITS WHEN TIME OF EXCURSIONS INTO CREEP RANGE IS SHORT

1F I(r:) k ~ 0.1 FoR s • i.5 sy

AND 2_(E:) !> 0.002 FOR s • 1.25 Sy

AND (pl+ PB+ Q)R~ 3Sm AND 3Sm

THEN, THE LOW TEMP. RULES OF SEC. 111 CAN BE USED WITH

\..(-N") _.$. 0.9 AND ONLY ELASTIC ANALYSIS L'.:\' d 1 IS REQUIRED

Fig. 15. Temperature - time limits for simplified high temperature analysis.

One-dimensional inelastic analyses using two­bar and thick cylinder models 27,2 Bree (ref. 6) and many others (ref. 28-31) have developed simple raLchet und creep ratchet: solutions for simple structure. The ASME Code CasP N-47 has incorporated the O'Donnell­Porowski (ref. 7) method based on lhe cylinder model.

27.3 The Code case points out in T-1324 that the 0-P method is tor axlsy1111·11ctr;l.c etr11rt11res object to axisymmetric loading away from local structural discontinuities.

27. 4 Two-bar and ..:ylinder model solutions are \tseful in studying the effects on ratcheting of (a) str.ess states, levels and their distribu­tion, (b) material,properties, yield strength, creep •nd T.P.laxation and strain hardening, (c) type and geometry of stru'cture and, (d) loatl order and time sequences.

7

27.5 Two very useful one-dimensional inelastic computer programs were developed by Chern (ref.32 and 33). These programs allow detailed materials models to be used with a thick cylinder struc-tural· model. ·

Primary plus secondary stress P + Q limits for elastic analysis 27. 6 The ASME Code for design of nuclear components operating at temperatures below the creep range has a limit on P + Q of ·3 Sm (or 2 Sy). This is based on a simplified shake-down concept. A similar simplified shakedown concept for operation in the creep range is shown in Figure. 17. Using this creep shake-down concept and other considerations, the preliminary design P + Q limits for pipelines was developed by Severud (ref. 8), see Fig. 18. Subsequent detail­ed inelastic analyses of the pipeline designed using the simple limits showed the limits served very well, (ref. 24).

Other approximate methods 27.7 To estimate creep buckling, the formulas for elastic buckling of structures were s·ome­times used with Young's modulus replaced by the tangent modulus of the isochronous stress-strain curve for the sustained stress and time values. 27.8 For creep-fatigue evaluations using the eiastlcally caleu.l:>.t:ed stre!O!? and strain ranges, simplified methods have also been used by Campbell and Severud (ref's. 27 and 4). 27.9 For estimates of creep damage during hold times, uni3xial and biaxial relaxation curves of the creep cumulative damage have been found very useful by Severud (ref. 8).

FOR 5 • 1.5 Sy

1050 FOR 5 • 1.25 Sy

...

.: w ... «

" ....

~ ~

sso

• •• ••' ••' ••' , .. TIME ABOVE soo•p, t. (HOURS)

Fig. 16. Time temperature limits 316 ss.

Fig. 17. Elevated temperature shakedown.

•

et: 0 ti <( u.. z ~ t3 => Q UJ et: UJ

!:: ...J

UJ => <:> .... <( u..

TEMPERATURE (0c) 4Zl 482 533 I'll 649 704

~...--.--....~...,....~r'---.-~..-....,.~....-~~..;..,;..~~......;,

2ll

ELASTIC I ANALYSIS PRELIMINARY DESIGN LIMIT, ~ 13Sm1

ID --1--LOW !!,f:l~·I ~~~!!_TEMP. CRITrRIA CRITrRIA

• ; LINl:S Wlllt DE7All£0 INl:lASTIC ANALYSES ANO PASSED CODE LIMITS

300

100

o.___._~..._--'~-'---''---'-~...__,_~_._--'~-'----'o 100 soo om 1000 1100 1200 l.300

TEMPERATURE l°F1

Fig. 18. Preliminary design primary plus secondary stress limits for 316 stainless steel pipelines.

l!Y

lrl-

101

0 10

10-l 10°120.3 ksil 101! 17. 9)

TENSILE HOLD TIME (h l

Fig. 19. Variation of calculated· fatigue­life reduction factors with tensile hold time for 0.5% strain-range tests of Type 304 stainless steel at 1100°F (593°C).

