Thermal Analysis and Management of High...

Thermal Analysis and Management of

High-Performance Electrical Machines

SHAFIGH NATEGH

Doctoral Thesis

Stockholm, Sweden 2013

TRITA-EE 2013:022ISSN 1653-5146ISBN 978-91-7501-733-4

KTH School of Electrical EngineeringSE-100 44 Stockholm

SWEDEN

Akademisk avhandling som med tillstånd av Kungl Tekniska högskolan framläggestill offentlig granskning för avläggande av teknologie doktorsexamen torsdag den13 juni 2013 klockan 14.00 i F3, Kungl Tekniska högskolan, Lindstedtsvägen 26,Stockholm.

© Shafigh Nategh, juni 2013

Tryck: Universitetsservice US AB

This thesis is dedicated to peace and freedom

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Abstract

This thesis deals with thermal management aspects of electric machineryused in high-performance applications with particular focus put on electricmachines designed for hybrid electric vehicle applications.

In the first part of this thesis, new thermal models of liquid (water andoil) cooled electric machines are proposed. The proposed thermal models arebased on a combination of lumped parameter (LP) and numerical methods.As a first case study, a permanent-magnet assisted synchronous reluctancemachine (PMaSRM) equipped with a housing water jacket is considered. Par-ticular focus is put on the stator winding and a thermal model is proposedthat divides the stator slot into a number of elliptical copper and impregna-tion layers. Additionally, an analysis, using results from a proposed simplifiedthermal finite element (FE) model representing only a single slot of the sta-tor and its corresponding end winding, is presented in which the number oflayers and the proper connection between the parts of the LP thermal modelrepresenting the end winding and the active part of winding are determined.The approach is attractive due to its simplicity and the fact that it closelymodels the actual temperature distribution for common slot geometries. Anoil-cooled induction machine where the oil is in direct contact with the statorlaminations is also considered. Here, a multi-segment structure is proposedthat divides the stator, winding and cooling system into a number of an-gular segments. Thereby, the circumferential temperature variation due tothe nonuniform distribution of the coolant in the cooling channels can bepredicted.

In the second part of this thesis, the thermal impact of using differentwinding impregnation and steel lamination materials is studied. Conven-tional varnish, epoxy and a silicone based thermally conductive impregnationmaterial are investigated and the resulting temperature distributions in threesmall induction machines are compared. The thermal impact of using differentsteel lamination materials is investigated by simulations using the developedthermal model of the water cooled PMaSRM. The differences in alloy con-tents and steel lamination thickness are studied separately and a comparisonbetween the produced iron losses and the resulting hot-spot temperatures ispresented.

Finally, FE-based approaches for estimating the induced magnet eddycurrent losses in the rotor of the considered PMaSRM are reviewed andcompared in the form of a case study based on simulations. A simplifiedthree-dimensional FE model and an analytical model, both combined withtime-domain 2D FE analysis, are shown to predict the induced eddy currentlosses with a relatively good accuracy compared to a complete 3D FE basedmodel. Hence, the two simplified approaches are promising which motivatesa possible future experimental verification.

Index Terms: Computational fluid dynamics, directly cooled electricmachines, finite element analysis, heat transfer, hybrid electric vehicle,induction machines, lumped-parameter thermal model, permanent-magnetassisted synchronous reluctance machines.

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Acknowledgements

This PhD thesis concludes the research that I have done at the Depart-ment of Electrical Energy Conversion, KTH Royal Institute of Technology,from September 2009 until June 2013.

First of all, I would like to thank my main supervisor and examiner Prof.Chandur Sadarangani for providing the opportunity to continue my academicstudies at the post-graduate level. I would also thank my supervisors As-soc. Prof. Oskar Wallmark and Lic. Eng. Mats Leksell for their guidance,supervision and encouragement throughout my PhD studies period.

I want to express my appreciation to all my current and former colleaguesat the Department of Electrical Energy Conversion (Electrical Machines andPower Electronics) where I spent four years of my life’s best time. Specialthanks go to my current and former officemates Andreas Krings, Hui Zhang,Ahmed Noman, Kashif Khan and Shuang Zhao for many nice and fruitfulconversations and also providing an excellent work environment.

