xiii symposium of magnetic measurements & … xiii_ smmm.pdf · stalprodukt s.a., bochnia – 3...

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Polish Society of Theoretical and Applied Electrical Engineering

Częstochowa Branch

XIII SYMPOSIUM

OF MAGNETIC MEASUREMENTS & MODELLING

Cracow – Wieliczka, 8th - 10th October 2018

ABSTRACTS

Under the auspices of

Polish Academy of Science Committee of Electrical Engineering

Rector of Czestochowa University of Technology prof. dr hab. inż. Norbert Sczygiol

Organized in the jubilee year of 70th Anniversary of Czestochowa University of Technology

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XIII SYMPOSIUM OF MAGNETIC MEASUREMENTS & MODELLING


ORGANIZED BY:

Polish Society of Theoretical and Applied Electrical

Engineering, Częstochowa Branch

Faculty of Electrical Engineering

Częstochowa University of Technology

Tele & Radio Research Institute, Warsaw

Faculty of Electrical and Computer Engineering

Cracow University of Technology

Ariel University, Israel

Institute of Materials Science

University of Silesia in Katowice

Stalprodukt S.A., Bochnia

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SCIENTIFIC COMITTEE

Honorary Chairman:

Jacek R. Przygodzki Retired Professor of Warsaw University of Technology

Warsaw, Poland

Philip Anderson Cardiff University, United Kingdom

Marcos F. de Campos Fluminense Federal University, Brazil

Ermanno Cardelli University of Perugia, Italy

Laurent Daniel Laboratoire de Génie Electrique de Paris, France

Andrzej Demenko Poznań University of Technology, Poland

Victorino Franco University of Sevilla, Spain

Octavio Guzman Universidad Nacional de Colombia, Colombia

Kay Hameyer RWTH Aachen University, Germany

Grzegorz Haneczok University of Silesia, Poland

Robert G. Harrison Carleton University, Canada

Adam Jagiełło Cracow University of Technology, Poland

Andrzej Kapłon Kielce University of Technology, Poland

Afef Kedous-Lebouc G2ELab Grenoble, France

Ivan Kityk Częstochowa University of Technology, Poland

Krzysztof Kluszczyński Cracow University of Technology, Poland

Miklós Kuczman Széchenyi István University of Győr, Hungary

Aminta Mendoza Universidad Nacional de Colombia, Colombia

Yevgen Melikhov Wolfson Centre for Magnetics, Cardiff University, UK

Kruno Miličević University of Osijek, Croatia

Andrzej Nowakowski Tele & Radio Research Institute, Poland

Katarzyna Oźga Częstochowa University of Technology, Poland

Yosef Pinhasi Ariel University, Israel

Helmut Pfützner Wien University, Austria

Marie-Ange Raulet Laboratiore Ampère, France

Pavel Ripka Czech Technical University in Prague, Czech Republic

Juliette Soulard University of Warwick, United Kingdom

Roman Szewczyk Warsaw University of Technology, Poland

Barbara Ślusarek Tele & Radio Research Institute, Poland

Manuel Vázquez Institute of Material Science of Madrid, Spain

Jerzy Wysłocki Częstochowa University of Technology, Poland

Asher Yahalom Ariel University, Israel

Sergey E. Zirka Oles Honchar Dnipro National University, Ukraine

Stan Żurek Megger Ltd., United Kingdom

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ORGANIZING COMITTEE

Jan Szczygłowski Chairman

Krzysztof Chwastek Co-Chairman

Mariusz Najgebauer Secretary

Adam Jakubas

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FRAME PROGRAM OF



Monday 08.10.2018

9.00 – 11.45 Registration of the participants

11.35 – 12:00 Official opening of the Symposium

12.00 – 13.30 Plenary Session

13.30 – 14.30 Lunch

14.30 – 16.30 Session 1: Magnetic Materials

16.30 – 16.45 Coffee break

16.45 – 18.30 Session 2: Electric Machines and Devices

19.00 – 24.00 Barbeque

Tuesday 09.10.2018

7.30 – 9.00 Breakfast

9.00 – 10.45 Session 3: Properties of Soft Magnetic Materials


11.00 – 13.00 Session 4: Hysteresis Modelling and Related Issues

13.00 – 14.00 Lunch

14.00 – 15.30 Session 5: Sensors and Actuators


16.00 – 17.30 Session 6: Electromagnetic Field Analysis

18.30 Setting out to the “Wieliczka” Salt Mine

19.00 – 22.00 Gala dinner in “Wieliczka” Salt Mine

Wednesday 10.10.2018


9.00 Setting out for excursion to the “Wieliczka” Salt Mine

9.25 – 11.00 Sightseeing of the “Wieliczka” Salt Mine

11.30 – 11.45 Closing remarks of the Symposium

11:45 – 12.00 Checking out

12.00 – 13.00 Lunch

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PROGRAM OF



Monday 08.10.2018

9.00 – 12.00 Registration of the participants

11.45 – 12:00 Official opening of the Symposium

12.00 – 13.30 Plenary Session I

Chairmen: Jan Sykulski, Katarzyna Oźga

1. P. Svec, I. Janotova, D. Janickovic, B. Kunca, J. Marcin, I. Matko,

I. Skorvanek, P. Svec Sr.: New developments in rapidly quenched soft

and hard magnetic alloys

2. M.F. de Campos: Methods for texture improvement in electrical steels

3. I. Mészáros, B. Bögre: Magnetic measurement of ferrite content of alloys

13.30 – 14.30 Lunch

14.30 – 16.30 Session 1: Magnetic Materials

Chairmen: Peter Svec, Adam Jakubas

1. M. Przybylski, B. Ślusarek, T. Bednarczyk, G. Chmiel: Magnetic and

mechanical properties of rubber bonded magnets with different type and

amount of hard magnetic powder

2. T. Garstka: A new parameter for the Barkhausen noise characterization

3. R. Gozdur, P. Gębara, K. Chwastek: Influence of DC-bias magnetic field on

dynamic magnetic properties of LaFeCoSi alloy

4. B. Guzowski, R. Gozdur, A. Kociubiński: Magnetic substrates made of sputtered

Y3Fe5O12

5. K. Kotynia, A. Chrobak, P. Pawlik: Structure and magnetic properties of the

rapidly solidified Gd3Zr10Fe55Co10Mo5W2B15 alloy

6. P. Kwapuliński, G. Haneczok: Magnetic relaxation in iron based melt spun

ribbons


16.45 – 18.30 Session 2: Electrical Machines and Devices

Chairmen: Kay Hameyer, Marek Przybylski

1. B. Koprivica, K. Chwastek, M. Koprivica: Short-circuit and load operation of

single-phase transformer at low frequencies

2. D. Kapelski, E. Kucal, A. Szymański: The magnetization curve of FeNiCo alloy

and the influence of its application on the penning effect in vacuum interrupter

3. D. Danielczyk, D. Janiszewski, C. Jedryczka, D. Kapelski, M. Krystkowiak:

Analysis of dual star permanent magnet synchronous motor with rotor back

iron made of soft magnetic composite

4. A. Kapłon, J. Rolek: Modeling of the magnetic field distribution in air gap of

the synchronous machine from permanent magnets in the rotor

5. W.A. Pluta: Surface isolation of modern electrical types for magnetic cores

19.00 – 24.00 Barbeque

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PROGRAM OF



Tuesday 09.10.2018


9.00 – 10.45 Session 3: Properties of Soft Magnetic Materials

Chairmen: István Mészáros, Tomasz Garstka

1. M.F. de Campos: Interpretation of loss separation with the Haller-Kramer

model

2. B. Koprivica, K. Chwastek: Verification of Bertotti’s loss model for non-

standard excitation

3. M. Bereźnicki, P. Jabłoński, M. Najgebauer, J. Szczygłowski: Analysis of the

skin effect in the calculation of power loss components in soft magnetic

materials

4. W.A. Pluta: Anisotropy of specific total loss components in Goss textured

electrical steel

5. N. Leuning, S. Steentjes, K. Hameyer: Evaluation of the interdependency of

mechanical cutting and magnetic anisotropy on the magnetic properties of

non-oriented FeSi electrical steel


11.00 – 13.00 Session 4: Hysteresis Modelling and Related Issues

Chairmen: Asher Yahalom, Ewa Łada-Tondyra

1. J. Eichler, M. Novak, M. Kosek: Experimental determination of Preisach model

for grain oriented steel

2. R. Jastrzębski, A. Jakubas, K. Chwastek: A comparison of two

phenomenological descriptions of magnetization curves based on T(x) model

3. W. Mazgaj, Z. Szular, M. Sierzega: Inverse model of the magnetic hysteresis

based on an exponential function

4. M. Novak: Difficulties cause by magnetic after-effect during identification of

the Preisach hysteresis model weighting function

5. M.F. de Campos, J.A. de Castro: Predicting recoil curves in Stoner-Wohlfarth

anisotropic magnets

6. L.F.T. Costa, G.J.L. Gerhardt, F.P. Missell, M.F. de Campos: Interpretation of

magnetic Barkhausen noise bursts in low frequency measurements

13.00 – 14.00 Lunch

14.00 – 15.30 Session 5: Magnetic Sensors and Actuators

Chairmen: Andrzej Nowakowski, Branko Koprivica

1. R. Szewczyk, A. Bieńkowski, M. Nowicki: Jiles-Atherton-Sablik model of

magneto-mechanical characteristics of soft magnetic materials - A review

2. D. Stachowiak, M. Kurzawa: A computational and experimental study of shape

memory alloy spring actuator

3. A. Lisowiec, A. Nowakowski, G. Kowalski, P. Wlazło: Miniature current sensor

for medium voltage networks

4. M. Woloszyn, S. Michalski, B. Potrac: Optimal flight direction of magnetic

system during object's detection on the Baltic Sea

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16.00 – 17.30 Session 6: Electromagnetic Field Analysis

Chairmen: Marco F. de Campos, Roman Gozdur

1. I. Chaimov, A. Yahalom: Correcting for FEL magnetic field distortions. The

method of bilinear shimming

2. A. Etinger, Y. Golovachev, G.A. Pinhasi, Y. Pinhasi: Propagation of Tera-Hertz

radiation in foggy conditions

3. E. Łada-Tondyra: The impact of applicator size on distribution of

electromagnetic field used in magnetotherapy

4. A. Cywiński, K. Chwastek, P. Gas: The influence of skin and proximity effects

on temperature distribution in multi-bundle cable lines

5. S. Nazrulla, E.G. Strangas, J.S. Agapiou, T.A. Perry: A Device for the Study of

Electrical Steel Losses in Stator Lamination Stacks

18.30 Setting out to the “Wieliczka” Salt Mine

19.00 – 22.00 Gala dinner in “Wieliczka” Salt Mine

The gala dinner of SMMM’2018 is organized in the Jan Haluszka Chamber

at the „Wieliczka” Salt Mine, which is a UNESCO World Heritage Site.

The Chamber is located about 135 meters underground and has an interesting,

vault-like shape and magnificent walls carved in green rock salt.

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PROGRAM OF



Wednesday 10.10.2018


9.00 Setting out for excursion to the “Wieliczka” Salt Mine

9.25 – 11.00 Sightseeing of the “Wieliczka” Salt Mine


11.30 – 11.45 Closing remarks of the Symposium

11.45 – 12.00 Checking out

12.00 – 13.00 Lunch

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ABSTRACTS

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ANALYSIS OF THE SKIN EFFECT IN THE CALCULATION

OF POWER LOSS COMPONENTS IN SOFT MAGNETIC MATERIALS

M. Bereźnicki, P. Jabłoński, M. Najgebauer and J. Szczygłowski

Częstochowa University of Technology, Faculty of Electrical Engineering

al. Armii Krajowej 17, 42-200 Częstochowa, Poland

e-mail: [email protected], [email protected], [email protected], [email protected]

Abstract. In the paper, the theoretical equations describing hysteresis and macroscopic eddy current losses in soft

magnetic materials are analyzed. Usually, the low frequency approximations are used to evaluate the components of

losses. The approximations assume that the magnetic flux is uniform throughout the whole sample. For higher

frequencies this may be not justified, especially for thicker electric sheets. Therefore, simplified formulas describing

hysteresis and eddy currents losses for two types of electrical steels 3% SiFe and 6.5% SiFe are analyzed. For both

components of losses, the limiting frequencies at which it is necessary to take the skin effect into account are

determined.

I. INTRODUCTION

Magnetic materials, including electric steels, are widely used in many electric devices. One of

the most important parameters of such materials is energy loss per magnetization cycle and mass

unit, called briefly specific loss. The specific loss is determined via measurements, but there are

also various theoretical and empirical formulas to express this loss as a function of frequency,

magnetic flux density, material parameters and sample dimensions [1, 2]. The fundamental

difficulty to obtain the theoretical formulas lies in nonlinearity and hysteresis of magnetization

process [3, 4]. These phenomena are often neglected when calculating magnetic fields, which leads

to simplified formulas describing power loss and hysteresis loops. As the frequency of magnetic

flux increases, the hysteresis loop becomes wider, as a result of increased energy dissipation due to

eddy current flows. In addition, the induced eddy currents generate their own magnetic field, which

changes the magnetic field distribution inside the sample. This effect is the stronger the smaller is

the skin depth. These formulas are even more simplified when analyzing two limiting cases

corresponding to the so-called weak and strong skin effect. In the first case, it is assumed that the

skin depth is much larger than the sample thickness. This allows one to neglect the skin effect and

to assume a uniform flux throughout the sample. In the second case, the skin depth is assumed to

be much smaller than the sample thickness, which leads to highly non-uniform flux distribution.

The aim of this paper is to analyze the application range of the theoretical formulas for power loss,

which take the weak and strong skin effects into account.

II. THEORETICAL CONSIDERATIONS

The theoretical formula for power loss in magnetic materials results from considering the

Maxwell equations. The Poynting theorem reveals two phenomena responsible for loss in magnetic

materials: the hysteresis and eddy current components [5]. The theoretical considerations assume

also that parameters such as conductivity and magnetic permeability are constant throughout the

sample. In the case of electrical steel, it is also assumed that the sample thickness is much lower

than its other dimensions.

Under the abovementioned assumptions, the loss due to hysteresis phenomena in a thin steel

lamination of thickness d and for a sinusoidal magnetic flux of density Bm throughout the sample

has the following form

coscosh

sinsinh2

m

hyst

sfBP , (1)

mailto:[email protected]



– 14 –

where: f is frequency of magnetizing field, is the skin effect ratio and s is a coefficient depending on the shape of hysteresis loop. Formula (1) is seldom used directly, but rather two cases of weak and strong skin effect are considered [5]:

2

mlow

hyst

2sfBP and

2/1

2

m

2/12/32/12

mhigh

hyst

dBsfsfBP . (2a / 2b)

In a similar way, the general formula for power loss due to eddy currents induced in the sample can

be derived [5, 6]:

coscosh

sinsinh

2

2

m

eddy

fBP , (3)

This loss component is called the macroscopic eddy current loss, because only macroscopic eddy

current flows are taken into account. As in the case of hysteresis loss, the formulas corresponding

to the weak and strong skin effect are as follows

66

2

m

222

2

2

mlow

eddy

BfdfBP

and

2/1

2

m

2/12/32/32

mhigh

eddy22

dBffBP . (4a / 4b)

Theoretical characteristic for the hysteresis loss (1, 2a, 2b) and eddy current loss (3, 4a, 4b) are

depicted in figure 1 and 2, respectively.

Fig.1. Hysteresis loss: Eq. (1) – solid line,

Eq. (2a) – dotted line, and Eq. (2b) – dashed line

Fig.2. Eddy current loss: Eq. (3) – solid line,

Eq. (3a) – dotted line, and Eq. (3b) – dashed line

The analysis of given formulas as well as determination of their application ranges will be

determined for 3% SiFe and 6.5% SiFe electrical steels.

REFERENCES

[1] Krings A., Soulard J., Overview and comparison of iron loss models for electrical machines, in

Proceedings of International Conference on Ecological Vehicles and Renewable Energies EVER 2010,

25-28 March 2010, Monaco, abridged version published in Journal of Electrical Engineering, vol. 10,

no. 3, 2010, paper 10.3.22

[2] Bertotti G., General properties of power losses in soft ferromagnetic materials, IEEE Trans. Magn., vol.

24, 1988, pp. 621-630

[3] Bertotti G., Some considerations on the physical interpretation of eddy current losses in ferromagnetic

materials, J. Magn. Mater., vol. 54-57, 1986, pp. 1556-1560

[4] Dąbrowski M., Analiza obwodów magnetycznych. Straty mocy w obwodach, Polska Akademia Nauk,

Oddział w Poznaniu, seria Elektrotechnika, tom III, PWN, Warszawa-Poznań, 1981

[5] Barranger J., Hysteresis and eddy-current losses of transformer lamination viewed as an application of

the Pointing Theorem, NASA Technical Note, D-3114, 1965

[6] Bertotti G., Hysteresis in magnetism, Academic Press, San Diego, 1998

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CORRECTING FOR FEL MAGNETIC FIELD DISTORTIONS

THE METHOD OF BILINEAR SHIMMING

I. Chaimov and A. Yahalom

Department of Electrical & Electronic Engineering

Ariel University, Ariel 40700, Israel, e-mail: [email protected]

One of the main requirements of a Free Electron Laser (FEL) instrument is to achieve nominal

B-field values with a high accuracy, along the main axis of the FEL’s permanent magnetic periodic

undulator, known as Wiggler device of the Halbach configuration. From practical reasons, mainly

due to magnets manufacturing, there are deviations of the magnetic field of the magnets bars which

construct the Wiggler device with reference to the theoretical magnetic field - in most cases, this

deviation reduces the efficiency of the radiation extraction - reducing the delivered power and

energy extracted from the device and hence, giving rise to undesirable heat absorbed at the device,

which in some cases can cause a de-magnetization affect at the magnet bars. The magnetic bars

magnetization could be treated as a random variable for which we assume to have a normal random

distribution having a standard deviation from 5% to 10% (depends on the quality of magnet’s

manufacturing). Our main purpose is to optimize the magnetic field by adding small dipole-like

magnets which correct for the noise in practical magnetic fields. This enable us to minimize the

unwanted field noise to minimum and achieve a more optimal Wiggler device.

The field generated by a single dipole of strength m located at position k at position n is:

(1)

where . and:

(2)

The actual noisy distribution of the magnetic field emerging from the Wiggler structure which is

denoted . The required magnetic field is denoted as . The difference between the actual and

required field at point n is given by:

(3)

The target function Tf with mk as the vector variable is thus defined as:

(4)

where the summation is performed for all the points n were the field is measured. This can be

written as a bilinear form:

(5)


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where:

(6)

The minimum of this target function is obtained as follows:

(7)

Hence we obtain the optimal dipole values:

(8)

And the minimum value of the target function:

(9)

In practice not every dipole value is achievable, hence one cannot obtain Tfmin. However, using

realizable discretization techniques one can obtain considerable improvement in the quality of the

magnetic field.

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INTERPRETATION OF MAGNETIC BARKHAUSEN NOISE BURSTS IN

LOW FREQUENCY MEASUREMENTS

L.F.T. Costa1, G.J.L. Gerhardt

2, F.P. Missell

2 and

M.F. de Campos

3

1 CEPEL, PFDF, Divisão de Planejamento e Fomento, Rio de Janeiro, RJ, Brazil

2 Universidade de Caxias do Sul, Caxias do Sul 95070-560, Brazil

3 Universidade Federal Fluminense, Volta Redonda 27255-125, Brazil, e-mail: [email protected]

Abstract. Experimental results allow the identification of three main Magnetic Barkhausen Noise (MBN) bursts, each

occurring at a different applied field. Magnetostrictive effects can be related to the 1st and 3

rd bursts, because closure

domain walls are created and/or eliminated. This gives an important insight on how stress may affect the losses and

MBN.

I. INTRODUCTION )

The Magnetic Barkhausen Noise can provide information about the main dissipative

mechanisms in the quasi-static hysteresis. In the present study, measurements performed at 0.5 Hz

are discussed. This situation is near that of quasi-static situation. The measurements were

performed in toroids, and this means that the exact field where the bursts take place can be

determined, and compared with the quasi-static hysteresis.

There are several possible dissipative mechanisms in a hysteresis curve:

i) creation/annihilation of domain walls, ii) domain wall displacement, iii) creation/annihilation of

closure domain walls with magnetostrictive effects, see Fig. 1 [1,2] and iv) domain rotation. In our

previous investigation, it was found that domain rotation produces small MBN [3]. MBN provides

insight about all these mechanisms

Fig.1. Due to magnetostictive effects, there is variation of volume in the direction of magnetization.

This gives rise to a misfit along 90o domain boundaries in iron [1,2].