RECENT CODE DEVEl.OPMENTS AND FUTURE TRENDS

28. ASME code conunittees are evaluating the various other approaches for creep-f atiguQ evaluation, but they desire more test data. A consensus does not yet exist that any of the pre­diction methods has shown itself as being

8

superior. The Code method uses the linear damage method, (Fig. 19). Accordingly, the.Code conunittee members are reluctant to change pre­diction methods at present. Long-time tests underway may show in a couple of years that a different method is better. 29. • Elastic creep-fatigue rules were modified to improve elastic analysis prediction of strain ranges (Severud, ref· 4) • Further improvements are desired in the d5sign fatigue curves by ex­tending them from 10 cycles to lOllcycles. 30. Other areas being addressed by the Code committees include; (a) simplifications to the buckling rules, (b) clarification of limits on third-stage creep, (c) improvements to the strain limits and rules, (d) considerations of actual versus assumed material behavior as effected by weldments, environmental effects and prior cold work, (e) guidance on intended margins of safety, and (f) inclusion of additional materials, such as 2 1/4 Cr-1.0 Mo, Inconel 718, ·and cast 304 stainless steel, into Case N-47. 31. The present Code rules sum creep and fatigue damage fractions as a measure of the total damage. Other approaches have been proposed in the litera­ture; for example, Manson et al. (ref. 34) have proposed a "strain range partitioning" method. Coffin (ref. 35) has offered the "frequency modified" method and others (ref's. 36, 37, 38) have described methods such as the "characteristic slopes," "hysteresis energy," and "damage rate." Some have tried applying "fracture mechanics" type crack-growth methods to high-temperature cyclic life predictions (ref's: 39-42). 32. Design analysts recognize that it is not uncommon for different methods to yield better predictions in different types of applications and that the best method for one application may not be the best for another. For example, cyclic life predictions of a structure that has a known fabrication flaw, crack, or inclusion may be more accurately assessed using fracture mechanics crack-growth techniques than using the present stress-strain range Code methods. 33. The Code Case N-47 rules include strain limit rules that arc held by some design analysts to be necessary to validate the creep-fatigue procedures. Others feel that the limits on cumu­lative positive inelastic strain help guard against creep-rupture failure and thus should be a direct part of the creep-fatigue evaluation procodura~. Whatever the method, or methods, deemed best and recommended for design use, it is important to recognize the need of very clearly delineating the mechanics and procedures of the method.

CONCLUDING REMARKS

34. The most difficult aspects of applying the American design codes for elevated tempera­ture environment relate to satisfying Appendix T of ASME Code Case N-47, rules for strain, defor­mation, and fatigue liini.t.s. Appr<;>aches to satisfying Appendix T limits range from very simple screening methods to detailed inelastic analyses. The U. S. industry has developed oon~iQQ~~ble skill in performin~ these analyses. 35. 'The rules and timits of the Code assuma that the design analyst can analytically predict the structural component stresses and strains in response to an imposed load history. In the interest of cost, the design analyst wants to do the minimlllll and simplest analysis that is necessary to demonstrate design adequacy.