I further would like to thank the Termoos project steering committee,Mr. Jan Folkhammar at BEVI AB, Mr. Viktor Lassila at BAE Systems,Prof. Mats Alaküla at LTH, Mr. Magnus Lindenmo at Surahammars Bruk,Mr. Svante Bylund formerly at BAE Systems, Lic. Eng. Zhe Huang atLTH, Assoc. Prof. Anders Malmquist at KTH, Dr. Lars Kvarnsjö at Vacu-umscmelze, Mr. Tom Sundelin at BAE Systems, Kristoffer Nilsson at Borg-Warner/Haldex Traction and Jerker Andersson at Dahréntråd, who guidedme with great comments and questions at the many meetings held during thisproject.

Prof. Stefan Östlund and Prof. Philip T. Krein are acknowledged forproviding the possibility for me to visit the University of Illinois at Urbana-Champaign during autumn 2012 and part of winter 2013. I would particularlythank the Magill family, V. Tutku Buyukdegirmenci, Hao Zhu, Siming Guo,Enver Candan, Joyce Mast, Giovanni Massa and Kai Van Horn for giving mea rewarding stay in the US.

I am very grateful to Eva Pettersson and Celie Geira for all of administra-tive aid, Peter Lönn for his help regarding my computers and software, andalso the E2C laboratory technicians Olle Brännvall and Jesper Freiberg.

I would also like to extend my thanks to my friends outside the departmentfor their support and kind help in different ways: Seyedali, Ebrahim, Afshin,Mohamad, Mohammadreza, Yelena, Angela, Ara, Amin, Majid, Reza, Royaand Saeid.

Last, but certainly not least, I would like to thank my lovely wife, Asiyeh,without whose love, continual support and patience my contribution to thisproject would not have been possible. Thanks also to my father, AbdolrezaNategh, for showing me what a hardworking man truly is, and my motherwho without her encouragement and endeavours I would never be able toachieve my goals.

Shafigh NateghStockholm, SwedenJune 2013

Contents

Contents ix

1 Introduction 1

1.1 Thermal Analysis and Management of Electric Machinery . . . . . . 11.1.1 Thermal Modeling of Electric Machines: Analytical Approaches 21.1.2 Thermal Modeling of Electric Machines: Numerical Methods 4

1.2 Objectives and Scope of Thesis . . . . . . . . . . . . . . . . . . . . . 51.3 Outline of Thesis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 71.4 Scientific Contributions . . . . . . . . . . . . . . . . . . . . . . . . . 81.5 List of Publications . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

1.5.1 Contributions of Individual Authors . . . . . . . . . . . . . . 11

2 Thermal Modeling of Liquid-Cooled Electric Machines 13

2.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 132.2 Thermal Modeling of Water-Cooled PM Machines . . . . . . . . . . 14

2.2.1 LP Thermal Model of the Winding . . . . . . . . . . . . . . . 142.2.2 LP Thermal Model of the Rotor . . . . . . . . . . . . . . . . 222.2.3 Complete LP Thermal Model . . . . . . . . . . . . . . . . . . 232.2.4 Experimental Evaluation . . . . . . . . . . . . . . . . . . . . 26

2.3 Thermal Modeling of Directly Cooled Electric Machines . . . . . . . 292.3.1 Thermal Modeling Approach . . . . . . . . . . . . . . . . . . 302.3.2 Complete LP Thermal Model . . . . . . . . . . . . . . . . . . 362.3.3 Experimental Evaluation . . . . . . . . . . . . . . . . . . . . 36

2.4 Summary of Chapter . . . . . . . . . . . . . . . . . . . . . . . . . . . 38

3 Thermal Effects – Material Selection 41

3.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 413.2 Thermal Effects of Using Different Impregnation Materials . . . . . . 42

3.2.1 Impregnation Materials Studied . . . . . . . . . . . . . . . . . 423.2.2 Impregnation Process Using SbTCM . . . . . . . . . . . . . . 433.2.3 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43

3.3 Thermal Effects of Using Different Steel Lamination Materials . . . 473.3.1 Comparison Between Different Steel Laminations . . . . . . . 49

ix

x CONTENTS

3.3.2 Iron Loss Estimation . . . . . . . . . . . . . . . . . . . . . . . 503.3.3 Thermal Impact of Using Laminations with Different Qualities 50

3.4 Summary of Chapter . . . . . . . . . . . . . . . . . . . . . . . . . . . 52

4 Magnet Eddy Current Loss Estimation 55

4.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 554.2 Complete 3D Electromagnetic FEM . . . . . . . . . . . . . . . . . . 564.3 2D and Partial 3D Electromagnetic FEA . . . . . . . . . . . . . . . . 574.4 2D Electromagnetic FEA and Analytical Modeling . . . . . . . . . . 594.5 Evaluation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 604.6 Summary of Chapter . . . . . . . . . . . . . . . . . . . . . . . . . . . 60