II. RESULTS AND DISCUSSION

Three main bursts can be summarized as follows (see the arrows in Fig.s 2 and 3):(i) CDF

(closure domain formation) at applied field H near 0. (ii) DWM (domain wall movement) for

applied field near the coercive field. (iii) CDA (closure domain annihilation) at higher applied

field. As magnetostrictive effects strongly affect closure domain walls, then the position and height

of the bursts can be altered by applied stresses. In Figs 2. and 3, the green curve is denoted dB/dH

curves. Note the difference between the blue curves (MBN) and dB/dH. The main burst is due to

180o domain wall movement and appears exactly at the coercive field. Although the main burst is

strong, the burst due to domain wall formation is difficult to detect. But if closure domain wall

elimination produces a burst, then it is also expected that creation of closure domain walls also

generates a burst.

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Fig.2. Sample 1045 steel. The arrows indicate the three main bursts. F=0.5 Hz.

Fig.3. Sample 1030 steel. The arrows indicate the three main bursts. F=0.5 Hz.

REFERENCES

[7] Kittel C., Physical Theory of Ferromagnetic Domains, Rev. Mod. Phys., vol. 21, 1949, pp. 541

[8] Hosford W. F., Iron and Steel, CRC Press, 2012

[9] Costa L.F.T., de Campos M.F., Gerhardt G.J.L., Missell FP, Hysteresis and Magnetic Barkhausen

Noise for SAE 1020 and 1045 Steels with Different Microstructures, IEEE Transactions on Magnetics,

vol. 50, 2014, pp. 2001504.

– 19 –

THE INFLUENCE OF SKIN AND PROXIMITY EFFECTS ON

TEMPERATURE DISTRIBUTION IN MULTI-BUNDLE CABLE LINES

A. Cywiński1, K. Chwastek

2 and P. Gas

3

1 Design Bureau Omega-Projekt, ul. Topolowa 1, 43-100 Tychy, Poland

2 Faculty of Electrical Engineering, Częstochowa University of Technology,

Al. Armii Krajowej 17, 42-201 Czestochowa, Poland, e-mail: [email protected] 3 AGH University of Science and Technology, Department of Electrical and Power Engineering

Al. Adama Mickiewicza 30, 30-059 Kraków, Poland, e-mail: [email protected]

Abstract. The paper focuses on Finite-Element-Method based computations of coupled electromagnetic-thermal effects

in multi-bundle cable lines.

I. INTRODUCTION

The fundamental document for Polish designers of low voltage cable lines is the standard

PN-HD 60364-5-52:2011 Low voltage electrical installations. It includes a number of correction

factors for computation of ampacity of multi-bundle cable lines to take into account the conditions

of thermal exchange between the buried cables and their environment. However the standard does

not take into account the possible increase of temperature in multi-bundle cable lines due to skin

and proximity effects. In order to gain insight on the possible coupled effects between

electromagnetic and thermal fields in such systems it is necessary to carry out Finite-Element

computations [1,2].

II. FEM COMPUTATIONS OF THERMAL EFFECTS

IN MULTI-BUNDLE CABLE LINES

Contemporary commercial software like Ansys-Maxwell-Icepack suite allows one to carry out

FEM simulations that take into account coupled electromagnetic-thermal effects, cf. Fig. 1.

Fig.1. Temperature distribution in a single phase cable

Such piece of software may be an indispensable tool for the designers of multi-bundle cable

lines.



– 20 –

In the full version of the paper exemplary results of computations using FEM method as well

as results of their experimental verification on a self-designed lab stand (Fig. 2) shall be provided

for different spatial configurations of cable lines.

Fig.2. Self-designed laboratory stand for examination of coupled electromagnetic-thermal effects

in multi-bundle cable lines.

REFERENCES

[1] Skibko Z., Obciążalność prądowa przewodów ułożonych wielowarstwowo, Rozprawa doktorska,

Politechnika Białostocka, Maj 2008 r.

[2] Szczegielniak T., Analiza elektromagnetycznych i termicznych pól sprzężonych w jednobiegunowych

torach wielkopradowych, Rozprawa habilitacyjna, Politechnika Częstochowska, 2018 (na prawach

rękopisu)

– 21 –

ANALYSIS OF DUAL STAR PERMANENT MAGNET SYNCHRONOUS

MOTOR WITH ROTOR BACK IRON MADE

OF SOFT MAGNETIC COMPOSITE

D. Danielczyk2, D. Janiszewski

2, C. Jedryczka

2, D. Kapelski

1 and M. Krystkowiak

2

1 Tele and Radio Research Institute, Ratuszowa 11 St., 03-540 Warsaw, Poland,

e-mails: [email protected]; 2Institute of Electrical Engineering and Electronics, Piotrowo 3a St., 60-965, Poznan, Poland,

e-mails: [email protected], [email protected],

[email protected], [email protected]

Abstract. The research conducted at Poznan University of Technology with cooperation the Tele- and Radio Research

Institute deals with finite element analysis of six-phase, dual star permanent magnet synchronous motor. To reduce

eddy current losses in the rotor of the machine the rotor back iron segments have been made of soft magnetic

composite (SMC). SMC are composites of iron powder particles separated with an electrically insulated layer. This

technology has many advantages in relation to classical laminated core solutions; among other lower manufacturing

costs due to simpler technology and reduced eddy current losses out of order times lower conductivity. The

mathematical model of machine utilizes field circuit approach assuming planar symmetry of the machine. The

magnetic properties of the applied SMC material have been introduced into the model basing on BH curves and unit

losses vs. frequency characteristics measured at Tele- and Radio Research Institute. Accuracy of developed numerical

model has been verified by measurements of the machine performance tested on the elaborated research stand.

I.INTRODUCTION

In the recent years increased interest in permanent magnet synchronous machines (PMSM)

with fractional slot contracted windings (FSCW) can be observed in many research teams [1,2]. By

shorten end windings (in relation to machines of distributed windings) such machines offers

possibility to reduce the copper losses and thus increase of efficiency with simultaneous winding

cost reduction. Nevertheless PMSM with FSCW suffer of increased eddy current losses in the rotor

due to high level of sub and super harmonics in spatial distribution of the magnetomotive force

excited by such type of winding [3]. One of the methods to reduce level of distortion mmf spatial

distribution is to increase the number of phases of the winding [2]. From the other hand the losses

in the rotor can be reduced by using appropriate material for the construction of rotor magnetic

circuit. In the presented approach a dual three-phase machine with rotor yoke made of SMC

material has been studied by means of finite element method (FEM). The view rotor of studied

machine and structure of the SMC material have been shown in fig. 1a) and 1b), respectively.

Fig.1. Spoke type magnet SMC based rotor of the machine a); illustration of the SMC structure b)

The rotor of the machine is composed of wedges made of the Somaloy 500 - the SMC

concept brand from Swedish company Höganäs - and NdFeB sintered magnets magnetized as

a) b)






– 22 –

shown in fig 2a). Exemplary magnetic flux density plot at no load state has been shown in

fig. 2b.

Fig.2. Magnetization direction a); exemplary magnetic flux density plot b); measured current waveforms c)

The model accuracy has been verified by comparing the simulation results with the

measurements performed on the special designed research test stand that consists of studied

machine driven by six phase inverter controlled by ALS-1369 DSP based control system, data

acquisition system (DAQ) based on National Instrument NI9220 DAQ and DC power system

utilizing dual channel TTIQPX600DP programmable DC supply. The diagram and photo of the

developed test stand have been shown in fig. 3a) and b) respectively.

Fig.3. Developed test stand a) block diagram; b) photograph (1 - PC for programming and data processing;

2 - supply, DAQ and control system; 3 - six phase PMSM)

More details about simulation techniques as well as discussion about conducted research and

obtained results will be provided during the conference and included in full version of the paper.

REFERENCES

[1] El-Refaie A.M., Fractional-Slot Concentrated-Windings Synchronous Permanent Magnet Machines:

Opportunities and Challenges, IEEE Transactions on Industrial Electronics, vol. 57, no. 1, 2010,

pp. 107-121

[2] Jedryczka C., Comparative analysis of the three- and six-phase fractional slot concentrated winding

permanent magnet machines, COMPEL - The International Journal For Computation and

Mathematics in Electrical and Electronic Engineering, vol. 36, no. 3, pp. 811-823.

[3] Magnussen F., Lendenmann H., Parasitic Effects in PM Machines With Concentrated Windings, IEEE

Transactions on Industry Applications, vol. 43, no. 5, 2007, pp. 1223-1232

a) b)

a) b) c)

– 23 –

INTERPRETATION OF LOSS SEPARATION

WITH THE HALLER-KRAMER MODEL

M.F. de Campos

UFF- Federal Fluminense University – Av dos Trabalhadores 420 – 27255-125 Volta Redonda RJ, Brazil

e-mail: [email protected]

Abstract. The Haller-Kramer model for domain structure is reviewed in detail. Magnetic domains and domain walls

can be interpreted as Prigogine dissipative structures. As consequence, processes of reversal of magnetization in a

hysteresis cycle can be modeled with the minimum energy production principle. The Haller-Kramer model gives

physical basis for the loss separation procedure.

I. INTRODUCTION

The loss separation model has been widely used since at least 1936 [1]. With the aim of

clarifying loss separation, the Haller-Kramer model [2],[3] will be discussed in detail. The

minimum energy production principle of Prigogine [4] is the basis of the Haller-Kramer model

[2],[3].

II. THE HALLER-KRAMER MODEL AND THE LOSS SEPARATION

Haller and Kramer [2] based their analysis on the existence of different dissipative processes,

one due to eddy currents (Eeddy) and another due to creation and annihilation of domain walls

(Ewall), see Eqs. (1,2). Ewall = n A where A is domain area area, which is given by the product A= e

w and is domain wall energy. Haller and Kramer [2] assumed that the system is in quasi-

stationary state, see Eq. (3). The number of domain walls (n) is function of frequency (f), as

experimentally observed, see Eq. (4) [3]. The model of Haller and Kramer [2] is for only one grain

with length L, thickness e and width w, see Fig. 1. is a constant found from the WSK theory [5].

For a half cycle (f/2), Eeddy is given by Eq. (5).

Fig.1. Scheme of the domain structure in the Haller-Kramer model, with neighbor domains magnetized

in opposite directions.

AnEn

Em

1

1 (1)

DLDMEs

/)/(7.1 2 (2)

0)(

dn

EEd

dn

dE walleddy

(3)

n

fLE

eddy4

22 (4)

dtvnE

f

eddy

2

2/

0

(5)

– 24 –

The domain wall velocity v is given by Eq. (6), for sinusoidal waveform. The constant is given

by Eq. (7) [5]. c is speed of the light, is resistivity, Bs is the induction of saturation. From Eq. (3),

the equilibrium distance between domain walls D(f) is found, see Eq. (8).

)2cos( tfn

fLv

(6)

23

21605.1

c

BeAs

(7)

f

A

fn

LfD

2

)()( (8)

One of the main predictions of Eq. (8) is that D(f) varies with 1/√f. This law has been

experimentally verified [3]. This is used in the model for anomalous losses given by Eq. (9) where

a is an experimental parameter [6].

2/32

max

22/1 11fBeL

naP

an

(9)

The total losses are Pt=Ph+Pcl+Pan. where Pt is the total experimental losses, and Pclas is

(10), and Ph is Eq (11). Expression (10) assumes perfect flux penetration (i.e., no skin effect),

constant permeability, and is valid only for small frequencies, less than 400 Hz

22

max

2

2 1

6fBeP

clas

(10)

HdBfPh (11)

Alternative expressions for the total losses Pt are given by the three losses terms are Eqs. (12)

and (13). Ch, Ce and Ca are experimental constants. q and m are non-dimensional, and

experimentally determined.

2/322 fBCfBCfBCP m

ae

q

ht (12)

2/3

)2/3(

2

)1(

dt

dBBC

dt

dBC

dt

dBBCP m

ae

q

ht (13)

REFERENCES

[1] Legg V.E., Bell System Technical Journal, vol. 15, no. 1, 1936, pp. 39-62

[2] Haller T.R., Kramer J.J., J. Appl. Phys., vol. 41, 1970, pp. 1034

[3] Haller T.R., Kramer J.J., J. Appl. Phys., vol. 41, 1970, pp. 1036

[4] Prigogine I., Introduction to the Theory of Irreversible Processes, 3rd

Edition, John Wiley and

sons, New York, 1967, p. 83

[5] Williams H.J., Shockley W., Kittel C., Phys. Rev., vol. 80, 1950, pp. 1090-1094.

[6] de Campos M.F., Teixeira J.C., Landgraf F.J.G., J. Magn. Magn. Mat., vol. 301, 2006, pp. 94

http://onlinelibrary.wiley.com/doi/10.1002/bltj.1936.15.issue-1/issuetoc

– 25 –

PREDICTING RECOIL CURVES IN STONER-WOHLFARTH

ANISOTROPIC MAGNETS

M.F. de Campos and J.A. de Castro

Universidade Federal Fluminense, Volta Redonda 27255-125, Brazil, e-mail: [email protected]

Abstract. It is possible to predict recoil curves in magnetic materials following the Stoner-Wohlfarth behavior. In the

presented example, it is discussed the shape of recoil curves predicted for anisotropic Stoner-Wohlfarth magnets,

which follow distribution of type f=cosn(), with n=15.

I. INTRODUCTION

The Stoner-Wohlfarth (SW) model [1] has several assumptions, among them phases with

uniaxial easy axis and non-interacting particles. High coercivity Sm2Co17 type magnets have shown

behavior close to the SW prediction [2,3]. Here it is discussed how to predict recoil curves for

anisotropic Stoner-Wohlfarth magnets [4,5].

II. MODEL FOR RECOIL CURVES OF HYSTERESIS

According to the SW model, there are regions of reversible and irreversible rotation. First it is

defined h=H/HA the reduced field and m=M/Ms the reduced magnetization. HA is the anisotropy

field and MS is the magnetization of saturation. The critical field hc for irreversible rotation as

function of grain orientation is given by Eq. (1), with t given by Eq. (2). is the angle between

applied field and crystal easy axis. The hc is plotted in Fig. 1.

2

2/142

1

)1(

t

tth

c

(1)

3/1)(tant (2)

Fig.1. Reduced field hc as function of an angle

From Fig. 1, it is clear that irreversible rotation only takes place for h>0.5, see Fig. 2. This

permits the prediction of recoil curves, shown in Fig. 3. As the predicted recoil curves are near the

experimentally observed for 2:17 SmCo magnets [6], this implies that the coercivity mechanism is

coherent rotation for these magnets.

0 15 30 45 60 75 900,4

0,6

0,8

1,0

hc

re

du

ce

d c

ritica

l fie

ld

Angle (degrees)

– 26 –

Fig.2. Start of irreversible rotation for an anisotropic Stoner-Wohlfarth magnet, which follows distribution f=cos

n(),

with n=15. Jr/Js=n+1/n+2, then Jr/Js=0,94 for this magnet.

Fig.3. Several possible recoil curves. It is assumed an anisotropic Stoner-Wohlfarth magnet, which follows distribution

of orientation f=cosn(), with n=15.

REFERENCES

[1] Stoner E.C., Wohlfarth E.P., IEEE Trans. Magn., vol. 27, 1991, pp. 3475

[2] Sampaio da Silva F.A. et al., J. Magn. Magn. Mat., vol. 328, 2013, pp. 53

[3] Romero S.A., et al., J. Alloys Compds., vol. 551, 2013, pp. 312

[4] de Campos M.F., et al., Materials Science Forum, vol. 775-776, 2014, pp. 431

[5] de Campos M.F., et al., J. Magn. Magn. Mat., vol. 345, 2013, vol. 147

[6] Bavendiek G., et al., in Proc. of WMM´18, Dresden, Germany, 12-14 June 2018, p. 148

-1,0 -0,5 0,0 0,5 1,0

-1,0

-0,5

0,0

0,5

1,0

-1,0 -0,5 0,0 0,5 1,0

-1,0

-0,5

0,0

0,5

1,0

start of irreversible rotation

m -

re

du

ce

d m

ag

ne

tiza

tio

n

h - reduced field

-1,5 -1,0 -0,5 0,0 0,5 1,0 1,5

-1,0

-0,5

0,0

0,5

1,0

m

h

Recoil Curve

Recoil Curve

Recoil Curve

– 27 –

METHODS FOR TEXTURE IMPROVEMENT IN ELECTRICAL STEELS

M.F. de Campos

Universidade Federal Fluminense, Volta Redonda 27255-125, Brazil, e-mail: [email protected]

Abstract. Aiming the development of high efficiency electric motors for electric vehicles, there is strong pressure for

improvement of the magnetic properties of electrical steel sheets. One of the clearest possibilities is texture

enhancement. In this review, diverse methods for texture improvement are presented and discussed. All of them have

the drawback of increasing the cost of material processing.

I. INTRODUCTION

The autonomy of electrical cars can be increased by basically three ways: (i) batteries

improvement, (ii) vehicle weight reduction (iii) increase of motor efficiency. As batteries

improvement is very difficult and slow, due to need of long time tests, the increase of motor

efficiency is the most clear alternative. Among the options for increasing motor efficiency is the

texture enhancement. There are many ways for improving the texture in electrical steels. However,

all of them may increase the cost of the steels. Other clear options to motor performance

enhancement, from the material point-of-view, are increase of resistivity, thickness reduction and

materials with zero magnetostriction.

II. METHODS FOR TEXTURE ENHANCEMENT

The ideal texture for non-oriented electrical steels is 100 <0vw>. Annealing in vacuum can

develop the 100 <001> cube on the face texture [1]. This method only can be applied for small

thickness, and is considered as very expensive. Other possible method is strip casting [2]. The

preferential direction of dendrite growth in the bcc structure - as well as in the fcc structure - is

<100> [3]. Thus, as cast materials can have the easy magnetization axis in the direction of the heat

extraction. The problem is to keep the <100> direction parallel to the sheet plane during the cold

rolling process. Cold rolling is interesting for developing the texture for deep-drawing steels 111

<uvw> [4]. This texture is opposite to that of ideal for non-oriented electrical steels.

However, rotated cube 100 <011> is stable orientation after cold rolling. Thus, if a strong

100 <011> is developed at the hot band, this texture component can be kept after the cold rolling

process [4].

The texture of recrystallization of austenite is cube on face 100 <001> [4]. Thus,

recrystallization of austenitic steels at high temperatures can generate a very favorable texture.

Alloying elements that increase the fcc region, as Mn can increase the austenite field in the Fe-C

carbon phase diagram. .However, Si and Al reduce the austenite region. Silicon is so effective as

alpha-iron stabilizer that a 2.5% Si alloy is always bcc.

Cross-rolling is a possibility [5], many times neglected because it is difficult for large scale

application. But if the sheet is cut as a square, and the cross rolling is done just before the stamping

step, then the cross-rolling may be possible.

An example of competition vehicle has chosen a FeCo alloy, named 1J22 [6], with 0.1 mm

thickness. This alloy has chemical composition 49%Fe-49%Co-2%V. This shows the relevance of

high magnetization of saturation (2.35 Tesla) for electric machines. It is important to add that it is

very easy to develop 100 textures in iron-cobalt alloys [7,8].

For non-oriented semi-processed electrical steels, however, the typically obtained texture is far

from ideal 100 <001> [9]. The typical recrystallization texture has as the most relevant

– 28 –

components Goss 110 <001> and 111 <uvw> [9,10]. The presence of Goss increases the

magnetic induction at the rolling direction, but also introduces strong anisotropy on the sheet [11].

As main conclusion, there is significant space for texture improvement in commercial

electrical steels, however, with the drawback of increasing the cost of processing.

ACKNOWLEDGEMENTS

CNPq, FAPERJ.