•

L. K. ~everud

However, since prediction accuracy is often sacrificed with the application of simplifying assumptions, the less expensive elastic methods of ten. have to be supplemented with more expensive and complex inelastic analyses. Nonetheless, there is a strong desire to be able to design using only the elastic methods. 36. Most applications of inelastic analysis for FFTF and LMFBR components have consisted of actually computing inelastic response for only up to two or three types of events. Then, the total design events are assessed by extrapolation of the computed damage factors for two or three cycles of each event to the hundreds of cycles of all the design events. Although this may appear like a grossly oversimplified procedure, this approach often is very satisfactory, since the damage per cycle usually decreases with increas­ing number of cycles and a conservative extrapo­lation can then be made. A great deal of useful information has been obtained from inelastic analyses performed to date. It is expected that more efficient computer codes will be developed in the future so that the time and cost for in­elastic analyses should be significantly reduced. 37. Moreover, there appears to be a definite need for creep-fatigue and cumulative strain criteria and procedures for use with both elastic and inelastic analysis. Hopefully, the elastic analysis methods, which inherently must be more conservative than the inelastic rules because of the less informative nature of elastic analysis, are or can be made conservative enough to be safe but not so conservative that they are overly restrictive, unrealistic, or impractical.

REFERENCES

1. Class 1 Components in Elevated Temperature Service, Class III, ASME Boilers and Pressure Vessel Code, Case Interpretations, Code Case N-47 (1592), American Society of Mechanical Engineers, New York. 2. CAMPBELL, R. D., "Creep-Fatigue Calcula-tion Procedures for Code Case 1592 "presented at the Second National Congress on Pressure Vessels and Piping, San Francisco, CA., June 23-27, 1975; published in Advances in Design for Elevated Temperature Environment, ASME, New York. :.L Criteria .tor Uesign o.t Elevated Temperature Class 1 Components in Section III of ASME Boiler and Pressure Vessel Code, published by the American Society of Mechanical Engineers, New York. 4. SEVERUD, L. K., "Background to the Elastic Creep-Fatigue Rules of the ASME Code Case 1592," Nuclear Engineering and Design 45 (1978) 449-455. 5. BERMAN, I., GANGADHARAN, A. C. and ' GUPTA, G. D., "Buckling and Instability at Elevated Temperature," presented at the Second National Congress on Pressure Vessels and Piping, San francisco, CA., June 23-27, 1975; published in Advances in Design for Elevated Temperature Environment, ASME, New York. 6. BREE, J., "Incremental Growth Due to Creep and Plastic Yielding of Thin Tubes Subjected to Internal Pressure and Cyclic Thermal Stresses," J. Strain Analysis, 1968, 3, No. 2, 122-127. 7. O'DONNELL, W. J., and POROWSKI, J., "UppGir Bounds for Accumu~:ited Straina Due to Creep Ratcheting," Welding Research Council · Bulletin No, 195, June 1974, also Trans ASME, Journal of Pressure Vessel Technology, Vol. 96, p. 126, 1974.

9

8. SEVERUD, L. K., "Simplified Methods and Applications to th~ Preliminary Design of Piping for Elevated Temperature Service," presented at the Second National Congress on Pressure Vessels and Piping, San Francisco, CA., June 23-27, 1975; published in Advanced in Design for Elevated Temperature Environment, ASME, New York. 9. PUGH, C. E., "Constitutive Equations for Creep Analysis of LMFBR Components," presented at the Second National Congress on Pressure Vessels and Piping, San Francisco, CA., June 23-27, 1975, published in Advances in Design for Elevated Temperature Environment, ASME, New York. 10. BIGELOW, C. C. , "Experience in the Imple­mentation of Current High Temperature Struc­tural Design Technology, presented at the Third International Conference on Structural Mechanics in Reactor Technology, London, U.K., Sept. 1-5, 1975. 11. GANGADHARAN, A. C., PAI, D. H., "Non-Linear Creep Fatigue Analysis of a Sodium Heat Exchanger Component for the Fast Flux Test Facility," paper C215/73, presented at the International Confer~ ence on Creep and Fatigue in Elevated Temperature Applications, Philadelphia, PA., Sept. 1973. 12. DHALLA, A. K., "Effect of Yield Strength Variation on the Inelastic Response of a C-Ring" ASME paper 75-PVP-31, presented at the Second National Congress on Pressure Vessels and Piping Technology, San Francisco, CA., June 23-27, 1975. 13. DHALLA, A. K. and ROCHE, R. V., "Inelastic Analysis and Satisfaction of Design Criteria of a High Temperature Comppnent," presented at the Second. National Congress on Pressure Vessels and Piping, San Francisco, CA., June 23-27, 1975; published in Advanced in Design for Elevated Temperature Environment, ASME, New York. 14. WU, c. G., "Inelastic Analysis in LMFBR Reactor Vessel Design," paper Gl-5, presented at the Second International Conference on Structural Mechanics in Reactor Technology, Berlin, Germany, Sept. 10-15, 1973. 15. HIBBITT, H. D., SORENSON, E. P., and MARCAL, P. v., "The Elastic-Plast.ic and Creep Analysis of Pipelines by Finite Elements," Pro­ceedings of the Second International Conference on Pressure Vessel Technology (Part 1. Design and Analysis), October lY/], pp. l]Y-~l, ASME, New York. 16. PAN, Y. S. and JETTER, R. I, "Inelastic Analysis of Pipelines in FFTF CLS Module," Pressure Vessel and Piping Conference, Miami Beach, FL., June 24-28, 1974, published in Pressure Vessels and Piping; Analysis and Computers, ASME, New York, N. Y. . • 17. MINAMI, H. M., "Application of Section III Class 1 Design Rules to an Elevated Temperature Component," presented at the Second National Congress on Pressure Vessels and Piping, San Francisco, CA., June 23-27, 1975p published in Advances in Design for Elevated Temperature Environment, ASME, New York, N. Y. 18. WEINER, E. O., "A Three-Dimensional Inelastic Finite Element Analysis of a Solid Y-Type Cylinder Interaction," presented at the Fourth International Conference on Structural Mechanics in Reactor Technology, San Francisco, CA., Aug. 15-19, 1977. 11 19. SAMPSON, R. C. and JAGELS, R. E., Stress Analysis for the Design of Liquid Metal Piping in the Fast Flux Test Facility," presented at the Joint ASME/CSME Pressure Vessels and Piping Conf. Montreal, Canada, June 25-30, 1978, ASME papers 78-PVP-21.