REFERENCES

[1] Assmus F., Detert K., Ibe G., Über eisen-silizium mit würfeltextur, Z. Metallk., vol. 48, 1957, p. 344-

349

[2] Landgraf F.J.G., Yonamine T., Takanohashi R., Silva F.Q., Tosetti J.P.V., Beneduce Neto F.,

Albertin E., Mazzarella V.N.GFalleiros., I.G.S., Emura M., Magnetic properties of silicon steel with

as-cast columnar structure. J. Magn. Magn. Mat., vol. 254-255, 2003, pp. 364–366

[3] Dantzig J.A., Rappaz M., Solidification, CRC Press, 2009

[4] Ray R.K., Jonas J.J., Transformation textures in steels, Int. Mat. Rev., vol. 35, no. 1, 1990, pp. 1-36

[5] Mekhiche M., Waeckerlé T., Cornut B., Influence of low Al content on anomalous growth in 3% Si-Fe

magnetic sheets, J. Magn. Magn. Mat., vol. 133, 1994, pp. 159-162

[6] Xuanyang Hu, Hong Guo, Hao Qian, Xiaofeng Ding, Yanling Yang, Development of a high-power-

density motor for Formula SAE electric race car, in Proceedings of IECON 2017 - 43rd Annual

Conference of the IEEE Industrial Electronics Society, 29 Oct.-1 Nov. 2017, Number: 17432419

[7] Foster K., Thornburg D.R., Magnetic properties of oriented iron‐cobalt alloys, AIP Conference

Proceedings, vol. 24, 1975, pp. 709

[8] Heck C., Magnetic Materials and their Applications, Newnes-Butterworth, 1974

[9] de Campos M.F., Landgraf F.J.G., Falleiros I.G.S., Fronzaglia G.C., Kahn H., Texture Evolution during

the Processing of Electrical Steels with 0.5% Si and 1.25% Si, ISIJ International, vol. 44, 2004,

pp. 1733-1737

[10] de Campos M.F., Yonamine T., Fukuhara M., Landgraf F.J.G., Achete C.A., Missell F.P., Effect of

frequency on the iron losses of 0.5% and 1.5%Si non-oriented electrical steels, IEEE Trans. Magn.,

vol. 42, 2006, pp. 2812

[11] de Campos M.F., Anisotropy of Steel Sheets and Consequence for Epstein Test: I Theory, in XVIII

IMEKO WORLD CONGRESS Metrology for a Sustainable Development, 17-22 September 2006, Rio

de Janeiro, Brazil

Available at: http://www.imeko.org/publications/wc-2006/PWC-2006-TC4-037u.pdf

https://ieeexplore.ieee.org/xpl/mostRecentIssue.jsp?punumber=8168197


https://aip.scitation.org/author/Foster%2C+K

https://aip.scitation.org/author/Thornburg%2C+D+R

https://www.jstage.jst.go.jp/article/isijinternational1989/44/10/44_10_1733/_article/-char/en

https://www.jstage.jst.go.jp/article/isijinternational1989/44/10/44_10_1733/_article/-char/en

http://www.imeko.org/publications/wc-2006/PWC-2006-TC4-037u.pdf

– 29 –

EXPERIMENTAL DETERMINATION OF PREISACH MODEL FOR

GRAIN ORIENTED STEEL

J. Eichler, M. Novak and M. Kosek

Technical University of Liberec, Studentska 2, 46117 Liberec I. Czech Republic, e-mail: [email protected]

Abstract. Full material characteristics in Preisach model of hysteresis is the weighting function. It can be determined

experimentally from systematic measurement of partial hysteresis loops by derivation of their decreasing parts.

Because of measurement errors, the derivation is not correct. Nevertheless basic material features can be obtained

either from incomplete measurement that uses Preisach triangle respecting the measurement errors.

I. INTRODUCTION

The Preisach model [1] is a very suitable one for complete description of hysteresis. Its basic

elements are hysterons exhibiting ideal rectangular loop. The magnetic field strength for switching

up and down are Hu and Hd (Hu >= Hd), respectively. They are systematically arranged in Preisach

triangle according to Hu and Hd. Increasing external field is represented by horizontal line moving

up that switches hysterons up. Decreasing field moves the vertical line moving left that switches

hysterons down. By this way the hysteresis is ensured.

A full description of material is given by the weighting function that defines magnetic

momentum of hysterons. The weighing function can be determined from systematic measurement

of partial loops by the FORC (First Order Reverse Curves) method [1]. The decreasing branches of

the loops (FORC) are used to form an Everett surface. Weighting function is given by partial

derivations of the Everett surface by both the field strengths Hu and Hd. Due the experimental

inaccuracy the derivation exhibits errors. The accuracy of weighting function and Preisach model

application is a subject of the paper.

II. BASIC RESULTS

The current source was used to ensure the harmonic field strength excitation. The well defined

starting level was the negative saturation. The time varying field strength for partial loop is Hd,

while its systematically increasing amplitude (up to the positive saturation) is Hu. The amplitude Hu

should increase by the smallest possible step because of precise determination of values especially

in the neighborhood of the weighting function main peak. Basic experimental limitation of the

procedure is demonstrated in Fig. 1.

Fig.1. Excitation and response in time domain.


– 30 –

Total 1800 loops ware measured. Several of them in time domain with loop number increasing

by 4 are in Fig. 1. In upper part there is the excitation, in lower the response. In the left hand part

there are details at maximum t = T/2, while in the right hand part the area for t = 3/4.T is shown.

The points in vertical cuts are in graphs in order to represent the quality of excitation and

measurement. For excitation at FORC starting the curves are well defined FORC centre the

excitation has only small deviation. The response is measured with lower accuracy. The curve

numbers are preserved at FORC start. But in the FORC centre the monotonic increase does not

exist and either impossible negative derivation takes a place.

The result of numeric derivation of Everett surface is in Fig. 2. The only cut at the plane in Hu

= 0 is shown for clarity. It contains main peak and values near zero. Insets in Fig. 2 reveal that the

peak is sharp and the most of area exhibits noise containing negative values.

Fig.2. Weighting function determined by numeric derivation

Hysteresis loops reconstructed from reduced weighting function are in Fig. 3. The loop

trajectory is not considerably changed by reduction except quantization. The reduction causes

major deviations in the central part, while in the area of saturation is negligible, see insets.

Fig.3. Hysteresis loops with strong reduction of number of hysterons (rows and columns).

III. CONCLUSION

Experimental inaccuracy limits the number of points in Preisach triangle. The distance

between rows and columns should be greater than the estimated experimental error. Fortunately,

this limitation does not affect the Preisach model prediction considerably.

REFERENCES

[1] Bertotti G., Mayergoyz I., The science of hysteresis, Elsevier, 2006 (1st ed.)

– 31 –

PROPAGATION OF TERA-HERTZ RADIATION IN FOGGY CONDITIONS

A. Etinger1, Y. Golovachev

1, G.A. Pinhasi

2 and Y. Pinhasi

1

1 Department of Electrical and Electronic Engineering, Ariel University, P.O. Box 3, Ariel 40700, Israel

e-mail: [email protected] 2 Department of Chemical Engineering, Ariel University, P.O. Box 3, Ariel 40700, Israel

Electromagnetic radiation at millimeter and sub-millimeter (Tera-Hertz) wavelengths are being

considered for various applications, including remote sensing, wireless communications and radars.

However, wireless links implemented in millimeter wavelengths above 30GHz suffer from

absorption and dispersion effects in air, which emerge mainly due to Oxygen molecules, humidity

and suspended water droplets. Such frequency dependent atmospheric propagation effects become

more severe as the frequency is raised to the Tera-Hertz regime. Moreover, weather conditions like

haze, fog and rain cause a further decrease in the overall link-budget leading to a degradation in the

channel performance.

In this paper, we analyze the performance of a link operating in the J-band within the

sub-millimeter wavelengths. Expressions for the attenuation and group delay are presented as

a function of the density of the fog. The analytical estimations are verified experimentally in

a controlled artificial fog chamber. Attenuation and group-delay were measured using a wide band

sub-millimeter radar for several degrees of visibility even below 1m.


– 32 –

– 33 –

A NEW PARAMETER FOR THE BARKHAUSEN NOISE

CHARACTERIZATION

T. Garstka

Częstochowa University of Technology, Faculty of Production Engineering and Materials Technology

Al. Armii Krajowej 19. 42-200 Częstochowa, Poland, e-mail: [email protected]

Abstract. In this paper, a new empirical parameter for magnetic Barkhausen noise characterization has been

described. The definition, method of measurement and results of its study in the function of the stresses, microstructure

state (grain diameter) and magnetization conditions were presented. It peculiarity is the fact, that is more sensitive to

the microstructure changes than to changes in applied or residual stress state. For this reason can be useful for solving

the main problem during the residual stress measurements that is the taking into consideration microstructure's

influence on Barkhausen noise.

I. INTRODUCTION

The sensitivity of the Barkhausen phenomena to the changes in the material properties as

microstructure or internal stress state is utilized in many magnetic non-destructive methods of

testing ferromagnetic materials and products. Unfortunately, Barkhausen noise (BN) parameters

most used for calibration procedure during residual stress measurements, as the Root Mean Square

(RMS) value, number of counted Barkhausen jumps (BJ) or the BN power spectrum [1-3] are

sensitive for both of these properties. It causes the results of the residual stress investigations may

be disrupted and falsified if the tested product has heterogeneous microstructure. For this reason,

calibration should be conducted not only as the function of the stress state and magnetization

conditions but also prepared on the samples with different microstructure. During the proper

measurements on the real objects, to recognize microstructure state in investigated region, its

measurable indicator is needed. As an effect of wide research on it, new empirical parameter of BN

has been developed which seems to be more sensitive to the changes in microstructure than to the

stress state.

II. NEW PARAMETER CHARACTERIZATION

The elaborated new parameter for BN characterization utilizes and joins two parameters,

mentioned above and mainly used for calibration during stress measurements by BN method: so-

called “digital” – amount of Barkhausen jumps (pulses) with amplitude over reference voltage and

energetic - integrated RMS value of the Barkhausen noise. The main idea during its development

was to replace the constant threshold voltage level for BJ discrimination by the dynamically

changing with stress BN RMS voltage.

By definition, this new parameter NRMS can be described as: The amount of measured

Barkhausen pulses with voltage amplitude greater or equal than RMS value of Barkhausen noise,

counted in specified period of time; and expressed mathematically by equation (1)

e

st

BNBj

BNBjiiRMS

t

RMS<Aif

RMSAif=s;s=N

0

1 (1)

where:

si – binary parameter of Barkhausen jump, depends on its amplitude and root mean square value of

BN,

ABj – amplitude of the particular Barkhausen jump,

ts – measurement start time,

te – measurement end time.


– 34 –

III. EXPERIMENTAL AND CONCLUSION

For practical determination value of described new BN parameter and its testing, a special

electronic circuit within measurement apparatus for Barkhausen noise measurement [4] was

created. It consists of integrated RMS converter, comparator and counter. As the measurement

period, time of one magnetization cycle was used. During initial experiment, two samples made

from the same steel grade S235JGR2 but with different average ferrite grain size (11 μm in sample

1 and 19 μm in sample 2 respectively) created by heat treatment were used. To apply tensile and

comprehensive stresses, the specimens were deflected in special equipment for uniaxial bending

[5]. Observation changes of the course of NRMS with stress (Fig.1) let to conclude, that in wide

range their course seems to be flat and exists expressive differentiation between both lines. Due to

this, its value can be assumed as invariant from the stress state. Further investigations confirmed its

usefulness in accurate investigations of residual stress by the BN method with multiparameter

calibration taking into consideration microstructure state.

Fig.1. Variation of the NRMS with applied stress σ in two samples with different grain size

IV. ACKNOWLEDGMENTS

This paper was financed within scientific work No. BS/PB-201-304/2018

REFERENCES

[1] Jagadish C., Clapham L., Atherton D.L.: Influence of uniaxial elastic stress on power spectrum and

pulse height distribution on power spectrum and pulse height distribution on surface Barkhausen noise

in pipeline steel, IEEE Trans. Magn., vol. 26, no 3, 1990, pp.1160-1163

[2] Matzkanin G.A., Gardner C.G.: Measurement of residual stresses using magnetic Barkhausen noise,

Proceedings of ARPA/AFML Rev. Quant. NDE, AFML-TR-75-212, 1976, pp. 791-813

[3] Grum J., Zerovnik P.: Use of the Barkhausen effect in the measurement of residual stresses in steel,

INSIGHT, vol. 42, no 12, 2000, pp. 796-800

[4] Garstka T., The complex system for residual stress determination based on Barkhausen noise

measurement, Proceedings of 5th International Conference in Barkhausen Noise and Micromagnetic

Testing, Petten, The Netherlands, 2005, pp. 219-228

[5] Garstka T., Microstructure state and heat treatment influence on Barkhausen noise parameters and

residual stress measurement, Solid State Phenomena, vol. 165, 2010, pp. 50-55

– 35 –

INFLUENCE OF DC-BIAS MAGNETIC FIELD ON DYNAMIC MAGNETIC

PROPERTIES OF LaFeCoSi ALLOY

R. Gozdur1, P. Gębara

2 and K. Chwastek

3

1 Department of Semiconductor and Optoelectronics Devices, Lódź University of Technology, Wólczańska 211/215,

Lódź, 90-924, Poland, e-mail: [email protected] 2 Institute of Physics, Częstochowa University of Technology, Armii Krajowej 19, Częstochowa, 42-200, Poland,

email: [email protected] 3 Faculty of Electrical Engineering, Częstochowa University of Technology, A1. Armii Krajowej 17, Częstochowa,

42-200, Poland, e-mail: [email protected]

Abstract. The paper presents an experimental study of magnetic properties of LaFeCoSi alloy in the ferromagnetic

state close to Curie temperature of 306 K. Influence of DC bias magnetic field on dynamic hysteresis loops and power

losses was determined. The investigated alloy has tenfold drop of losses during biased magnetization.

I. INTRODUCTION

The interest in the magnetocaloric effect (MCE) has been steadily increasing since giant MCE

was discovered in gadolinium [1]. MCE is characterized by a high value of entropy changes in the

room temperature range only in very few alloys. The most promising magnetocaloric materials are

alloys containing Gd, La and MnAs [1-3]. MCE observed in La-containing alloys is slightly

weaker in comparison to MCE in Gd [4]. However, reasonable price, excellent physical properties

and low environmental impact distinguish this material for further development and its

applications. The experimental study of LaFeCoSi magnetocaloric alloy in DC bias magnetic field

gives better insight into estimation of magnetic power losses under real operating conditions [5].

II. SAMPLE AND MEASUREMENTS

The cast of the sample was made of LaFe10.92Co1.08Si1.2 alloy (Fig.1a) with Curie point at a

temperature of 306 K. The final-form of the tested ring core with overall dimensions

OUT=8.9 mm, IN= 3.7 mm, h=5.8 mm has been obtained after mechanical processing (Fig.1b).

The weight of the core was 2.021 g.

Fig. 1. a) XRD pattern of the alloy from Bruker D8 X-ray diffractometer with LynxEye detector.

b) View of the ring core applied for the tests.

The measurements of the hysteresis loops, magnetic polarization and power losses were

carried out at a temperature of 292 K and range of magnetizing frequency from 0.1 Hz to 10 Hz.

The experimental study has been done according to IEC 60404-6 standard.

a) b)




– 36 –

Fig. 2. Influence of frequency and DC bias on dynamic hysteresis loops of LaFe10.92Co1.08Si1.2; a) Sinusoidal

magnetic field strength without DC bias. b) Sinusoidal magnetic field strength with 0.5APk-Pk DC bias.

Fig. 3. Influence of DC bias magnetic field on specific power losses Ps/f in the range of frequency from 0.1Hz to 10Hz.

Operation conditions of magnetocaloric refrigerators are based on magnetization-

demagnetization cycles in rotating and reciprocating magnetizing systems. Biased magnetizing

field (unipolar waveform) is in compliance with real waveforms while the approach based on

bipolar magnetizing field is recommended during tests of soft magnetic materials. The

measurements of magnetic properties (Fig.2a, 2b, 3) were carried out with the same HPk-Pk of

magnetic field strength.

II. SUMMARY

The experimental study of LaFe10.92Co1.08Si1.2 alloy confirms strong influence of DC bias

magnetic field on its magnetic properties. The most significant effect of biased magnetization is

illustrated by the curves of specific power losses. Direct application of IEC, ASTM measuring

requirements is not appropriate for testing the magnetic properties of LaFeCoSi magnetocaloric

materials.

REFERENCES

[1] Pecharsky V. K., Gschneidner Jr. K. A., Giant magnetocaloric effect in Gd5(Si2Ge2), Phys. Rev. Lett.,

vol. 78, 1997, pp. 4494

[2] Brück E., Ilyn M., Tishin M., Tegus O., Magnetocaloric effects in MnFeP1−xAsx-based compounds,

J. Magn. Magn. Mater., vol. 290-291, Part 1, 2005, pp. 8-13

[3] Fujieda S., Fujita A., Fukamichi K., Large magnetocaloric effect in La(FexSi1-x)13 itinerant-electron

metamagnetic compounds, Appl. Phys. Lett., vol. 81, 2002, pp. 1276-1278

[4] Bjørk R., Bahl C.R.H., Katter M., Magnetocaloric properties of LaFe13−x−yCoxSiy and commercial grade

Gd, J. Magn. Magn. Mater., vol. 322, no. 24, 2010, pp. 3882-3888

[5] Sandeman K.G., Magnetocaloric materials: The search for new systems, Scr. Mater., vol. 67, no. 6,

2012, pp. 566-571

a) b)

– 37 –

MAGNETIC SUBSTRATES MADE OF SPUTTERED Y3Fe5O12

B. Guzowski1, R. Gozdur

2 and A. Kociubiński

3

1 Lodz University of Technology, Department of Semiconductor and Optoelectronics Devices, Wolczanska 211/215,

90-924 Lodz, Poland, [email protected] 2 Lodz University of Technology, Department of Semiconductor and Optoelectronics Devices, Wolczanska 211/215,

90-924 Lodz, Poland, [email protected] 3 Lublin University of Technology, Institute of Electronics and Information Technology, Nadbystrzycka 38A,

20-618 Lublin Poland, [email protected]

Abstract. The paper presents detailed investigation of sputtered Y3Fe5O12 (YIG) films dedicated to spintronic devices.

Magnetic properties of developed films were analyzed and compared with magnetic properties of pure YIG target.

I. INTRODUCTION

Yttrium iron garnet – Y3Fe5O12 (YIG) is a widely used garnet because of its excellent

parameters such as: low microwave loss, high resistivity and good transparency. Therefore YIG is

irreplaceable in microwave [1], optoelectronics [2], magneto-optical [3] and magnetic [4]

applications.

In recent years YIG became attractive in another field of science – spintronic [5, 6]. YIG with

Pt has high efficiency of the spin Hall effect and because of high resistivity, generation of pure spin

waves in YIG is not disturbed by Nernst-Ettingshausen effects [7].

Nowadays YIG films are fabricated by various methods: liquid phase epitaxy (LPE) [8], RF

magnetron sputtering [9] or pulsed laser deposition (PLD) [10] and one of the main objective of the

conducted research is to decrease the cost of YIG substrates. In this paper detailed investigation of

the sputtered Y3Fe5O12 substrates is given.

II. SAMPLE PREPARATION AND MEASUREMENTS

During research three samples shown in Fig. 1. were investigated. In Sample 1 200 nm thick

YIG on glass was used, while in Sample 2 and Sample 3 100 nm thick YIG film was sputtered on

Al2O3 and glass respectively. As a reference sample piece of pure YIG target was used.

Fig.1. Sample 1: 200 nm YIG on glass (a), Sample 2: 100 nm YIG on Al2O3 (b), Sample 3: 100 nm YIG on glass (c)

TABLE I: The nominal and the measured YIG composition, wt. %

Element O Fe Y

Nominal 26.75 24.56 38.7

Measured 28.38 56.82 14.8

The thin films of YIG were sputtered by deposition system Nano 36 from Kurt J-Lesker. The

developed substrates were characterized by EDS probe X-MAX N80 from Oxford Instruments and

scanning electron microscope. The measurements and calculations of percentage weights of YIG

films are collected in Tab. 1. The magnetic properties of sputtered YIG films were recorded by

VersaLab System (Quantum Design). The measured results for YIG thin films are shown in Fig. 2a

while in Fig. 2b reference hysteresis loop measured for pure YIG target is given.




– 38 –

Fig. 2. Set of hysteresis loops measured for developed YIG films (a), measured hysteresis loop of pure

YIG target (b)

III. SUMMARY

In this paper magnetic properties of thin YIG films were fabricated with magnetron sputtering

process. YIG films were deposited on glass and Al2O3 in the same technological process. Based on

the obtained results it can be concluded that YIG films on glass have much worse magnetic

properties in comparison to YIG films deposited on Al2O3. YIG films sputtered on ceramic have

very similar magnetic properties to measured properties of pure YIG target. Magnetron sputtering

seems to be a low cost method to fabricate the YIG substrate for spintronic devices, however,

further research must be taken to improve composition of the deposited films.

REFERENCES

[1] Sharma V., Saha J., Patnaik S., Kuanr B.K., YIG based broad band microwave absorber: A perspective

on synthesis methods, Journal of Magnetism and Magnetic Materials, vol. 439, no. 1, 2017, pp. 277-

286

[2] Ghosh S., Keyvavinia S., Van Roy W., Mizumoto T., Roelkens G., Baets R., Ce:YIG/Silicon-on-

Insulator waveguide optical isolator realized by adhesive bonding, Optics Express, vol. 20, no. 2, 2012,

pp. 1839-1848

[3] Boudiar T., Payet-Gervy B.,. Blanc-Mignon M.-F, Rousseau J.-J., Le Berre M., Joisten H., Magneto-

optical properties of yttrium iron garnet (YIG) thin films elaborated by radio frequency sputtering,

Journal of Magnetism and Magnetic Materials, vol. 284, 2004, pp. 77-85

[4] Bandyopadhyay A.K., Rios S.E., Fritz S., Garcia J., Contreras J., Gutierrez C.J., Ion beam sputter-

fabrication of Bi-YIG films for magnetic photonic applications, IEEE Transactions on Magnetics,

vol. 40, no. 4, 2004, pp.2 805-2807

[5] Uchida K., Xiao J., Adachi H., Ohe J., Takahashi S., Ieda J., Ota T., Kajiwara Y., Umezawa H.,

Kawai H., Bauer G.E.W., Maekawa S., Saitoh E., Spin Seebeck insulator, Nature Materials, vol. 9,

2010, pp. 894-897

[6] Uchida K., Adachi H., Ota T., Nakayama H., Maekawa S., Saitoh E., Observation of longitudinal spin-

Seebeck effect in magnetic insulators, Applied Physics Letters, vol. 97, 2010, pp. 172505-3

[7] Boona S.R., Myers R.C., Heremans J.P., Spin caloritronics, Energy Environmental Science, vol. 7,

2014, pp. 885-910

[8] Blank S.L., Nielsen J.W., The growth of magnetic garnets by liquid phase epitaxy, Journal of Crystal

Growth, vol. 17, no. 302, 1972, pp. 302-311

[9] Stadler B., Gopinath A., Magneto-optical garnet films via reactive sputtering, Transactions on

Magnetics, vol. 36, 2000, pp. 3957-3961

[10] Karim R., Oliver S.A., Vittoria C., Laser ablation deposition of YIG films on semiconductor and

amorphous substrates, Transactions on Magnetics, vol. 31, 1995, pp. 3485-3487

– 39 –

A COMPARISON OF TWO PHENOMENOLOGICAL DESCRIPTIONS

OF MAGNETIZATION CURVES BASED ON T(X) MODEL

R. Jastrzębski, A. Jakubas and K. Chwastek

Czestochowa University of Technology, Faculty of Electrical Engineering,

Al. Armii Krajowej 17, 42-201 Czestochowa, e-mail: [email protected]

Abstract. The paper considers the effect of compaction pressure on the shape of magnetization curves of soft magnetic

composite cores compacted at different compaction pressures. Two versions of the phenomenological Takács T(x)

model were taken into account.

I. INTRODUCTION

The paper focuses on the issue of phenomenological modeling of magnetization curves for soft

magnetic composite cores (SMC) with two descriptions derived from the Takács hysteresis model

[1]. The first description includes a linear term to take into account reversible magnetization

processes, as suggested by the model developer. The other model neglects the reversible

magnetization term, however it considers mutual interactions between magnetic domains within the

material, given in the first approximation as the so called Weiss’ coefficient [2,3]. The aim of the

paper is to elucidate which description is more adequate for modeling.

The Takács model is a relatively simple description based on hyperbolic tangent mapping

between the output and input variables. Its foundations are well described in the textbook [1]. The

description was used previously for modeling hysteresis curves of commercial SMC cores in the

publications [4,5]. However only the second version of the model was considered.

REFERENCES

[1] Takács J., Mathematics of hysteretic phenomena, J. Wiley & Sons, Weinheim, 2003.

[2] Chwastek K., A dynamic extension to the Takács model, Physica B, vol. 407, no. 17, 2010, pp. 3800-

3802

[3] Jakubas A., Modeling of the effect of grain size on hysteresis curves using the Takács model, in Progress in Applied Electrical Engineering (PAEE), 18-22.06.2018 Kościelisko, Poland,

[4] Ślusarek B., Chwastek K., Jankowski B., Szczygłowski J., Modeling hysteresis loops of SMC cores,

Solid State Phenomena, vol. 220-221, 2015, pp. 652-660

[5] Ślusarek B., Szczygłowski J., Chwastek K., Jankowski B., A correlation of magnetic properties with

material density for soft magnetic composite cores, COMPEL, vol. 34, no. 3, 2015, pp. 636-646





– 40 –

– 41 –

THE MAGNETIZATION CURVE OF FeNiCo ALLOY AND THE

INFLUENCE OF ITS APPLICATION ON THE PENNING EFFECT IN

VACUUM INTERUPTOR

D. Kapelski, E. Kucal and A. Szymański

Tele and Radio Research Institute, Ratuszowa 11 St., 03-540 Warsaw, Poland,

e-mails: [email protected], [email protected]

Abstract.The research conducted at the Tele- and Radio Research Institute covers a developing a new portable vacuum

gauge for using in a vacuum circuit breaker. A step after that is its integration with new generation vacuum circuit

breakers.The new device is supposed to use the Penning phenomenon. In this phenomenon, glow charge is amplified by

an external magnetic field. In the Pening method, a constant magnetic field is used to generate a magnetic field. The

glow discharge and its strengthened is proportional to vacuum pressure and magnetic field induction inside vacuum

interrupter chamber.Vacuum chambers based on a glass insulator, must be made of FeNiCo alloy called kovar. Kovar

has a high magnetic permeability, but its mechanical and magnetic parameters can strongly depend on heat and

mechanical treatment [1]. The paper presents results of investigation on influence of complex thermal-mechanical

treatment on magnetization curve of FeNiCo alloy. Applied samples have undergone a similar process of heat,

chemical and plastic treatment as the structural element of vacuum chambers with glass insulators. In addition,

simulation studies of magnetic induction in a chamber with kovarconstructional elements were presented.

I. INTRODUCTION

In the vacuum chambers with a glass insulator in vacuum circuit breaker, it is necessary to use

a construction elements and a vapor condensation shieldmade ofkovar. Construction of a vacuum

chamber was shown in figure 1. The condensation screen protects the insulating material from

dusting the metal evaporated from the contacts. Protection of the internal insulating surfaces

against metal sputtering, condensation of metal vapors[2].

Fig.1. Cross section of vacuum interrupter: a) fixed contact stem, b) ceramic or glass insulator,

c) vapor condensation shield, d)Copper-Tungstencontacts discs, f) moving contact stem, g) bellows

Kovar is a nickel–cobalt-ferrous alloy compositionally identical to Fernico 1, designed to have

substantially the same thermal expansion characteristics as borosilicate glass, in order to allow

a tight mechanical joint between the two materials over a range of temperatures. Kovar was

invented to meet the need for a reliable glass-to-metal seal, which is required in electronic devices



– 42 –

such as light bulbs, vacuum tubes, cathode ray tubes, and in vacuum systems in chemistry and

other scientific research[2]. It finds application in glass-to-metal seals in vacuum interrupter made

in Tele and Radio Research Institute.

Table 1. Typical composition of kovar given in percentages of weight.

Fe Ni Co C Si Mn

balance 29% 17% < 0.01% 0.2% 0.3%

The research conducted at the Tele- and Radio Research Institute covers a developing a on-line

detection of residual gases inside the chamber of the vacuum interrupter. This method is using of

the Penning effect.

The Penning effect is used in vacuum gauges, but also for testing vacuum interrupter[3]. The

Panning gauge is an ionization gauge with an unheated cathode in which a discharge is maintained

between two electrodes with a potential difference of a few kilovolts. Pressure is converted from

discharge current. Magnetic field is applied to increase the number of ions produced during

discharge. An axial magnetic fields cause electrons to move in spiral path and increase the

ionization current. The longer path length of an electron from cathode to anode increased

possibility of generate another electron by impacting on a gas molecule to maintain the discharge.

Penning vacuum measurement method based on generating an axial field inside the vacuum

interrupter and applying high voltage to one of the contacts. The magnetic field is a gain for the

ionic current resulting from the emission of electrons.

The vacuum interrupter manufactured in ITR are mainly made of kovar, glass and copper.

Kovar from which the housing elements and sometimes screens are made can influence the

distribution of the magnetic field during the Penning method test. Research include measurements

of magnetization characteristics of samples made of kovar. In addition, simple FEM simulation of

the magnetic field distribution in the chamber with elements made of kovar were carried out.

REFERENCES

[1] https://en.wikipedia.org/wiki/Kovar

[2] Sibilski H., Dzierżyński A., Berowski P., Hejduk A., Krasuski K., Grodziński A., Szymański A., AMF

contact research in a dismountable vacuum chamber, Elektronika: konstrukcje, technologie,

zastosowania, vol. 52, 2011, pp. 8

[3] Huiyong M., Guang Ch., Xuegui Z., Wang, Y., On-line monitoring of pressure in vacuum interrupters,

IEEE Transactions on Dielectrics and Electrical Insulation, vol. 14, 2017, pp. 179-184

– 43 –

MODELING OF THE MAGNETIC FIELD DISTRIBUTION IN AIR GAP

OF THE SYNCHRONOUS MACHINE FROM PERMANENT MAGNETS

IN THE ROTOR

A. Kapłon and J. Rolek

Department of Power Electronic, Electrical Machines and Drives

Faculty of Electrical Engineering, Automatic Control and Computer Science

Kielce University of Technology

al. Tysiąclecia Państwa Polskiego 7, 25-314 Kielce, Poland

e-mail: [email protected], [email protected]

Abstract. The paper presents the magnetic field distribution from permanent magnets in the rotor core for different

rotor configurations in the sense of material and shape allowing to obtain near-sinusoidal distribution of magnetic

induction in the air gap of the machine.

I. INTRODUCTION

In synchronous machines with permanent magnets, classic (PMSM) or Line-Start (LSPMSM),

it is important the distribution of the magnetic field in the air gap from these magnets. In the

majority of currently used constructional solutions, the distribution is rectangular. Such

a distribution is unfavorable from the point of view of higher harmonics generated in SEM,

currents and electromagnetic torque of the machine. The limitation of these unfavorable

phenomena can be obtained by a sinusoidal distribution of the magnetic field from permanent

magnets. Taking into consideration the material and technological possibilities in the production of

both soft and hard magnetic materials, the following solutions are possible:

a) proper profiling of a homogeneous soft magnetic material of the rotor core ensuring,

at magnets with rectangular characteristic, the desired induction distribution in the air gap,

b) appropriate shaping of the soft magnetic material properties of the rotor core that ensures

in the air gap the desired induction distribution from permanent magnets with rectangular

characteristic,

c) suitable construction of the permanent magnet (eg. induction distribution, shape) providing

the desired induction distribution in air gap of the machine with a homogeneous

magnetically soft rotor material.

The most commonly used solution (a) comes to cutting in the rotor sheets properly profiled

additional gaps. Such a solution is not favorable from the point of view of mechanical properties of

the rotor. Solution (b) is difficult to perform from a technological point of view in the case of

classical construction from packaged sheets. It becomes feasible, however, when using

a composition of appropriately selected powders from soft magnetic materials in the 3-D printing

technology of the rotor magnetic circuit. Solution (c) comes to the proper implementation of the

magnet, using in the construction of a single magnet both the appropriate composition of hard

magnetic materials and their appropriate configuration.

II. THE MAGNETIC FIELD DISTRIBUTION IN A ROTOR

The subject of the article is the analysis of the magnetic field distribution from permanent

magnets in the machine rotor for the three presented solutions. Magnetic field distribution was

determined by MES using the FEMM and ANSYS software. The figures show examples of the

magnetic field distributions in the rotor cross-section and the normal magnetic induction

component in the machine's air gap.



– 44 –

Fig.1. Magnetic field distribution in powder hybrid rotor.

Fig.2. Magnetic field distribution in rotor with segmented permanent magnets.

CONCLUSIONS

The presented analysis shows that in each of the discussed solutions it is possible to obtain the

optimal, from the point of view of the desired properties, distribution of the magnetic field in the

machine's air gap. The best solution presented seems to be solution (c). Modern material and

technological capabilities allow the construction of this type of magnets.

REFERENCES

[1] Jedryczka C., Wojciechowski R.M., Andrzej Demenko A., Influence of squirrel cage geometry on the

synchronisation of the line start permanent magnet synchronous motor, Selected papers from the

International Conference on Computational Electromagnetics (CEM), IET Science, Measurement and

Technology, 2014, pp. 197-203

[2] Jedryczka C., Wojciechowski R.M., Demenko A., Finite element analysis of the asynchronous torque

in LSPMSM with non symmetrical squirrel cage winding, Int. J. Appl. Electromagn. Mech, vol. 49,

no. 2, 2014, pp. 367-373

[3] Mingardi D., Bianchi N., Line-Start PM-Assisted Synchronous Motor Design, Optimization, and Tests,

IEEE Transactions on Industrial Electronics, vol. 64, no. 12, 2017, pp. 9739-9747

[4] Ugale R.T., Chaudhari B.N., Rotor Configurations for Improved Starting and Synchronous

Performance of Line Start Permanent-Magnet Synchronous Motor, IEEE Transactions on Industrial

Electronics, vol. 64, no. 1, 2017, pp. 138-148

[5] Łyskawiński W., Jędryczka C., Szeląg W., Influence of magnet and cage shape on properties of the

line start synchronous motor with powder hybrid rotor, 978-1-5386-0359-8/17/$31.00 ©2017 European

Union, 2017, pp.155-163

[6] Barański M., Idziak P., Łyskawiński W., Analiza porównawcza stanów pracy silników indukcyjnego

i synchronicznego z magnesami trwałymi i klatką rozruchową, Electrical Engineering, vol. 77, 2014,

pp. 155-163

– 45 –

VERIFICATION OF BERTOTTI’S LOSS MODEL

FOR NON-STANDARD EXCITATION

B. Koprivica1 and K. Chwastek

2

1 University of Kragujevac, Faculty of Technical Sciences in Cacak

Svetog Save 65, 32000 Cacak, Serbia, e-mail: [email protected] 2 Czestochowa University of Technology, Faculty of Electrical Engineering

Armii Krajowej 17, 42-200 Czestochowa, Poland, e-mail: [email protected]

Abstract. The paper focuses on the possibility to use the Bertotti’s loss theory to describe energy losses in a cylindrical

core made of grain-oriented steel under non-standard excitation conditions.

I. INTRODUCTION

Bertotti’s theory [1,2] of loss dissipation in ferromagnetic materials still attracts the attention

of scientific community [3]. However the formulas developed by the author are generally valid for

sine induction waveform. In reality the excitation conditions may differ significantly from those

prescribed in international standards, therefore it is crucial to examine the possibility to use the

formalism for generic B-waveforms.

In the present paper we examine the possibility to use the Bertotti’s formulas for classical and

excess loss computation for triangular H-waveforms in a cylindrical core made of grain-oriented

steel.

REFERENCES

[1] Bertotti G., General properties of power losses in soft ferromagnetic materials, IEEE Trans. Magn.,

vol. 24, no. 1, 1988, pp. 621-630

[2] Bertotti G., Hysteresis in magnetism, Academic Press, San Diego, 1988

[3] Zhao H., Ragusa C., de la Barrière O., Wang Y., Fiorillo F., Energy losses in soft magnetic materials

under symmetric and asymmetric induction waveforms, IEEE Trans. Power Electron., 2018,

DOI: 10.1109/TPEL.2018.2837657



https://doi.org/10.1109/TPEL.2018.2837657

– 46 –

– 47 –

SHORT-CIRCUIT AND LOAD OPERATION OF SINGLE-PHASE

TRANSFORMER AT LOW FREQUENCIES

B. Koprivica1, K. Chwastek

2 and S.M. Koprivica

3

1 University of Kragujevac, Faculty of Technical Sciences in Cacak

Svetog Save 65, 32000 Cacak, Serbia, e-mail: [email protected] 2 Czestochowa University of Technology, Faculty of Electrical Engineering

Armii Krajowej 17, 42-200 Czestochowa, Poland, e-mail: [email protected] 3 University of Kragujevac, Faculty of Technical Sciences in Cacak

Svetog Save 65, 32000 Cacak, Serbia, e-mail: [email protected]

Abstract. The aim of this paper is to present experimental results on the short circuit and load test of the single-phase

transformer at low frequencies. The paper gives information on PC based measurement setup and presents the results

of measurement of primary and secondary voltages and currents and electric power of the transformer. Also, a proper

discussion of the results obtained is given in the paper.

I. INTRODUCTION

Previous research on the operation of a single-phase power transformer at low frequencies was

conducted under no-load conditions [1]. It was found that primary current of the transformer

increases with the decrease of the frequency (at constant amplitude of primary voltage). Also, this

current becomes significantly distorted, which indicates increase of the magnetic flux density in the

transformer core and approaching to the magnetic saturation [2].

The aim of this paper is to present results of continuation of research on the operation of the

transformer at low frequencies in the case of short-circuit or load conditions. General requirements

on such testing are given in the international standard [3]. Short overview of these tests exists in the

literature [4, 5].

The results of the short-circuit and load test of power transformer at low frequencies are

presents in this paper. The transformer under study is a single-phase unit, rated at 1 kVA,

230 V/12 V, 50 Hz. It has an EI core. Tests are performed under the sinusoidal primary voltage at

different frequencies from 5 Hz to 50 Hz. The primary voltage amplitude has been maintained at

55 V. PC based measurement setup is used to record time waveforms of the primary voltage, the

secondary voltage, the primary current and the secondary current. The input power is also recorded

during the tests. The paper presents results obtained and gives their discussion.

II. EXPERIMENTAL SETUP AND MEASUREMENTS

EI shaped magnetic core of tested transformer is made of M530-50A non-oriented electrical

steel sheets.

Scheme of electrical connections for PC based measurement is presented in Fig. 1. Power

supply generates time-varying voltage of sinusoidal shape with adjustable frequency. This voltage

is supplied to the primary side of transformer over the non-inductive resistor R. Primary and

secondary currents i1 and i2 are calculated using measured voltages uR1 and uR2. Primary and

secondary voltages u1 and u2 are measured at the ends of primary and secondary winding.

LabVIEW application is made and used in the measurement of these voltages. Its appearance of

during the measurements under load conditions at a frequency of 5 Hz is presented in Fig. 2.

According to this figure, it can be seen that the primary current of the transformer is highly

distorted from sinusoidal shape, while the secondary current and voltages have very low distortion.

This is caused by the increase of the magnetising current of the transformer core.




– 48 –

Fig.1. Measurement setup based on personal computer (SC - short-circuit).

Fig.2. LabVIEW application during experiment at 5 Hz.

REFERENCES

[1] Koprivica B., Milunovic Koprivica S., No-Load Operation of Single-Phase Power Transformer at Low

Frequencies, International Scientific Conference - UNITECH 2017 - Vol. 1, Gabrovo, Bulgaria, 2017,

pp. I-75 - I-79

[2] Langella R., Testa A., Emanuel A.E., On the Effects of Subsynchronous Interharmonic Voltages on

Power Transformers: Single Phase Units, IEEE Transactions on Power Delivery, vol. 23, no. 4, 2008,

pp. 2480-2487

[3] IEC 60076-1 Edition 3.0, 2011-04, Power transformers - Part 1: General, IEC, Geneva, Switzerland,

2011

[4] Winders Jr. J.J., Power Transformers - Principles and Applications, Marcel Dekker, NY, USA, 2002

[5] Harlow J.H., Electric Power Transformer Engineering, CRC Press, Boca Raton, FL, USA, 2004

1i 2i

NI cDAQ-9172

1Ru

2u1R

PC

1uPower

Supply

TR

2RLoadSC

2Ru

– 49 –

STRUCTURE AND MAGNETIC PROPERTIES OF THE RAPIDLY

SOLIDIFIED Gd3Zr10Fe55Co10Mo5W2B15 ALLOY

K. Kotynia1, A. Chrobak

2 and P. Pawlik

3

1 Institute of Physics, Częstochowa University of Technology, Av. Armii Krajowej 19, 42-200 Czestochowa, Poland,

e-mail: [email protected] 2 Institute of Materials Science, University of Silesia, 75 Pułku Piechoty 1, 41-500 Chorzów, Poland,

e-mail: [email protected] 3 Institute of Physics, Częstochowa University of Technology, Av. Armii Krajowej 19, 42-200 Czestochowa, Poland,


Abstract. The paper presents a study of structural and magnetic properties of the Gd3Zr10Fe55Co10Mo5W2B15 glassy

alloy. The solidified ribbon of the Gd3Zr10Fe55Co10Mo5W2B15 alloyof the amorphous structure confirmed by X-ray

diffraction (XRD) was studied. The Curie temperature TC was determined from the curve representing the dependence

of magnetic polarization J on temperature T. The isothermal magnetic entropy change |-ΔSM| and relative cooling

power RCP were calculated from the experimental magnetization curves to assess the possibility of potential

application.