- .. •

L. K. Severud . .

20. CHEN, W. L. and WEINER, E. O., "Inelastic Analysis of Pipeline in FFTF Heat Transport System," presented at the Joint ASME/CSME Pres­sure Vessels and Piping Conference, Montreal, Canada, June 25-30, 1978, ASME special publica­tion PVP-PB-028. 21. HUANG, S. N., "Inelastic Analysis of Two Pipelines in the Fast Flux Test Facility," presented at the Third U. S. National Congress on Pressure Vessels and Piping, San Francisco, CA., June 25-29, 1979, published in ASME special publication PVP-36. 22. SAMPSON, R. C. , "Stress Analysis of Conical Flued Heads for FFTF Liquid Metal Piping Anchors," presented at the Joint ASME/CSME Pressure Vessels and Piping Conference, Montreal, CAnada, June 25-30, 1978, ASME special publication PVP-f>B-028. 23. WINKEL, B. V., "Experience with Simplified Inelastic Analysis of FFTF Test Assemblies," pre­sented at the Joint ASME/CSME Pressure Vessels and Piping Conference, Montreal, Canada, June 25-30, 1978, ASME special publication PVP-PB-028. 24. SEVERUD, L. K., "Experience with Simplified Inelastic Analysis of Piping," presented at the Nucelear Engineering Division ASME Century 2 Conference, San Francisco, CA., August 18, 1980. 25. ANDERSON, M. J., HYDE, L. L., WAGNER, S. E., and SEVERUD, L. K., "Insulated Pipe Clamp Design," presented at the ASME P&PV Conf., San Francisco, CA., Aug. 1980. . 26. LINDQUIST, M. R., and ANDERSON, M. J., "Pipe Clamp Effects on Thin-Walled Pipe Desigt)," presented at the ASME B&PV Conference, San Francisco, CA., Aug., 1980. 27. CAMPBELL, IL D., "Creep-Fatigue Calculation Procedures for Code Case 1592," presented at the Second National Congress on Pressure Vessels and Piping, San Francisco, CA., June 23-27, 1975; published in Advances in Design for Elevated Temperature Environment, ASME, New York. 28. MILLER, D. R., 'Thermal-Stress Ratchet Mechanism in Pressure Vessels;• Trans. ASME, Journal of Basic Engineering, Vol. 8, June 1959. 29. BURGREEN, D., "Structural Growth Produced by Thermal Cycling," ASME Journal of Basic Engineering, Dec., 1968. 30. EDMONDS, H. G., and BEER, F. J., "Notes on Incremental Collapse in Pressure Vesc:cls," Journal of Mechanical Engineering Science, Vol. 3., No. 3, 1961. . 31. MULCAHY, T·. M., "Thermal Ratcheting of a Beam Element Having an Idealized Bauschinger ' Effect," ASME paper 75-WA-Mat-4, presented at the 1975 ASME WAM, Houston, TX, Nov. 30-Dec. 4, 1975. 32. CHERN, J.M., and PAI, D. H., "A Simplified Tool for the Elevated Temperature Cyclic Analysis of Pressure Components," Second International Conference on Pr·essure Vessel Technology, Part I: Design and Analysis, San Antonio, 1973, pp . . 263-275. 33. CHERN, J. M. and PAI, D. H., "Inelastic Analysis of a Straight Tube under Combined Bend­ing, Pressure and Thermal Loads; ASME paper 75-PVP-19, presented at the Second National Congress on Pressure Vessels and Piping, San Francisco, CA., June 23-27, 1975. 34. MANSON, S. S., "The Challenge to Unify Treatment of High Temperature Fatigue - A Partisan Proposal Based on Strainrailge Pa<titfon­ing," Fatigue at Elevated Temperatures, STP-520, pp. 744-82, ASTM, 1973.

10

35. COFFIN, L. F., "Fatigue at High Tempera­ture - Prediction and Interpretation," James Clayton Lecture, presented at 1973-1974 International Conference on Creep and Fatigue, Sheffield, England; published in Proc. Inst. Mechanical Engineering 188 9/74, 109-27 (1974). 36. CONWAY, J. B., STENTZ, R. H. and BERLING, J. T., "Fatigue Tensile, and Relaxation Behavior of Stainless Steels," TID-26135, pp. 92-101. 37. FONG, J. T., "Energy Approach for Creep and Fatigue Interactions of Metals at High Tempera­ture," ASME paper 75-PVP-30, presented at the Second National Congress on Pressure Vessels and Piping Technology, San Francisco, CA., June 23-27, 1975. 38. MAJUMDAR, S. , and MAIYA, P. S. , "A Damage Equation for Creep-Fatigue Interaction," presen­ted at ASME WAM, Dec., 1976, published in 1976 ASME-MPC Symposium on Creep-Fatigue Interaction, MPC-3, ASME, New York. 39. SHAHINIAN, P., SMITH, H. H. and WATSON, H. E., "Fatigue Crack Growth in Type 316 Stainless Steel at High Temperature," ASME Paper 71, PVP 25, 1971. 40. JAMES, L. A., and SCHWENK, E. B., "Fatigue­Crack Propagation on Type 304 Stainless Steel at Elevated Temperatures," Met. Trans. 2, 491 (1971). 41. CARDEN, A. E., "Parametric Analysis of Fatigue Crack Growth," Int. Con£. on Creep and Fatigue at Elevated Temperature· Applications, Philadelphia, PA, Sept. ,1973. 42. SPEIDEL, M. O., "Fatigue Crack Growth at High Temperature," in High Temperature Materials in Gas Turbines, e.d. by R. Ji'. Sahim and M .. 0. Speidel, Elsevier Scientific Publishing Company, 1974.