I. INTRODUCTION

The growing interest in magnetocaloric materials was caused by using them in cooling systems

operating near room temperature. Moreover, in recent years, magnetocaloric effect was also used in

power industry to convert industrial waste heat to electricity [1]. There are three physical quantities

that designate the usefulness of magnetocaloric materials: the magnetic entropy change |-ΔSM|, the

adiabatic temperature change |-ΔTad| and the relative cooling power (RCP) [2]. An ideal magnetic

refrigerant material should exhibit large values of both |-ΔSM| and |-ΔTad|, as well as high RCP

around room temperature at the low magnetic field [3].

High magnetic entropy change was found in Gd5(Si2Ge2) [4], MnAs [5], MnFe(P, As) alloys

[6]. The first order magneto-structural phase transition and the appearance of hysteretic losses were

observed for this alloys. Although these materials reveal significant |-ΔSM| and |-ΔTad| their major

drawbacks are low RCP and complex processing route. Despite relatively low |-ΔSM| and |-ΔTad|

values due to second order magnetic phase transition the amorphous alloys seems to be interesting

alternative for their large RCP and low processing costs.

The Gd3Zr10Fe55Co10Mo5W2B15 alloy seems to be a good candidate for application

as a refrigerant. This alloy exhibits a broader |-ΔSM| peak over a wide range of temperature. It

shows no thermal hysteresis and its electrical resistivity is larger than for crystalline materials [7].

II. TECHNICAL INSTRUCTIONS

The Gd3Zr10Fe55Co10Mo5W2B15 glassy alloy was obtained by arc-melting of the mixture of

high purity (99.98 %) constituent elements Gd, Zr, Fe, Co, Mo, W with the addition of pre-alloyed

Fe-B. To protect oxidation of the alloy, titanium was used as a getter. The ingot was re-melted

seven times to guarantee the homogeneity of the alloy.

The ribbon was prepared by melt-spinning technique at surface velocity of the copper roll of

32 m/s. The phase structure was investigated by X-ray diffractometry (XRD) using Bruker D8

Advance diffractometer with CuK radiation and the LynxEye semiconductor detector.

The data were recorded using the step-scanning method in 2Ɵ range from 30 to 100 degrees.

The J(T) curve was obtained using the Faraday balance operating at a magnetic field

of 0.87 T and temperature range from 300 K to 750 K with heating rate of 10 K/min.

The field dependences of magnetization were recorded in field cooled mode (FC) by SQUID

MPMS XL-7 Quantum Design. The magnetic measurements M(H) were performed in the




– 50 –

temperature range 150 K to 350 K and Arrot plots were constructed at constant temperature values

from these curves. Magnetocaloric effect MCE was estimated by calculation of the temperature

dependences of magnetic entropy change |-ΔSM| for various changes of external magnetic fields

according to the Maxwell thermodynamic formula [8] :

∆SM (T, H) = ∂M(T,H)

∂T

HdH

H

0 (1)

where: T – temperature, M(T, H) – magnetization, H – external magnetic field.

To assess the applicability of the Gd3Zr10Fe55Co10Mo5W2B15 alloy as a refrigerant, the relative

cooling powers were calculated according to the formula [9]:

RCP = | − ∆SMmax | ∙ δTFWHM (2)

for various changes of external magnetic field, where | − ∆𝑆𝑀𝑚𝑎𝑥 | denotes maximum entropy

change and 𝛿𝑇𝐹𝑊𝐻𝑀 – the full width at half maximum of |-ΔSM| function versus temperature.

REFERENCES

[1] Vuarnoz D., Kitanovski A., Gonin C., Borgeaud Y., Delessert M., Meinen P.W., Egolf M.,

Quantitative feasibility study of magnetocaloric energy conversion utilizing industrial waste heat,

Applied Energy, vol. 100, C, 2012, pp. 229-237

[2] Chaudhary V., Repaka D.V.D., Chaturvedi A., Sridhar I., Ramanujan R.V., Magnetocaloric properties

and critical behavior of high relative cooling power FeNiB nanoparticles, Journal of Applied Physics,

vol. 116, 2014, pp. 163918

[3] Shamba P., Zeng R., Wang J.L., Campbell S.J., Dou S.X., Enhancement of the refrigerant capacity in

low level boron doped La0.8Gd0.2Fe11.4Si1.6, Journal of Magnetism and Magnetic Materials, vol. 331,

2013, pp. 102-108

[4] Pecharsky A.O., Gschneidner K.A., Pecharsky V.K., The giant magnetocaloric effect of optimally

prepared Gd5Si2Ge2, Journal of Applied Physics, vol. 93, 2003, pp. 4722-4728

[5] Wada H., Asano T., Effect of heat treatment on giant magnetocaloric properties of Mn1+δAs1−xSbx,

Journal of Magnetism and Magnetic Materials, vol. 290, 2005, pp. 703-705

[6] Yibole H., Guillou F., Zhang L., Dijk N. H. van, Brück E., Direct measurement of the magnetocaloric

effect in MnFe(P,X)(X = As, Ge, Si) materials, Journal of Physics D: Applied Physics, vol. 47, no. 7,

2014, pp. 1-9

[7] Kotynia K., Pawlik P., Hasiak M., Pruba M., Pawlik K., Structural and Magnetic Studies of the Fe–Co–

Zr–Mo–W–B Amorphous Alloy, Acta Physica Polonica A, vol. 131, 2017, pp. 1204-1206

[8] Gschneidner K. A., Pecharsky Jr. and V. K., Magnetocaloric Materials, Annual Review of Materials

Science, vol. 30: 387-429, 2000, pp. 387-429

[9] Zhang Q., Thota S., Guillou F., Padhan P., Hardy V., Wahl A., Prellier W., Magnetocaloric effect and

improved relative cooling power in (La0.7Sr0.3MnO3/SrRuO3) superlattices, Journal of Physics:

Condensed Matter, vol. 23, 2011, pp. 052201

https://econpapers.repec.org/article/eeeappene/

http://scitation.aip.org/content/contributor/AU1101434;jsessionid=FQKZagYKSEqJ4aCSdd6kY9YO.x-aip-live-03




http://aip.scitation.org/journal/jap

http://www.sciencedirect.com/science/article/pii/S030488531200933X





https://www.sciencedirect.com/journal/journal-of-magnetism-and-magnetic-materials

http://www.sciencedirect.com/science/journal/03048853/331/supp/C

http://iopscience.iop.org/journal/0022-3727

http://iopscience.iop.org/volume/0022-3727/47

http://iopscience.iop.org/issue/0022-3727/47/7




– 51 –

MAGNETIC RELAXATION IN IRON BASED MELT SPUN RIBBONS

P. Kwapuliński and G. Haneczok

University of Silesia, Institute of Materials Science, 40-500 Chorzów, 75 PułkuPiechoty 1A, Poland


The paper concerns detail examinations of thermal/time instabilities of macroscopic properties

of iron based amorphous alloysby making use of magnetic relaxation technique. Experiments were

carried out for Fe74Cu1 Cr3Si13B9 melt spun ribbons with thickness andwidth of about 20 mm and

5mm, respectively. In order to study the structural relaxation in the context of free volume

diffusion samples in the as quenched state were annealed at temperatures ranging from 300 K to

650 K for one hour. Such annealing slightly changes the amorphous microstructure and what

follows changes the degree of advancement of structural relaxation. Measurements of magnetic

reluctivity at low field (0.1 A/m) versus time at room temperature were carried out for samples

after demagnetization by applying precision RLC meter – Agilent E4980A. The obtained curves

r(t) show that the reluctivity increases with time reaching at least a partial saturation at times over

80 ks (over 22 h). Numerical analysis allows concluding that the observed effect consists of two

components attributed to the reversible and irreversible component of the structural relaxation. The

reversible component was described by the so-called coupling model referring to diffusion in

correlated systems. The theoretical basis of this model is also presented and discussed in detail.

The main conclusions of the present paper can be summarized as follows: i) reversible

component of magnetic relaxation in iron based amorphous alloys (e.g. Fe74Cu1Cr3Si13B9) can be

well described by the coupling model which allows monitoring initial stages of structural

relaxation, ii) relaxation time of the irreversible component of magnetic relaxation is at least three

orders of magnitude longer that the relaxation time of the reversible component which means that

for long times this component can be approximated by a straight line, iii) the observed

disappearance of thermal/time instabilities of magnetic properties caused by the preliminary

annealing is quantitatively documented.


– 52 –

– 53 –

EVALUATION OF THE INTERDEPENDENCY OF MECHANICAL

CUTTING AND MAGNETIC ANISOTROPY ON THE MAGNETIC

PROPERTIES OF NON-ORIENTED FESI ELECTRICAL STEEL

N. Leuning, S. Steentjes and K. Hameyer

Institute of Electrical Machines (IEM), RWTH Aachen University, D-52062 Aachen, Germany


Abstract. Due to present-day challenges for the improvement of the operating characteristics of rotating electrical

machines, the core material is regarded as increasingly important. Non-oriented (NO) electrical steel sheets are often

used for the construction of magnetic cores. The most appropriate choice of material and machine design is based on

standardized material data obtained from Epstein frames or single sheet testers. Both, low losses as well as good

magnetizability in any spatial direction of the sheet plane are requested. Despite their name, NO electrical steels do

show a magnetic anisotropy. A significant increase of loss and decrease of permeability can additionally be induced by

mechanical cutting processes to shape the magnetic circuit. Magnetic anisotropy and induced mechanical stress can

affect the local iron loss distribution, magnetizability, acoustic behavior and therefore needs to be considered in

numerical simulations of electrical machines. In this paper, the interdependency of the magnetic anisotropy with the

effect of cutting is studied in order to improve the necessary understanding of the material behavior.

I. INTRODUCTION

Machine modeling is the standard tool to design electrical machines. Accurate modeling

enables the possibility to achieve high power densities and low losses without unnecessary

oversizing, at still compact geometries. The potential for improvement by constructive measures is

largely exploited and subsequently, the focus for future progress is laid on material design and

optimization. To maximize the energetic efficiency, iron losses need to be minimized. The

mechanical processing of electrical steel components is often disregarded in machine design, by

using data from material testing, that is standardized for single sheet tester (SST) or Epstein frame

(EPF) measurements with fixed geometries and required gentle processing of samples according to

international standards, e.g. IEC60404-3. However, the geometry and processing majorly affect the

properties of the NO in its application and are ideally incorporated already during the design stage

[1]. Cutting is highly detrimental to the electromagnetic properties of the steel sheets, e.g., losses

and magnetizability [2][3][4]. Different cutting techniques as well as the cutting parameters

influence the properties to a different extend, dependent on the magnitude of the physical impact,

primarily the induced mechanical stress [5]. The insufficient knowledge regarding the

quantification of material deterioration and the cutting impact motivates consecutive studies on this

topic.

The magnetic anisotropy is a further subject, which is mostly neglected by standardized

material measurements. Properties are either determined in rolling (RD) or transverse direction

(TD) separately or joint, in one measurement with Epstein frames of both orientations. The angular

Fig.1. Influence of cutting and magnetic texture of a conventional 2.9 wt.-% FeSi at 50 Hz.

-2.0

-1.0

0.0

1.0

2.0

-2000 0 2000Mag

netic

p

ola

riza

tio

n

J 1

.5 T

, 50

Hz

in T

Magnetic field strength H in A/m

a) Influence of cutting

Water jet, uncut Water jet, cutGuillotine, uncut Guillotine, cut

0.0

0.5

1.0

1.5

2.0

1 100 10000

Mag

netic

p

ola

riza

tio

n

J max

, 50

Hz

in T

Magnetic field strength Hmax in A/m

b) Effect of magnetic anisotropy

0° 30° 45° 60° 90°

RD 0 TD (90 )


– 54 –

dependence of properties in between those directions is however not linear and therefore needs to

be studied as well [6]. Dependent on the NO grade, the extend of the magnetic texture can be

different [7].

II. APPROACH AND RESULTS

In this paper, the interdependency of the effect of cutting and the magnetic anisotropy is

studied, in order to gain knowledge on the phenomenology of both effects. Both effects are

separately displayed in Fig. 1 a) and b). A conventional 2.9 wt.-% FeSi with 0.5-mm thickness is

characterized by using a 120 mm x 120 mm SST. Samples are cut to different strip widths in

different angles relative to RD. Two different cutting techniques are studied, which are water jet

and guillotine. The measurements are performed between 0.1 T and 1.8 T at different frequencies

from 20 Hz to 1000 Hz. The deterioration of magnetic properties due to the cutting can thereby be

evaluated in different spatial directions of the sheet plane and the sensitivity of certain directions

can be quantified and discussed, as displayed in Fig. 2, and evaluated regarding their frequency and

induction dependence, their initial material properties, i.e. crystallographic texture and regarding

the impact of the cutting procedure, i.e., micro hardness measurements.

REFERENCES

[1] Martin F., Aydin U., Sundaria R., Rasilo P., Belahcen A., Arkkio A., Effect of Punching the Electrical

Sheets on Optimal Design of a Permanent Magnet Synchronous Motor, IEEE Trans. Mag., vol. 54,

no. 3, 2018, pp. 1-4

[2] Emura M., Landgraf F.J.G., Ross W., Barreta J.R., The influence of cutting technique on the magnetic

properties of electrical steels, JMMM, vol. 254-255, 2003, pp. 358-360

[3] Moses A.J., Derebasi N., Loisos G., Schoppa A., Aspects of the cut-edge effect stress on the power loss

and flux density distribution in electrical steel sheets, JMMM, vol. 215-216, 2000, pp. 690-692

[4] Schoppa A., Schneider J., Roth J.-O., Influence of the cutting process on the magnetic properties of

non-oriented electrical steels, JMMM, vol. 215-216, 2000, pp. 100-102.

[5] Weiss H.A., Leuning N., Steentjes S., Hameyer K., Andorfer T., Jenner S., Volk W., Influence of shear

cutting parameters on the electromagnetic properties of non-oriented electrical steel sheets, JMMM,

vol. 421, 2017, pp. 250-259.

[6] Emura M., de Campos M.F., Landgraf F.J.G., Teixeira J.C., Angular dependence of magnetic

properties of 2% silicon electrical steel, JMMM, vol. 226-230, 2001, pp. 1524-1526

[7] Leuning N., Steentjes S., Hameyer K., On the Homogeneity and Isotropy of Non-Grain-Oriented

Electrical Steel Sheets for the Modeling of Basic Magnetic Properties from Microstructure and Texture,

IEEE Trans. Mag., vol. 53, no. 11, 2017, pp. 1-5

Fig.2. Relative loss increase at 1.5 T and 50 Hz, compared to the uncut sample in different orientations relative to

the rolling direction for a) guillotine and b) water jet cut samples

0%

50%

100%

150%

120 m

m10 m

m7.5

mm

5 m

m

120 m

m10 m

m7.5

mm

5 m

m

120 m

m10 m

m7.5

mm

5 m

m

120 m

m10 m

m7.5

mm

5 m

m

0° 30° 60° 90°

Rel

ativ

e lo

ss i

ncre

ase

ΔP

s

due

to c

uttin

g

Angle θ relative to RD and strip width dS

a) 1.5 T, 50 Hz, guillotine

0%

50%

100%

150%

120 m

m10 m

m7.5

mm

5 m

m

120 m

m10 m

m7.5

mm

5 m

m

120 m

m10 m

m7.5

mm

5 m

m

120 m

m10 m

m7.5

mm

5 m

m

0° 30° 60° 90°

Rel

ativ

e lo

ss i

ncre

ase

ΔP

s

due

to c

uttin

g

Angle θ relative to RD and strip width dS

b) 1.5 T, 50 Hz, water jet

– 55 –

MINIATURE CURRENT SENSOR FOR MEDIUM VOLTAGE NETWORKS

A. Lisowiec, A. Nowakowski, G. Kowalski and P. Wlazło

Tele and Radio Research Institute, 11 Ratuszowa, 03-450 Warsaw, Poland

e-mail: [email protected], [email protected], [email protected],

[email protected]

Abstract. The paper presents the construction and electrical parameters of miniature current sensor designed for use

in signal processing paths of protection devices working with current transformers in power substations. The sensor is

in a form of double solenoid. The aim of designing such a sensor was to replace the current transformer with magnetic

core and by that to achieve better electrical parameters such as dynamic range, bandwidth and linearity.

I. INTRODUCTION

Current measurement in power networks is increasingly done with Rogowski coils. These coils

measure directly primary circuit current and their advantages are well known. However, there is a

lot of legacy measurement and protection equipment still in use that measure current with the use

of current transformers connected to primary circuits. Refurbishing the older substations must be

done with protection relays that are matched at the input of their current measuring circuits to

current transformers output.

II. CONSTRUCTION OF THE SENSOR

The miniature current sensor described in the paper has been developed especially for signal

conditioning circuits of measurement and protection devices working with current transformers.

Modern power protection equipment is required to have wide measurement dynamic range, wide

bandwidth and good accuracy. Accordingly, the requirements put on the sensor were wide dynamic

range, good linearity, wide bandwidth and small size. The simplest design that fulfilled all of these

requirements appeared to be an air-core transformer, cylindrical in shape, where the secondary

winding is placed inside the primary winding, fig. 1.

Fig.1. Construction of the miniature current sensor

The current flowing through the primary circuit induces a voltage in the secondary circuit. By

the Faraday law, the voltage leads in phase the current by 90 degrees. The design goal was to

achieve as high voltage output as possible within space constraints that in this case were 22 mm x





– 56 –

12 mm (length x diameter). Other design goals were good repeatability of electrical parameters.

Table 1 shows the sensitivity of 10 sensors chosen randomly from prototype production lot.

Table 1. Sensitivity spread of miniature current sensors

Sensor Current in primary circuit [A] Output voltage [mV] Sensitivity Dispersion

1 5 29,12 5,8240 -0,212%

2 5 29,00 5,8000 -0,624%

3 5 29,53 5,9060 1,193%

4 5 29,20 5,8400 0,062%

5 5 29,21 5,8420 0,096%

6 5 29,34 5,8680 0,541%

7 5 29,00 5,8000 -0,624%

8 5 29,04 5,8080 -0,487%

9 5 28,99 5,7980 -0,658%

10 5 29,54 5,9080 1,227%

Sensitivity averaged 5,8364

The spread of the sensitivity is within 1.3% and is low enough to be easily calibrated out in the

protection device. The signal output of the sensor is rather low, and equals approximately 30 mV

for 5 A current in the primary winding of the sensor. For such low output voltage the noise and

interference in the signal amplification path have to be considered. The noise sources are the

thermal noise of the secondary winding resistance of the sensor (equal to 10 Ω) and the noise of the

operational amplifier. Low noise operational amplifiers are available with input voltage noise equal

to 0,7 µV/Hz. The bigger problem is the influence of the magnetic fields generated by currents in

other wires than the measured one.

The magnetic field created by the outer coil of the sensor attains its maximum inside the coil

but the magnetic field lines close outside the coil. As the sensors have to be placed inside the

protection relay, close to each other, their mutual interaction has to be determined. According to

theoretical calculations, carried out similarly as in [1] (presented in full paper), and experimental

data, the best placement of the sensors within predefined area is as shown in figure 2.

Fig.2. Placement of the sensors that minimizes their mutual influence

REFERENCES

[1] Lisowiec A., Nowakowski A., Kowalski G. Influence of primary conductor position on Rogowski coil

measurement accuracy, Proceedings of ISEF 2017, 13-16 September 2017, Łódź

– 57 –

THE IMPACT OF APPLICATOR SIZE ON DISTRIBUTION OF

ELECTROMAGNETIC FIELD USED IN MAGNETOTHERAPY

E. Łada-Tondyra

Wydział Elektryczny, Politechnika Częstochowska, Al. Armii Krajowej 17, 42-200 Częstochowa,


Abstract. Electromagnetic field is used in magnetotherapy. The therapy effectiveness depends on the value of magnetic

induction and its distribution in applicators. In the case of real objects, uniformities inside the applicator are higher

than calculated in the model. The modeled distribution of induction inside the applicators are similar to characteristics

obtained from the measurements.

I. INTRODUCTION

The electromagnetic field is used both in ad hoc operations and in long-term therapy or

rehabilitation [1,2]. Magnetotherapy is the most popular method used in bone diseases [3]. The

magnetotherapy devices consist of a control device and applicators. In Poland, the most commonly

used are coils with diameters from approx. 15 cm to 70 cm and a length of approx. 20-30 cm. The

electromagnetic field used in magnetotherapy generated by the solenoid has

a frequency of 10 to 100 Hz and magnetic induction from 0.1 mT to 20 mT. The effectiveness of

therapy using the electromagnetic field depends primarily on the value of magnetic induction and

its distribution inside the applicator [4,5].

An important problem is the appropriate choice of an applicator. It is practiced that the choice

of applicator size is determined by the size of a body part being treated. Applicators used in

physiotherapy are connected to the same control device. Practice shows that the intensity level of

the treatment is not adjusted to the size of the applicator.

II. NUMERICAL ANALYSIS

The knowledge of the electromagnetic field distribution of applicators is required to plan the

magnetotherapy and to evaluate its effectiveness, because the magnetic component generates eddy

currents in the human body [6]. Providing exact geometry influences into greater accuracy of field

distribution. The distribution of the magnetic induction module (Figure 1) allows the observation of

relatively small changes in the induction value inside the applicator. The maximum value of

induction is measured at the edge of the solenoid in its half-length.

For all applicators, the same excitation conditions were applied that allowed observing the

differences in the maximum value of induction for analyzed sizes of solenoid. The value of

induction module inside the analyzed applicators does not change significantly. However, the

difference in value at the edge and inside the solenoid increases with the increase in the radius of

the solenoid.

III. VERIFICATION OF THE MODEL

In order to verify the numerical model, the measurements of induction distribution were

carried out around solenoid applicators used in magnetotherapy. In the tests, the MAGNETRONIC

MF-10 power supply (usually used in magnetotherapy) and solenoid applicators with different

diameters were applied. The measurements of magnetic induction were made using the CK-1

Halleter teslameter.

Magnetic field induction were measured for applicators with a radius 0.095 m, 0.15 m and

0.245 m. For all applicators, the same parameters such as waveform shape (sinusoidal), frequency


– 58 –

(50 Hz) and intensity (maximum according to the manufacturer's scale) were used. Similar to the

numerical models, the induction value at the edge and inside the solenoid increases with the

increase in the radius of the solenoid.

Fig.1. The distribution of induction module inside the applicators a) 0.095 m, b) 0,15 m, c) 0,245 m

In the case of real objects, the heterogeneity of the induction distribution inside the applicators

is greater than the case of modeled one. The differences between the modeled and measured values

are from 1% up to 5%, depending on the solenoid diameter.

IV. CONCLUSIONS

The results presented in the paper shows that the applicators size determines the magnetic

induction distribution inside the solenoid, generated for the same excitation conditions. It indicates

that not only the position of the treated body part, but also the applicator size have

a great impact on the value of induction used in magnetotherapy.

REFERENCES

[1] Sieroń A., Zastosowanie pól magnetycznych w medycynie, αmedica Press, Bielsko-Biała, 2002 [2] Gas P., Transient Temperature Distribution inside Human Brain during Interstitial Microwave

Hyperthermia, Przeglad Elektrotechniczny, vol. 89, no. 3a, 2013, pp. 274-276

[3] Krawczyk A., Miaskowski A., Łada-Tondyra E. Ishihara Y., Healing of orthopaedic diseases by means

of electromagnetic field, Przegląd Elektrotechniczny, vol. 86, no. 12, 2010, pp. 72-75 [4] Cieśla A., Syrek P., Parameters and position of the applicator’s effect on magnetic field distribution

during magnetotherapy, Przegląd Elektrotechniczny, vol. 88, no. 12b, 2012, pp. 124-127 [5] Cieśla A., Kraszewski W., Skowron M., Syrek P., Analiza rozkładu pola magnetycznego

generowanego przez urządzenia do fizykoterapii. Przegląd Elektrotechniczny, vol. 91, no. 2, 2015,

pp. 162-165 [6] Cieśla A., Kraszewski W., Tadeusiewicz R., Visualization of field generated by portable coil designed

for magnetotherapy, Przegląd Elektrotechniczny, vol. 88, no. 10a, 2012, pp. 127-131

– 59 –

INVERSE MODEL OF THE MAGNETIC HYSTERESIS BASED

ON AN EXPONENTIAL FUNCTION

W. Mazgaj, Z. Szular and M. Sierzega

Cracow University of Technology, e-mail: [email protected], [email protected], [email protected]

Abstract. In many cases it is profitable to apply an inverse model of the magnetic hysteresis presenting the relationship

between the field strength and the flux density. The paper presents a relatively simple method of approximation of field

strength changes during magnetization of electrical steel sheets. It was assumed that the field strength changes are

a sum or a difference of the function which describes one curve of the limiting hysteresis loop and a certain

“transient” component.

I. INTRODUCTION

The mechanics of the magnetic hysteresis phenomenon is quite well known [1,2]. However,

a formulation of suitable mathematical relations on the basis of the physics of this phenomenon is

still a relatively difficult problem, despite the fact that scientific literature contains a lot of papers

presenting different mathematical models of the hysteresis phenomenon. In some numerical

calculations it is profitable to use an inverse model of the magnetic hysteresis, which presents

changes of the field strength as a function of the flux density, especially when the flux densities are

treated as unknown quantities in numerical calculations. The most well-known models of the

hysteresis are the Preisach model and the Jiles-Atherton model [3, 4]; in practice, only the inverse

Jiles-Atherton model was formulated [5]. The advantage of this model is short calculation time.

However, the calculation algorithm is quite complicated and the determination of model parameters

is difficult.

II. APPROXIMATION OF THE FIELD STRENGTH CHANGES

Any point P with the co-ordinates (H, B) can move along a certain trajectory to one of the

limiting magnetization curves B=f(H), depending on the field strength changes (Fig. 1a). Similarly,

considering an inverse model of the magnetic hysteresis, point P with the co-ordinates (B, H) can

move to one of the limiting magnetization curves H=f(B), depending on changes of the flux density

(Fig. 1b).

Fig.1. Hysteresis loops: a) as function B=f(H), b) as function H=f(B)

When the flux density B increases then changes of the field strength Hr(B) can be written in the

following form:




– 60 –

)](exp[])([)()(0000

BBkHBHBHBHBruur

(1)

where: Hu(B) – upper curve of the limiting hysteresis loop (as a function H=f(B)), H0 – initial value

of the field strength, B0 – initial value of the flux density, kBr – attenuation coefficient of the

“transient” component when B increases.

For decreasing values of the flux density the relationship describing changes of the field

strength Hd(B) can be written as follows:

)](exp[)]( [)()(0000

BBkBHHBHBHBdbbd

(2)

where: Hb(B) – lower curve of the limiting hysteresis loop (as a function H=f(B)), kBd –attenuation

coefficient of the “transient” component when B decreases.

The calculations with the use of the inverse model of the magnetic hysteresis were made for

different electrical steel sheets. For example, Figure 2a shows the comparison between measured

and calculated hysteresis loops of the dynamo sheet M530-50A, and Figure 2b presents a similar

comparison regarding the loops of the transformer sheet M120-27S.

Fig.2. Comparison of the measured and calculated hysteresis loops: a) dynamo sheet M530-50A,

b) transformer sheet M120-27S; measured loops – black lines, calculated loops as B=f(H) – blue lines,

calculated loops as H=f(B) – red lines

III. CONCLUSIONS

The field strength changes as the dependence of the flux density are written by means of

simple formulas. Therefore, they can be relatively easily inserted into equations of the magnetic

field distribution. The times of numerical calculations are shorter than in other models. In order to

apply this method the limiting hysteresis loop of the given electrical steel sheet has to be known.

Additionally, some minor loops should be measured to choose the attenuation coefficients of the

“transient” components correctly.

REFERENCES

[1] Bertotti G., Mayergoyz I.D., The science of hysteresis, vol. I, Elsevier, Oxford, 2006

[2] Tumański S., Handbook of magnetic measurements, CRC/Taylor & Francis, Boca Raton, 2011

[3] Iványi A., Hysteresis models in electromagnetic computation, Akadémiai Kiadó, Budapest, 1997

[4] Liorzou F., Phelps B., Atherton D.L., Macroscopic models of magnetization, IEEE Trans. on

Magnetics, vol. 36, no. 2, 2000, pp. 418-427

[5] Sadowski N., Batistela N.J., Bastos J.P.A., Lajoie-Mazenc M., An inverse Jiles-Atherton model to take

into account hysteresis in time-stepping finite-element calculations, IEEE Trans. on Magnetics, vol. 38,

no. 2, 2002, pp. 797-800

– 61 –

MAGNETIC MEASUREMENT OF FERRITE CONTENT OF ALLOYS

I. Mészáros and B. Bögre

Budapest University of Technology and Economics, Department of Materials Science and Engineering,

H-1111 Bertalan L. u. 7., Budapest, Hungary, e-mail: [email protected]

Abstract.In this paper three different magnetic measurement methods were compared. The tested measuring techniques

were AC magnetometer, DC magnetometer and a so called Ferritscope device. They were used to determine the ferrite

content of alloys. For this investigation a model sample series was prepared from 2507 type super-duplex stainless

steel by cold rolling and heat treatment. The above-mentioned methods were used to determine the δ-ferrite content of

the samples. The results of the different electromagnetic methods were compared with each other. The limits,

disadvantages and advantages of the applied methods were analyzed.

I. INTRODUCTION

Nowadays, the importance of nondestructive magnetic measurements increases rapidly. The

aim of the fast and widely useable NDT can be defect (cracks, voids etc.) detection or study of

material properties without damaging the sample.The magnetic- and electromagnetic

measurements are especially useful for determining the structural changes of alloys caused by

technological- or material deterioration processes due to service.Several NDT methods are used in

industrial practice from which those electromagnetic methods are investigated in this paper which

are suitable to determine ferrite content. Alternating current (AC) magnetometer, direct current

(DC) magnetometer and Ferritscope were applied to measure the δ-ferrite content of cold rolled

and heat treated super-duplex stainless steel (SDSS) samples.

SDSS is a particular category of stainless steels characterized by a double-phase

microstructure with about equal proportions of austenite and ferrite phases. The combination of

properties, including high strength and excellent resistance to corrosion and stress corrosion

cracking in chloride ion containing environments make SDSS very attractive for many

applications.Unfortunately, there are several disadvantages as well.The most important phase

transformation process in duplex stainless steel is the eutectic decomposition of ferrite which

means the transformation of the δ-ferrite into sigma phase and secondary austenite due to heat

treatment (𝛿 → 𝜎 + 𝛾2) [1].If the well-adjusted ferrite-austenite phase ratio changes due to heat

input these benefic properties can disappear. Some percentage decrease of the ferrite content can

significantly decrease the corrosion resistance and impact energy of SDSS.Therefore, the

determination of ferrite content is essential in heat treated or welded duplex stainless steel

structures.

The aim of this study was to compare the capabilities of three different electromagnetic

methods which are suitable for ferrite content determination.

II. TESTED SAMPLES

For studying the capabilities of the before mentioned electromagnetic methods the 2507 grade

SDSS was chosen as a model material. This SDSS contains about 25% chromium and 7% nickel as

main alloying elements. From the original sheet material 35 uniform samples were cut with the size

of 15x10x100 mm. Samples were cold rolled at room temperature with six different reduction rates

(0, 10, 20, 30, 40, 50, 60%). The rolled samples were heat treated at 700°C, 750°C, 800°C, 850°C

temperatures for 30 minutes and cooled in normal air. At the end of the preparation process all

samples were milled for the same geometry (3,4x10x100 mm) which was suitable for the applied

AC magnetometer and Ferritscope devices. The applied DC magnetometer requires bulk


– 62 –

specimens, so the milled samples were cut into more pieces and fixed into a rectangular cuboid

(10x10x10,2 mm).

III. APPLIED MAGNETIC MEASUREMENTS

As it is well known the magnetic saturation polarization is directly proportional to the

ferromagnetic phase ratio (ferrite content in our case) of the specimen [2].

The AC magnetometeris suitable to measure the hysteresis and normal magnetization curves of

the specimen from which among others the maximal polarization, remnant induction, coercive field

and initial permeability can be determined.200 minor hysteresis loops were measured in case of

each specimens, the normal magnetization curves were determined from the peak points of the

minor hysteresis loops. The maximal excitation field of the AC magnetometer was about 128 A/cm

which definitely was not enough to saturate the samples. The saturation polarization values were

calculated by an extrapolation method based on the multiphase hyperbolic model [3].

The so called Stablein-Steinitz DC magnetometer is a magnetic bridgewhich has two

symmetrical yokes and a small cross-section cross bridge[4]. The maximum excitation field

strength was about 2,700 A/cm. Therefore, this setup is capable to excite the bulk steel samples

into magnetic saturation which makes it one of the most precise way of the ferrite content

measurement.Unfortunately, this set-up is not portable it is only for laboratory use because of its

extensive size.

Samples were also measured by a commercial Fischer FERITSCOPE FMP30 type Ferritscope

equipment[5]. The equipment contains a data acquisition device, a probe and an etalon series. This

user friendly, portable measuring device especially useful for quick determination of ferrite

content. Because of its physical limitations its excitation level is very low. The Ferritscope derives

the ferrite content from the initial permeability of the sample.

IV. RESULTS

The ferrite phase ratio values determined by AC and DC magnetometerswere close toeach

other in case of all deformation extents and heat treatments. In contrast, to the Ferritscope device

which gave significantly lower ferrite contents especially in case of plastic deformed samples. The

stronger the cold rolling reduction was the lower the measured ferrite content was.

This phenomenon was explained by the change of the shapes of magnetization curves. It

allowed us to develop a hysteresis model-based calculation method for eliminating this

measurement error of the Ferritscope device. The details of this correction method will be

presented.

REFERENCES

[1] Breda M., Brunelli K., Grazzi F., Scherillo A., Calliari I., Effects of Cold Rolling and Strain-Induced

Martensite Formation in a SAF 2205 Duplex Stainless Steel, Metallurgical and Materials Transactions

A-Physical Metallurgy and Materials Science, vol. 46A, 2015, pp. 577-586

[2] Fiorillo F., Measurement and Characterization of Magnetic Materials, Elsevier, Amsterdam, 2004

[3] Takacs J., Mészáros I., Separation of magnetic phases in alloys, Physica B, vol. 403, 2008, pp. 3137-

3140

[4] Stablein F., Steinitz, Ein neuer Doppeljoch-Magnetsthalprüfer, R. Arch Eisenhüttenwesen, vol. 8, 1935,

pp. 549-554

[5] http://www.fischer-technology.com/fileadmin/documents/broc/EN/BROC_FMP30_FERITSCOPE_902-

039_en.pdf

– 63 –

DIFFICULTIES CAUSE BY MAGNETIC AFTER-EFFECT DURING

IDENTIFICATION OF THE PREISACH HYSTERESIS MODEL

WEIGHTING FUNCTION

M. Novak

Technical University of Liberec, Studentská 2, CZ 46117 Liberec, The Czech Republic, e-mail: [email protected]

Abstract. The time dependency of the magnetic hysteresis cause by thermal activation over an energy barrier is called

after-effect, magnetic viscosity or magnetic relaxation. Magnetic after-effect influences the hysteresis loop shape

depending on the excitation filed rate of change, sample geometry and state. The magnetizing loop changes, especially

rounding of peaks at steep part of the loop, severely influences process of the Preisach model weighting function

identification from experimental data. This article concern on the after-effect measuring with the aim of determines

limiting excitation speed.

I. INTRODUCTION

The time dependence of a ferromagnetic material magnetization under a constant magnetic

exciting field is called magnetic after-effect. This phenomenon occurs due to thermal activation of

the irreversible magnetization processes [1]. It is a result of approaching the thermodynamic

equilibrium with minimum free energy in material where the energy distribution is complicated.

Magnetic domain walls can be trapped in local minimums of energy for a long time until they are

excited by thermal fluctuations and they overcome the energy barrier. The material exhibits

thermally activated Barkhausen jumps and it moves step by step towards the low energy state. The

rate of this relaxation depends on the distribution of energy barriers and temperature.

The overall magnetization compose of reversible and irreversible component M(t) = Mrev(t) +

Mirr(t). Although, the distribution of energy barriers can be general function the most ferromagnetic

materials exhibit logarithmic decay of magnetization over time Mirr(t) = M0 – Sln(t/t0), where t is

the time since changing the excitation field, M0 = Mirr(t = t0) and S is the relaxation coefficient. The

magnetic after-effect can be observed at change of the field and it requires much longer time to

attain new steady state compared to the much faster eddy current effect.

The First Order Reversal Curves (FORC) method is one of the methods used for identification

of the weighting function (WF) of the Preisach model of hysteresis (PMH) [2]. It is based on

measuring set of hysteresis loop. Each loop starts from negative saturation level. Consequent loops

have gradually increased maximum of the exciting field strength HU. The descending branch of the

loop, the first order reversal part, is inserted into the 2D matrix. The magnetization of descending

branch M(HD) creates rows of the matrix and are placed into the columns in accord to HU. Such

surface is called the Everett function. Second partial derivative of the Everett function resulting in

the WF of PMH:

DU

DU

DUHH

HHMHHw

),(

2

1,

2

. (1)

The rounding of magnetizing loops caused by after-effect takes place especially at the area of

steep part where irreversible magnetization process dominates. Rounded tops of the magnetizing

loop make impossible to determine values of the maximal filed strength HU. There are several

approaches how to find the HU in case of rounded curve e.g. from maximal field strength max(H),

maximal energy product max(HB), maximal flux density max(B) etc. Inaccurate determination of

HU causes misalignment of rows in the WF matrix and shift of data in rows leading to deterioration

of WF and simulated hysteresis loops. The only correct way how to obtain proper results is slow

down frequency of the excitation field during measuring. On the other side the extra low frequency


– 64 –

measurement brings other kind of problems e.g. with the offset drift during the magnetic flux

integration, offset drift of exciter and so on. The aim of this article is to determine the time constant

of the magnetic after-effect and proper frequency setting of FORC measurement.

II. MAGNETIC AFTER-EFFECT MEASUREMENT

The standard Epstein frame was used for the experiment with sample of grain oriented steel

M165-35S. First, the major loop was measured at a frequency of 0.1 Hz. This major loop was used

to determine the jumps of the excitation current so that the change in the flux density in one step

was a constant BP = 50 mT. Then, the magnetization loop was measured again by these

discontinuous jumps. The secondary voltage response was captured for five seconds after each

jump. Corresponding magnetic flux density transients are shown in Fig. 1 (right hand graph) for the

different positions on the magnetization loop (marked in the left hand graph). The flux density time

constant varies from milliseconds at saturation region up to more than two seconds in steep part of

the loop and again shorten when approaching to the negative saturation.

Fig.1. Time response of flux density to small jumps of field strength for different points at the magnetizing loop

The theory of the after-effect expects exponential response. The measurement revealed that

this theoretical assumption is valid in an area where time constants are small but at the steep part of

the loop the response is irregular with very long full stabilization time.

III. CONCLUSION

This experiment has shown that for the correct measurement of FORC it is necessary to use

a frequency of less than 1 mHz. An alternative can be a trapezoidal excitation signal with

persistence at HU value.

REFERENCES

[1] Abeywickrama N., Serdyuk Y. V., Gubanski S. M., Effect of Core Magnetization on Frequency

Response Analysis (FRA) of Power Transformers, IEEE Trans. on Power Delivery, vol. 23, no. 3,

2008, pp. 1432-1438

[2] Bertotti G., Mayergoyz I., The science of hysteresis, vol. 1, 2 and 3. (1st ed.), Elsevier, 2005

– 65 –

ANISOTROPY OF SPECIFIC TOTAL LOSS COMPONENTS IN GOSS

TEXTURED ELECTRICAL STEEL

W.A. Pluta

Czestochowa University of Technology, Al. Armii Krajowej 17, 42-200 Czestochowa, Poland


Abstract. The production process of electrical steel sheets (ES) can be carried out in such a way that in the final

product crystals are ordered in rolling direction (RD). As a result ES displays different properties in different

magnetization directions. An investigation of the specific power loss separation of electrical steel sheets in different

direction to rolling direction was performed. The measurements were carried out in the in non-standard Single Sheet

Tester at ten different frequencies and for different angles to rolling directions. The separation of the specific total loss

into three components was performed. The investigation shows applicability of three components specific total loss

model and influence of magnetic anisotropy on loss components.

I. INTRODUCTION

The production process of electrical steel sheets (ES) can be carried out in such a way that in

the final product crystals are ordered in rolling direction (RD). I this direction ES displays most

favorable magnetic properties and in directions 55 and 90 appears poor magnetic properties. The

amount of crystals oriented along RD in relation to whole amount of crystals decides about

directional properties of ES. This is usually described by degree of texture being a measure of

amount of crystals oriented along RD in relation to whole amount of crystals. Another way to

describe the directional properties is the anisotropy of magnetic properties e.g. flux density or

anisotropy of specific total loss. Generally, magnetic anisotropy is determined for a given magnetic

parameter at a given value of the abscissa y. For example, the anisotropy of specific total loss 090

5.1,

SAP is calculated for the magnetization angles x = 90 and x = 0 at the flux density 1.5 T. For

analysis of anisotropic properties of specific total loss different models as model based on:

Orientation Distribution Function (ODF), polynominal approximation, the reluctivity tensor or

phase Neel’s theory. The anisotropy phenomenon play important role in construction of magnetic

circuits. Cores made of grain oriented ES for construction of magnetic circuits of transformers,

generators and large rotating machines are used taking into account the direction of sheet

production. The quotient of losses of made magnetic circuit and losses of magnetic material,

measured by standardized methods, determines the quality of the magnetic circuit, the so-called

building factor. In the case of a simple single-phase magnetic circuit packaged with strips cut at an

angle of 90, the building factor is 1.45, and for cut at an angle of 45 the building factor is 1.1.

Therefore, taking into account the anisotropic magnetic properties of ES at the design stage,

significant energy and material savings can be achieved and the technical parameters of the device

can be improved, such as noise or vibration.

II. EXPERIMENTAL SETUP

Measurements were taken under axial examination in a non-standard Single Sheet Tester

(SST) on square samples of 100 mm width on conventional five grades of grain-oriented (GO) ES.

The ES grades varies by thickness form 0.27 mm to 0.35 mm and differs by specific total loss

anisotropy from about 50% to 60%. Measuremts were performed at 10 frequencies for 2 Hz to 100

Hz at different angles to RD.


– 66 –

III. RESULTS

It is overall accepted that the specific total loss PS consist of three components: hysteresis,

classical and excess eddy current. The frequency dependence of loss can be described by the three

components model [1] and it can be applied to any angles x to the RD as follows:

fP

/

p

x

ex

fP

x

pce

fP

p

x

h

x

S

xexce

xh

fBCfBCBCfP

/

212.3

/

2

/

/ (1)

where: Chx is the hysteresis loss coefficient, is the exponent of flux density, Cce is the classical

eddy current loss coefficient, Cexx is the excess loss coefficient.

In Fig.1 are shown experimental data (points) of energy loss fitted using (1) for different

magnetizing directions obtained for GO ES grade M150-35S at Bm = 1.0 T and 1.2 T according to

(1).

Fig. 1. Energy loss per unit mass versus frequency for different magnetizing directions obtained for ES grade M150-

35S at: a) Bm = 1.0 T, b) Bm = 1.2 T

As can be in Fig.1 only classical eddy current specific total loss shows isotropic characters.

This is due to the fact it is calculated for perfectly conducting infinite homogenous plate. The

hysteresis and excess eddy current loss components display anisotropic character. Additionally,

both components show similarity due to their common origin [1, 2]. In Fig. 1 is visible the non-

linearity of frequency dependence of anisotropy of specific total loss.

REFERENCES

[1] Bertotti G., Hysteresis in magnetism, Academic Press, 1998

[2] Pluta W.A., Angular properties of specific total loss components under axial magnetization in grain-

oriented electrical steel, IEEE Trans. on Magnetics, vol. 52, no 4, 2016, pp. 6300912

0 20 40 60 80 100 1200.00

0.01

0.02

0.03

0.04

0.05

0.06

f , Hz

Ph+a

/ f, J/kg

Ph+a

(0)

Ph+a

(90)

54o

90o

0o

Ph+a

(54)

0 20 40 60 80 100 1200.00

0.01

0.02

0.03

0.04

0.05

0.06

f , Hz

Ph+a

/ f, J/kg

Ph+a

(0)

Ph+a

(90)

54o

90o

0o

Ph+a

(54)

Pce Pce

– 67 –

SURFACE ISOLATION OF MODERN ELECTRICAL TYPES FOR

MAGNETIC CORES

W.A. Pluta

Czestochowa University of Technology, Al. Armii Krajowej 17, 42-200 Czestochowa, Poland


ABSTRACT

With the development of distributed energy sources, the challenge is the most effective

transformation of electrical energy. The conversion of electrical energy can be defined as the

conversion of current, voltage and frequency to a different set of such values [1]. Simultaneously it

is present tendency to increase energy density of newly build devices by increase of frequency.

The increase of frequency cause increase of importance of eddy currents and of the skin effect.

Also with increase of frequency increases electrical strain and the requirements to the quality of

insulation layer. The insulation layer must also be resistant to mechanical and temperature stresses

present as well as during production process as during magnetic core manufacturing. Additionally

it must be also economically justified. For example to obtain optimum magnetic properties in

magnetic cores made from nanocrystalline types toroidal cores are first wound in their final

configuration and then annealed with a circumferential or perpendicular magnetic field applied to

the toroid. This anneals serves to relieve stresses in the metallic glass ribbons resulting both from

the rapid quench during casting of the ribbons and from bending stresses in the ribbons due to the

curvature of the ribbon in the toroidal core. The applied magnetic field during the anneal serves to

induce an easy direction of magnetization along the field direction. By field annealing cores made

from metallic glass ribbons, cores with very square B-H loops can be produced..

For mentioned reasons more and more attention is devoted to the surface of isolation of

electrical tapes and the selection of optimal materials due to their properties and costs [2]. The

paper presents different methods of applying and testing the insulation layers on electrical tapes.

In this paper different issues concerning requirements, application and testing of insulation

layers on different magnetic material are described.

REFERENCES

[1] Shen W., Wang F., Boroyevich D., Tipton IV C.W., High-density nanocrystalline core transformer for

high-power high-frequency resonant converter, IEEE Trans. on Industry Applications, vol. 44, no.1,

2008, pp. 213-218

[2] Beckley P., Electrical steel for rotating machines, The Institution of Engineering and Technology,

Power and energy series No 37, Glasgow, 2002


– 68 –

– 69 –

MAGNETIC AND MECHANICAL PROPERTIES OF RUBBER BONDED

MAGNETS WITH DIFFERENT TYPE AND AMOUNT

OF HARD MAGNETIC POWDER

M. Przybylski1, B.Ślusarek

1, T.Bednarczyk

2 and G.Chmiel

2

1 Instytut Tele- i Radiotechniczny, 03-450 Warszawa, ul. Ratuszowa 11, e-mil: [email protected],

[email protected] 2 GUMET Sz. Geneja Spółka Jawna, 23-200 Kraśnik, ul. Kolejowa 12, e-mail: [email protected],

[email protected]

Abstract: Application of permanent magnets bonded by rubber is still growing, especially in automotive industry.

Magnetic and mechanical properties of rubber permanent magnets can be tailored by a production's method and a

magnet's composition. Permanent magnetsbonded by rubber are produced by a method called calendaring.Physical

properties of rubber bonded permanent magnets depend on a type and amount of hard magnetic powder in a mixture

with rubber. Anisotropic strontium ferrite powder and spherical isotropic Nd-Fe-B alloy powder obtained by

atomization were used in research. Results of measurements show thatwith increasing amount of ferrite powder

magnetic properties and Shore hardness increase whereas tensile strength decreases. Addition of Nd-Fe-B powder to

the mixture slightly increases magnetic properties of magnets.

I. INTRODUCTION

Application of permanent magnets is constantly growing. One ofindustry brancheswhere

application of permanent magnets is constantly growing is an automotive industry.Permanent

magnets in this industry are applied, among others, in rotary magnetic encodersfor ABS (Anti-Lock

Braking System) systems.

Multipole permanent magnets for rotating encoders for ABS systemsare,among others,

prepared by technology of bonding hard magnetic powder by rubber. Properties of this type of

permanent magnets depend mainly ona type and amount of hard magnetic powder and kind of

a cross-linker [1-2].

The aim of investigation isto show influence of a kind and amount of hard magnetic powder on

magnetic and mechanical properties of permanent magnets.

II. EXPERIMENTAL DETAILS AND RESULTS

Technology of production permanent magnets from hard magnetic powder bonded by rubber

consists in preparing a mixture of powder with rubber and a cross-linker with additives, then

a vulcanization process of this mixture is conducted. The last operation is magnetization of

samples. In the experiment magnetic powder of strontium ferrite and Nd–Fe- B powder were used.

The mixture of powder and rubber with a cross-linker were prepared by calendaring process.

The first set of mixtures contains from 76.3 weight % to 88.2 weight % of strontium ferrite. The

second set of samples with mixture of strontium ferrite from 83.3 to 69.0 weight % and spherical

atomized powder of Nd-Fe-B from 4.8 weight % to 19.0 weight % were prepared as well. The

powder of Nd-Fe-B was a powder designed for injection moulding technology.

A sheet of rubber with magnetic powder was prepared. Samples for measurement of magnetic

and mechanical properties were prepared in a vulcanization process in a temperature 160°C for 15

minutes.

The results of investigation are shown in Table 1.





– 70 –

Table 1. Magnetic properties of rubber permanent magnets

An amount of ferrite and Nd-Fe-B powder (weight %),

the rest is rubber, cross-linker and additives

Density

(g/cm3)

Br

(mT)

HcB

(kA/m)

HcJ

(kA/m)

BHmax

(kJ/m3)

Strontium ferrite –76.3 % 2,85 173 129 262 5,72






Strontium ferrite – 83.3 %, Nd-Fe-B – 4.8 % 3,48 237 158 222 10,44

Strontium ferrite -78.5 %, Nd-Fe-B – 9.5 % 3,52 240 159 239 10,54

Strontium ferrite – 69.0 %, Nd-Fe-B –19.0 % 3,60 251 175 318 11,46

As Table 1 shows with increase of an amount of strontium ferrite powder in mixture of powder

and rubber magnetic properties of samples grow. It was impossible to prepare a mixture of rubber

with a larger amount of strontium ferrite. The mixture of rubber with strontium ferrite and Nd-Fe-B

powder were prepared for increase of magnetic properties of magnets. The small increase of

magnetic properties is observed in samples with Nd-Fe-B powder. Powder Nd-Fe-B for injection

moulding technology has a value of median particle size about 35-55µm, but the value of strontium

ferrite powder is1.05 µm. In future experiments Nd-Fe-B powder with smaller median particle size

will have to be used. It should allow powders with rubber to be mixed better and, in consequence,

better magnetic properties of magnets will be obtained.

Mechanical properties of samples were measured. The result of measurements are shown in

Figure 1.

Fig.1. Mechanical properties of rubber permanent magnets

As Figure 1 shows with increase an amount of strontium ferrite powder hardness of samples

increase, whereas tensile strength decrease.

A rubber permanent magnet ring with 96 alternating magnetic poles was prepared and will be

applied in ABS system.

REFERENCES

[1] Soloman M.N.,Kurian P., Anantharaman M.R., Joy P.A., Cure characteristics and dielectric properties

of magnetic composites containing strontium ferrite, Journal of Elastomers& Plastics, vol. 37, Issue: 2,

2005, pp. 109–120

[2] Kruzelak J., Hudec I., Dosoudil R., Sykora R., Investigation of strontium ferrite activity in different

rubber matrices, Journal of Elastomer and Plastics, vol. 47, Issue: 3, 2015, pp. 277-290

– 71 –

A COMPUTATIONAL AND EXPERIMENTAL STUDY OF SHAPE

MEMORY ALLOY SPRING ACTUATOR

D. Stachowiak and M. Kurzawa

Poznan University of Technology, Piotrowo 3a, 60-965 Poznań, Poland,


Abstract. The paper presents the combined experimental and computational study of the shape memory alloy spring

actuator. The design strategy for a system consisting of two springs: a SMA spring and a steel spring has been

presented. The distribution of forces in the designed system for high and low temperature condition has been

calculated and measured. A prototype of linear actuator with SMA spring and a biasing steel spring and an

experimental setup have been designed to perform the electro-thermo-mechanical characterization of SMA spring. The

selected results of calculation and laboratory tests of the designed spring system have been given.

I. INTRODUCTION

Shape memory alloys (SMA), according to their ability to revert to their programmed shape

through thermal activation have a great potential for a wide range of actuator application [1,2,3].

The SMA have two stable phases - the high-temperature phase, called austenite and the low-

temperature phase, called martensite. In addition, the martensite phase can be in one of two forms:

twinned and detwinned [1, 2]. Transformation of phase which occurs between these two phases

upon heating or cooling is the basis for the unique properties of the SMA. The main effects of

SMA associated with the phase transformation are pseudoelasticity and shape memory effect [1].

The pseudoelasticity occurs when the martensitic phase transformation is stress-induced at

a constant temperature. This effect applies for most of nowadays shape memory applications in the

field of medical devices [3]. The shape memory effect refers to the ability of the material, initially

deformed in its low-temperature phase, to recover its original shape upon heating to its high

temperature phase. The shape memory effect may be one-way or two-way effect [1, 2]. To provide

the necessary reversible shape memory effect, two methods: intrinsic and extrinsic can be used.

Intrinsic methods consist of modifying the material microstructure so that certain martensitic

variants orientations will preferably nucleate upon cooling. The intrinsic methods are also called

training processes [2]. Extrinsic methods refer to the addition of an external element coupled to the

SMA material that provides the required stress to induce stress-oriented variants [2].

The shape memory effect of SMA provides possibilities of using it as actuators. The SMA can

be formed into almost any shaped actuator. In the paper the electro-thermo-mechanical characteristics

of the reversible shape memory effect SMA spring have been investigated.

II. ACTUATOR DESIGN, SELECTED RESULTS AND CONCLUSIONS

Usually a SMA actuator consists of at least one actuator element and at least one return

element. The return element can be either a dead-weight or a bias spring or another SMA

(antagonist configuration), etc. In the paper a SMA-spring actuators configuration using as active

element electrically driven SMA spring working against to a conventional steel spring have been

considered. Figure 1 shows the total system made up of a SMA spring and a biasing steel spring.

A SMA spring coupled to a bias spring is preloaded so that the system is under stress. Upon

heating the SMA spring transforms back to the high-temperature phase and pushes the steel spring

as it tries to recover its original shape. The return element actions the reformation of the actuator

element into its original shape during cooling (the two-way motion). The output of the mechanism

is taken between the SMA and the bias spring.

The software for designing calculation of SMA spring and for determining the distribution of

forces in a system consisting of two springs was elaborated. The distribution of forces in the

designed system for high and low temperature condition has been calculated using in house

software. The calculated and measured forces have been presented in Fig. 2.



– 72 –

Fig.1. The total system made up of a SMA spring

and a biasing steel spring Fig.2. Diagram of forces acting in the systems

The authors elaborated on the special experimental setup for the testing SMA spring actuator.

The view of the elaborated laboratory stand has been shown in Fig. 3. Selected field distributions

obtained from the thermal camera have been shown in Fig. 4. The selected dynamic characteristics

determined at the experimental setup have been shown in Fig. 5 and 6.

Fig.3. The experimental setup Fig.4. Test of the SMA spring using thermal camera

Fig.5. The stroke of the SMA spring vs. time at

different current values Fig.6. The temperature of the SMA spring vs.

time at different current values

The dynamic characteristics have been presented, taking into account changes in the length of

the SMA spring and temperature as a function of time with a linearly increasing load and at

stepping on and off the current. It has been found that SMA spring can successfully be employed to

provide linear displacement. This study could be useful in precisely controlling of SMA spring

actuator.

REFERENCES

[1] Lagoudas D.C., Shape Memory Alloys: Modeling and Engineering Applications, Springer, 2008

[2] Czechowicz A., Langbein S., Shape Memory Alloy Valves - Basics, Potentials, Design, Springer

Verlag, 2015

[3] Mohd Jani J., Leary M., Subic A., Gibson M.A., A review of shape memory alloy research,

applications and opportunities, Materials and Design, vol. 56, 2014, pp. 1078-1113

– 73 –

NEW DEVELOPMENTS IN RAPIDLY QUENCHED

SOFT AND HARD MAGNETIC ALLOYS

P. Svec1, I. Janotova

1, D. Janickovic

1, B. Kunca

2, J. Marcin

2, I. Matko

1,

I. Skorvanek2 and P. Svec Sr.

1

1 Institute of Physics, Slovak Academy of Sciences, Bratislava, Slovakia, e-mail: [email protected]

2 Institute of Experimental Physics, Slovak Academy of Sciences, Kosice, Slovakia, e-mail: [email protected]

Abstract. New trends in rapidly quenched soft magnetic materials with focus on enhanced physical properties

(saturation magnetization, coercivity, operating temperatures, frequencies, etc.) will be presented. The use of rapid

quenching will also be shown on the case of rare-earth free hard magnetic materials based on Mn-Al and Mn-Bi.

Importance of diverse aspects of processing for property optimization will be demonstrated on selected examples of

alloy systems.

I. SOFT MAGNETIC MATERIALS

The necessity for controlled use of critical elements stimulates research on rare-earth free or at

least rare-earth-poor soft and hard magnetic materials. In soft magnetics there is a demand for

materials with high saturation magnetization combined with low coercivity, which might surpass

the well known and still developed excellent rapidly quenched nanocrystalline soft magnetic

materials as FINEMET, NANOPERM and HITPERM. In our work attention will be put on

systems similar to NANOMET-type systems based on Fe-B with high Fe content and small

additions of specific elements which fulfill these requirements. Selected results will be presented

on Fe-B based system alloyed with Co, Si and P [1] together with results obtained on a new system

based on Fe-Sn-B [2]. It will be shown that using compositional tuning, special preparation and

processing algorithms leading to optimal crystal size, phase content, ribbon thickness and magnetic

domain structure it is possible to optimize the desired magnetic properties. An example of

compositional tuning to obtain grain-refined nanocrystalline structure in rapidly quenched Fe-Sn-B

is shown in Fig. 1, where the addition of Sn up to 7 at.% into Fe85B15 leads upon annealing to

transformation of amorphous structure into nanograins of ~20-25 nm in size embedded in

amorphous remains. Such morphology is nearly identical to that observed in nanocrystallized

FINEMET, NANOPERM or HITPERM systems.

Fig.1. Effect of Sn addition on microstructure refinement in rapidly quenched amorphous

Fe85-xSnxB15 for x = 3.5, 5 and 7 at. % after annealing at 700K for 30 min (left, middle and right images, respectively).

The marker in all three images (bottom left) is 50 nm.



– 74 –

II. RARE-EARTH FREE HARD MAGNETIC MATERIALS

In rare-earth free Mn-based permanent magnets the main issue is to obtain Al-Mn and Bi-Mn

alloys with maximized amounts of hard ferromagnetic tau-AlMn or alpha-BiMn phases,

respectively. The formation of these phases prepared by transformation from rapidly quenched

precursors [3, 4] will be shown by conventional in-situ transmission electron microscopy methods

(Fig. 2) together with the development of their microstructure and magnetic properties using high

magnetic field annealing.

In order to assess the micromechanisms of formation of hard magnetic phase tau-AlMn from

the as-quenched matrix containing mainly epsilon-AlMn phase details of microstructure evolution

on atomic level during phase transformation of rapidly quenched Al45Mn55 will be presented. Data

from atomically resolved scanning transmission electron microscopy and electron energy loss

spectroscopy on samples annealed isothermally ex-situ at selected temperatures as well as in-situ

using dedicated heating holders will be shown. Special features of the phase transformation and

chemically resolved local atomic arrangements will be presented indicating the processes

controlling the type of transformation.

Fig.2. In-situ isothermal transformation at 693 K of as-quenched epsilon-AlMn into ferromagnetic tau-AlMn phase.

Left image t = 0 min., right image t = 90 min, epsilon-AlMn phase growing gradually from left to right in form of

heavily twinned crystals

ACKNOWLEDGEMENT

Support of M-era.Net NEXMAG, APVV-15-0621 and VEGA 2/0082/17 projects is gratefully

acknowledged.

REFERENCES )

[1] Janotova I., Zigo J., Svec P., Matko I., Janickovic D., Svec Sr. P., Analysis of phase transformations in

Fe–(Co)–B–Si–(P), J. Alloys and Compounds, vol. 643, 2015, pp. S265-S269

[2] Matko I., Illekova E., Svec Sr. P., Svec P., Janickovic D., Vodarek V., Microstructural study of the

crystallization of amorphous Fe–Sn–B ribbons, J. Alloys and Compounds, vol. 615, 2015, pp. S462-

S466

[3] Palanisamy D., Srivastava Ch., Madras H., Chattopadhyay K., High-temperature transformation

pathways for metastable ferromagnetic binary Heusler (Al–55 at.%Mn) alloy, J. Mater. Sci., vol. 52,

2017, pp. 4109-4119

[4] Janotova I., Svec Sr. P., Svec P., Matko I., Janickovic D., Zigo J., Mihalkovic M., Marcin J.,

Skorvanek I., Phase analysis and structure of rapidly quenched Al-Mn systems, J. Alloys and

Compounds, vol. 707, 2017, pp. 137-141

– 75 –

JILES-ATHERTON-SABLIK MODEL OF MAGNETO-MECHANICAL

CHARACTERISTICS OF SOFT MAGNETIC MATERIALS - A REVIEW

R. Szewczyk, A. Bieńkowski and M. Nowicki

Institute of Metrology and Biomedical Engineering, Warsaw University of Technology,

ul. św. A. Boboli 8; 02-525 Warszawa, Poland, e-mail: [email protected]

Abstract. The paper presents recent advances in development of Jiles-Atherton-Sablik model of magnetic hysteresis

loops. Progress in modeling an anhysteretic magnetization curve concerning stress-induced anisotropy is described.

Moreover, different approaches to differential equations stating the hysteresis model are presented. Finally the

methods of parameters identification together with estimation of the influence of stresses on model parameters are

elaborated.

I. INTRODUCTION

The principles of Jiles-Atherton model of magnetic hysteresis loop were first introduced in

1984 and fully developed in 1986 [1]. Since then, this is one of the most popular models with wide

range of applications. Jiles-Atherton model is useful for modeling the characteristics of inductive

components for SPICE (Simulation Program with Integrated Circuits Emphasis), FEM (Finite

Elements Method) and MoM (Method of Moments) methods.

Moreover, Jiles-Atherton model was expanded by Sablik et al. in 1993 [2], to present one of

the first quantitative explanations of magnetoelastic phenomena. Since then, in spite of criticism

[3], so called Jiles-Atherton-Sablik model is the important frame for analyses of magnetic,

magnetoelastic and magnetostrictive effects.

Paper presents the review of recent advances in development of Jiles-Atherton-Sablik model

with special stress on explanation of magnetoelastic effects. Steps towards consideration of

macroscopic anisotropy (including stress induced anisotropy) of magnetic materials are explained.

Alternative concepts of differential equations stating hysteresis in the model are also analyzed.

Additionally, paper presents the methods of identification of parameters of Jiles-Atherton-Sablik

model, as well as the most important computational problems connected with solving the equations

of Jiles-Atherton-Sablik model.

II. JILES-ATHERTON-SABLIK MODEL

The Jiles-Atherton-Sablik model is based on the concept of anhysteretic magnetization curve

commonly described by the Langevin equation. However, due to the fact, that the model of this

curve considers Boltzman statistics [1], the Langevin form of anhysteretic curve is valid only for

strictly isotropic materials. Analyses confirm, that different types of anisotropy can be considered,

such as uniaxial anisotropy [4] or magneto-crystalline anisotropy energy of single cubic crystals

[5]. Moreover, stress induced, uniaxial anisotropy can be also modeled due to the fact, that average,

stress-induced anisotropy density Kan is given by the following equation:

Kan =3

2σλs σ sin2 ψ (1)

where s is saturation magnetostriction, are uniaxial mechanical stresses and is the angle

between stresses and direction of magnetization. However, it should be taken into account that

mechanical stress dependence of saturation magnetostriction s is quantum effects-based

phenomenon, which quantitative description is still not fully understood. Magnetic materials

subjected to uniaxial stresses are not isotropic and the Langevin form of anhysteretic curve is not


– 76 –

valid. Mechanical stresses influence the Bloch interdomain coupling [2] which significantly

changes the shape of anhysteretic curve of magnetic material subjected to mechanical stresses.

The origins of the ordinary differential equation stating the hysteresis in Jiles-Atherton model

are not clearly explained, especially that original calculations neglect the chain rule [6]. Due to this

fact, alternative forms of hysteresis description were presented by Venkataramann et al. [7], Cheng

et. al. [8].

III. COMPUTATIONAL PROBLEMS AND METHODS OF PARAMETERS

IDENTIFICATION

Solving of ordinary differential equation stating the Jiles-Atherton-Sablik model of magnetic

hysteresis is not a trivial task. Previously presented Riemann method based approach [9] lead to

significant numerical errors, which cumulates during the numerical integration. For this reason,

Runge-Kutta based algorithms are applied for calculations.

It should be stressed that possibility of application of optimization based methods of

identification of Jiles-Atherton model parameters is limited. Even with the use of two-steps

optimization method and differential evolution based optimization, the identification of parameters

may lead to ambiguous results [10]. As a result, the alternative, physical dependences based

methods are intensively developed [11].

IV. CONCLUSIONS

In spite of over thirty years of development of Jiles-Atherton-Sablik model, many problems

connected with this model seem to be still unsolved. However, works connected with this model

leads to better understanding of physical phenomena behind the magnetic hysteresis and magneto-

mechanical interactions.

To reduce the severity of problems connected with the numerical calculations of Jiles-

Atherton-Sablik model and enable validation of the results, open source OCTAVE/MATLAB

scripts were developed and freely distributed at: www.github.com/romanszewczyk/JAmodel

REFERENCES

[1] Jiles D., Atherton D., Theory of ferromagnetic hysteresis, J. Magn. Magn. Mater., vol. 61, 1986, pp. 48

[2] Sablik M., Jiles D., Coupled magnetoelastic theory of magnetic and magnetostrictive hysteresis, IEEE

Trans. Magn., vol. 29, 1993, pp. 2113

[3] Zirka S.E., Moroz Y.I., Harrison R.G., Chwastek K., On physical aspects of the Jiles-Atherton

hysteresis models, J. Appl. Phys., vol. 112, 2012, pp. 043916

[4] Ramesh A., Jiles D.C., Roderik J., A model of anisotropic anhysteretic magnetization, IEEE Trans.

Magn., vol. 32, 1999, pp. 4234

[5] Baghel A., Kulkarni S. V., J. Appl. Phys., vol. 113, 2013, pp. 043908

[6] Szewczyk R., Cheng P., Open Source Implementation of Different Variants of Jiles-Atherton Model of

Magnetic Hysteresis Loops, Acta Physica Polonica A., vol. 133, 2018, pp. 654

[7] Venkataraman R., Krisnaprasad P.S., Qualitative analyse of a bulk ferromagnetic hysteresis model,

Proceedings of the 37th IEEE Conference on Decision and Control, 1998

[8] Cheng P., Szewczyk R., Modified description of magnetic hysteresis in Jiles-Atherton model,

AUTOMATION 2018, AISC 743, 2018, pp. 648–654

[9] Calkins F., Smit R., Flatau A., Energy-based hysteresis model for magnetostrictive transducers, IEEE

Trans. Magn., vol. 36, 2000, pp. 429

[10] Szewczyk R., Nowicki M., Explicitness of Jiles-Atherton model parameters identified during the

optimization process, International Conference APCOM 2018, Slovakia

[11] Chwastek K., Szczyglowski J., Identification of a hysteresis model parameters with genetic algorithms,

Mathematics and Computers in Simulation, vol. 71, 2006, pp. 206-211

– 77 –

OPTIMAL FLIGHT DIRECTION OF MAGNETIC SYSTEM DURING

OBJECT'S DETECTION ON THE BALTIC SEA

M. Woloszyn1, S. Michalski

2 and B. Potrac

2

1 Gdansk University of Technology, Faculty of Electrical and Control Engineering,

G. Narutowicza 11/12, 80-233 Gdansk, e-mail: [email protected] 2 Gdansk University of Technology, Marine Military Technologies Centre

Abstract. The paper presents the problem of object’s detection on the Baltic Sea. The big magnetic anomalies on the

Baltic Sea hinder the detection of ferromagnetic objects by using a magnetic system installed in a gondola. A gondola

is towed by helicopter and during a flight is deviating perpendicular to the direction of a movement. The deviations

cause magnetic disturbances that make it difficult to detect the object. For this reason, the optimal direction of the

flight is vital for detecting objects.

I. INTRODUCTION

There are great magnetic anomalies on the Baltic Sea which amount to several thousand nT

over a dozen kilometers (Fig.1). A magnetic disturbance appears during the measurement of

a magnetic signal causes a track deviation of a gondola in which a magnetic sensor is installed

(Fig.2). The optically pumped magnetometers are most commonly used in measurements. These

magnetometers measure a modulus of the magnetic field density with a resolution of about

5 pT/Hz0.5. In order to use a high sensitivity of magnetometers, the influence of the track deviation

of a gondola into measurements should be minimized.

Fig.1. The magnetic anomalies on the Baltic Sea (near

Ustka city)

Fig.2. The track deviation of a gondola

II. OPTIMAL DIRECTION OF THE FLIGHT

The method of compensation of magnetic disturbances should be used in a magnetic system

installed in a gondola and also on another platform [1, 2]. This method does not take into account

disturbances caused by a great magnetic anomaly. The difference of the modulus magnetic flux

density along lx (Fig.1) for y = 10 m is shown in Fig.3 and along ly (Fig.1) for x = 10 m in Fig.4.

The deviation of the gondola depends on its aerodynamical properties, length of a cable-line and on

meteorological conditions. The amplitude of the deviation can take 5 m or more (when the wind is


– 78 –

perpendicular to the flight direction). The optimal direction of the flight is the direction

perpendicular to the magnetic isoclines (lx trajectory – Fig.1).

Fig.3. The difference of the modulus magnetic flux

density along lx

(Fig.1) for y = 10 m

Fig.4. The difference of the modulus magnetic flux

density along ly

(Fig.1) for x = 10 m

Fig.5. The disturbances causes deviation (amplitude 5 m) of the gondola

along lx and ly lines (Fig.1).

III. CONCLUSIONS

Searching for a sunken ship in the Baltic Sea requires an optimal flight for minimization of an

influence of great magnetic anomalies. The best direction is perpendicular to the magnetic

isoclines. In this case gondola’s deviations has minimal influence on the magnetic measurements.

REFERENCES

[1] Leliak P., Identification and Evaluation of Magnetic Field Sources of Magnetic Airborne Detector

Equipped Aircraft, IRE Trans. Aerospace and Navigational Electronics, vol. 8, 1961, pp. 95-105

[2] Allen G., Matthews R., Wynn M., Mitigation of Platform Generated Magnetic Noise Impressed on a

Magnetic Sensor Mounted in an Autonomous Underwater Vehicle, MTS/IEEE Oceans, 1999, pp. 63-71

– 79 –

A DEVICE FOR THE STUDY OF ELECTRICAL STEEL LOSSES

IN STATOR LAMINATION STACKS

S. Nazrulla1, E.G. Strangas

1, J.S. Agapiou

2 and T.A. Perry

2

1 Electrical Machines and Drives Laboratory, Departmentof Electrical and Computer Engineering, Michigan State

University, East Lansing, MI 48824-1226 USA, e-mail: [email protected] 2 Manufacturing Systems ResearchLaboratory, General Motors Technical Center,Warren, MI 48092 USA,

e-mail: [email protected]; [email protected]

Abstract. A new electromagnetic device to measure the electrical losses in stator lamination stacks, along with its

associated test procedure, is presented. This procedure provides the ability to distinguish between the qualities of

stators made of different types of materials, and can be employed to evaluate finished stator stacks prior to motor

assembly. The design and simulation of the proposed device is documented, along with experimental data supporting

our conclusions.

I. INTRODUCTION

Laminations of electrical machines are affected by variations in the raw material,

manufacturing process and subsequent handling. Both product and manufacturing engineering have

an interest in measuring the magnetic properties of lamination stacks as an assembly because it is

not simple to correlate the performance of the steel used in the laminations to the assembled stator

stack. This is critical to achieve the highest possible performance in electric machines at

a consistent quality, independent of material variations, and tooling wear out. This paper examines

a new method of measuring the losses in stator stacks prior to full motor assembly – post-assembly

evaluation is common, for instance, in conventional dynamometer efficiency testing – thus

avoiding complications from bearings, windings, or other elements.

Estimating losses from the magnetic steel characteristics is in itself a complex task with

inaccurate results. Although analytical and numerical tools showed a great improvement, the

accuracy of the calculations is not adequate to allow comparisons between different steels and

treatment of iron cores. To characterize the material a number of efforts have been made. Among

them some experimental procedures were developed, e.g. for characterization of electromagnetic

phenomena that occur in the end regions of large turbo-generators.

II. TECHNICAL DISCUSSION

A device utilizing a magnetic probe was developed to impose a time-dependent magnetic flux

of controlled amplitude and frequency in the stator teeth and back iron. The sensor is an

electromagnetic device comprising a magnetic core made of laser welded high quality steel

laminations, and a drive coil with a high current density and the requisite number of Ampere-turns.

The device was able to impose a large enough flux such that the flux density in at least some

regions of the part or piece of material being tested is well into the saturation region of the

material’s B-H characteristics. Testing was conducted on two groups of stators, and results are

compared between the two groups and to FEM computation of losses.

In order to obtain a good spatial picture of the steel, testing was done for each stator in

a number of locations, and in each location a measurements were taken with the sensor moving

slightly, in order to obtain an average. A set of data was collected for each of the following

conditions and test parameters: 5 distinct regions or sectors of each stator, 6 local within each

region, 2 flux density levels induced in the stator tooth: and a drive current frequency of 50 Hz.



– 80 –

Fig.1. Sensor and magnetic field during testing

Fig.2. Connection and placement Fig.3. Data acquisition system

The data were used to compute the total energy losses over four cycles (or equivalently, the

average power losses over the same period) for each set of data. Since only the (type of) stator

changed from one experiment to another and the sensor characteristics were common to all

experiments, it was possible to make a direct comparison between the power losses to determine

which (type of) stators exhibited greater overall losses, and therefore resulting in the conclusion

that there would have been correspondingly greater losses in the stator material in particular (as

opposed to the sensor laminations, which are a constant factor in all trials).

The electromagnetic probe and associated experimental procedure presented in this work

provide the ability to distinguish between stators made of the different types of material that were

tested.

REFERENCES

[1] Popescu M., Ionel D.M., A best-fit model of power losses in cold rolled-motor lamination steel

operating in a wide range of frequency and magnetization, IEEE Transactions on Magnetics, vol. 43,

no. 4, SI, 2007, pp. 1753-1756

[2] Cheng Z., Takahashi N., Forghani B., Du Y., Fan Y., Liu L., Zhao Z., Wang H., Effect of variation of

B-H properties on loss and flux inside silicon steel lamination, IEEE Transactions on Magnetics,

vol. 47, no. 5, 2011, pp. 1346-1349

[3] Rasilo P., Dlala E., Fonteyn K., Pippuri J., Belahcen A., Arkkio A., Model of laminated ferromagnetic

cores for loss prediction in electrical machines, IET Electric Power Applications, vol. 5, no. 7, 2011,

pp. 580-588.

[4] Mazurek R., Hamzehbahmani H., Moses A. J., Anderson P.I.,. Anayi F.J, Belgrand T., Effect of

artificial burrs on local power loss in a three-phase transformer core, IEEE Transactions on Magnetics,

vol. 48, no. 4, 2012, pp. 1653-1656

[5] Romary R., Jelassi S., Brudny J.F., Stator-interlaminar-fault detection using an external-flux-density

sensor, IEEE Transactions on Industrial Electronics, vol. 57, no. 1, 2010, pp. 237-243

[6] Gutierrez-Castaneda E.J., Salinas-Rodriguez A., Effect of annealing prior to cold rolling on magnetic

and mechanical properties of low carbon non-oriented electrical steels, Journal of Magnetism and

Magnetic Materials, vol. 323, no. 20, 2011, pp. 2524-2530

– 81 –

PARTICIPANTS OF



Bednarczyk Tomasz GUMET Sz. Geneja Spółka Jawna

ul. Kolejowa 12, 23-200 Kraśnik, Poland


Bieńkowski Adam Warsaw University of Technology

Institute of Metrology and Biomedical Engineering

ul. Andrzeja Boboli 8, 02-525 Warsaw, Poland


Chmiel Grzegorz GUMET Sz. Geneja Spółka Jawna

ul. Kolejowa 12, 23-200 Kraśnik, Poland


Chwastek Krzysztof Częstochowa University of Technology


Al. Armii Krajowej 17, 42-200 Częstochowa, Poland


de Campos Marco Flavio UFF- Federal Fluminense University

Av dos Trabalhadores 420, 27255-125 Volta Redonda RJ, Brazil


Demenko Andrzej Poznań University of Technology


Piotrowo 3A, 60-965 Poznań, Poland


Eichler Jakub Technical University of Liberec

Studentska 2, 46117 Liberec, Czech Republic


Gas Piotr AGH University of Science and Technology, Department of Electrical

and Power Engineering

Al. Adama Mickiewicza 30, 30-059 Kraków, Poland


Garstka Tomasz Częstochowa University of Technology

Faculty of Production Engineering and Materials Technology



Gozdur Roman Technical University of Lodz

Department of Semiconductor and Optoelectronic Devices

ul. Wólczańska 211/215, 90-924 Łódz, Poland









– 82 –

Guzowski Bartłomiej Technical University of Lodz

Department of Semiconductor and Optoelectronic Devices

ul. Wólczańska 211/215, 90-924 Łódz, Poland


Hameyer Kay RWTH Aachen University

Institute of Electrical Machines (IEM)

Schinkelstrasse 4, D-52062 Aachen, Germany


Haneczok Grzegorz University of Silesia, Institute of Materials Science

ul. 75 Pułku Piechoty 1A, 41-500 Chorzów, Poland


Jagiełło Adam Cracow University of Technology


ul. Warszawska 24, 31-155 Kraków, Poland


Jakubas Adam Częstochowa University of Technology




Jastrzębski Radosław Częstochowa University of Technology



Kapelski Dariusz Tele & Radio Research Institute

ul. Ratuszowa 11, 03-450 Warszawa, Poland


Kapłon Andrzej Kielce University of Technology

Power Electronic, Electrical Machines and Drives Chair

Aleja 1000-lecia Państwa Polskiego 7, 25-314 Kielce, Poland


Kluszczyński Krzysztof Cracow University of Technology




Koprivica Branko University of Kragujevac

Faculty of Technical Sciences in Cacak

Svetog Save 65, 32000 Cacak, Serbia


Kotynia Katarzyna Częstochowa University of Technology, Faculty of Production

Engineering and Materials Technology, Institute of Physics

al. Armii Krajowej 19, 42-200 Częstochowa, Poland




http://uranos.cto.us.edu.pl/~inom/index.php/instytut/kontakt






http://www.pcz.pl/english/faculites-and-other-units/materials-processing-technology-and-applied-physics




– 83 –

Kucal Ewelina Tele & Radio Research Institute



Kwapuliński Piotr University of Silesia, Institute of Materials Science

ul. 75 Pułku Piechoty 1A, 41-500 Chorzów, Poland


Leuning Nora RWTH Aachen University

Institute of Electrical Machines (IEM)

Schinkelstrasse 4, D-52062 Aachen, Germany


Lisowiec Aleksander Tele & Radio Research Institute



Łada-Tondyra Ewa Częstochowa University of Technology




Mazgaj Witold Cracow University of Technology




Mészáros István Budapest University of Technology and Economics

Department of Materials Science and Engineering

H-1111 Bertalan L. u. 7., Budapest, Hungary


Nadolski Roman Kielce University of Technology


Aleja 1000-lecia Panstwa Polskiego 7, 25-314 Kielce, Poland


Najgebauer Mariusz Częstochowa University of Technology




Novak Miroslav Technical University of Liberec

Studentská 2, 46117 Liberec, Czech Republic


Nowakowski Andrzej Tele & Radio Research Institute















– 84 –

Nowicki Michał Warsaw University of Technology


sw. Boboli 8, 02-525 Warsaw, Poland


Oźga Katarzyna Częstochowa University of Technology




Pinhasi Yosef Ariel University, Faculty of Engineering

Department of Electrical and Electronic Engineering

Kiriat Hamata POB 3, Ariel 40700, Israel


Pluta Wojciech A. Częstochowa University of Technology




Przybylski Marek Tele & Radio Research Institute



Przygodzki Jacek Retired Professor of Warsaw University of Technology


Rolek Jarosław Kielce University of Technology


Aleja 1000-lecia Państwa Polskiego 7, 25-314 Kielce, Poland


Sobczyk Tadeusz Cracow University of Technology




Strangas Elias G. Electrical Machines and Drives Laboratory

Department of Electrical and Computer Engineering

Michigan State University, East Lansing, MI 48824-1226 USA


Stachowiak Dorota Poznań University of Technology


Piotrowo 3A, 60-965 Poznan, Poland


Svec Peter Institute of Physics

Slovak Academy of Sciences

Dúbravská cesta 9, 845 11 Bratislava 45, Slovakia












– 85 –

Sykulski Jan University of Southampton

University Road, Southampton SO17 1BJ, United Kingdom


Szczurek Paweł Stalprodukt S.A.

ul. Wygoda 69, 32-700 Bochnia, Poland


Szczygłowski Jan Częstochowa University of Technology




Szewczyk Roman Warsaw University of Technology


ul. Andrzeja Boboli 8, 02-525 Warsaw, Poland


Ślusarek Barbara Tele & Radio Research Institute



Wołoszyn Mirosław Technical University of Gdańsk

Faculty of Electrical and Control Engineering

ul. G. Narutowicza 11/12, 80-233 Gdańsk, Poland


Yahalom Asher Ariel University, Faculty of Engineering

Department of Electrical and Electronic Engineering

Kiriat Hamata POB 3, Ariel 40700, Israel








– 86 –

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