5 Concluding Remarks 63

5.1 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 635.2 Proposal for Future Work . . . . . . . . . . . . . . . . . . . . . . . . 64

List of Figures 67

List of Tables 71

Bibliography 73

A Glossary of Symbols and Abbreviations 81

B Selected Publications 85

Chapter 1

Introduction

Thermal aspects of electric machines, particularly in high-performance applications,are concisely presented in this chapter. Also, traditional thermal modeling ap-proaches are reviewed, and strengths and weaknesses of each method are discussed.Then, the scope of this thesis and the major scientific contributions are highlighted.Finally, the publications originating from this project are listed.

1.1 Thermal Analysis and Management of Electric

Machinery

In high-performance applications such as hybrid electric vehicles and aerospace,there is a growing need for electric machines with high torque/power densities.A higher power density can be achieved by applying higher current densities tothe electric machine windings and/or running the machine at higher speeds. Ahigh current density in the stator winding results in significant copper losses and,in turn, high hot-spot temperatures [1, 2]. Also, high rotor speeds lead to highercurrent and voltage frequencies that increase the iron losses in the stator and rotorsteel laminations and the permanent magnet segments in permanent magnet (PM)machines. The increase in copper and iron losses may, if the resulting heat isnot properly dissipated, cause increased temperatures which may be particularlyproblematic in parts of the machine that are difficult to cool down (e.g. the rotor).

From the beginning of the twentieth century, considerable efforts have beenput on dealing with thermal issues of electric machines [3–6]. In this regard, aconsiderable amount of work has been carried out on the development of complexcooling systems that effectively extract losses from critical parts of the machine[7–11]. Forced air cooling has been employed to enhance the heat transfer fromthe housing fins and often also from the end winding and rotor surfaces [12, 13].However, for high current densities, air cooling may not be sufficient and some formof liquid cooling is required [11]. A housing water jacket enables an effective heattransfer from the active part of the stator winding to the coolant [Paper I], [14,15].

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2 CHAPTER 1. INTRODUCTION

However, water cooling does not provide a successful cooling of the end windingswhich can be particularly problematic for machines with long end windings, e.g.in machines with low pole numbers. Instead of water, oil may also be used as acooling medium and different oil cooling methods for electric machinery have beenproposed [7,16–18]. In directly-oil-cooled machines, oil is in direct contact with theinner parts of the machine and an effective cooling of both the stator and the endwinding body can be realized.

A variety of work has been done and published on improvements in the electro-magnetic design of different kinds of electric machines to reduce losses, e.g. the ironlosses in the stator and rotor laminations, and eddy current losses in the permanentmagnets [19–22].

In order to improve the thermal behavior of electric machines, a good knowledgeof heat transfer in different parts of the machine is required. Lumped parameter(LP) thermal analysis and numerical methods are the major approaches proposedto model thermal effects in electric machines [23].

1.1.1 Thermal Modeling of Electric Machines: Analytical

Approaches

From the time that thermal issues of electric machines first raised, engineers havetried to develop analytical methods to estimate the temperature distribution inelectric machinery [24, 25]. An early attempt made to implement a functioninganalytical thermal model to predict the temperatures in different parts of an electricmachine is presented in [26, 27] where a simple thermal network is developed fortotally enclosed non-ventilated induction motors. Following this work, the thermalmodel of a totally enclosed fan cooled electric machine was developed and testedby Mellor et al. on a medium (75 kW) and two small size (5.5 kW) inductionmachines [28, 29]. In this work, thermal models of different parts of inductionmachines were derived. In [30], a thermal model of high-speed induction machines(20000 – 200000 rpm) is implemented and discussed. In [31], Lindström appliedanalytical thermal modeling to the PM machines. Following this work, Aglén etal. [32] employed thermal modeling for high-speed PM machines and EL-Refaieet al. [33] developed a LP thermal model for multi-barrier interior PM machines.During the last decade, many attempts have been made mainly by Staton, Boglietti,Cavagnino et al. to develop more accurate models by solving difficult aspects ofelectric machines thermal modeling and identifying thermal parameters which aredifficult to estimate [34–40]. Moreover, software packages, e.g. Motor-CAD1 andPortunus2, with the aim of simplifying thermal modeling of electric machine forusers with only a basic knowledge of heat transfer have been developed.

LP modeling is the most commonly used analytical approach to model thermaleffects in electric machines. A LP model represents a simplification, where spatially

1Motor-CAD is a registered trademark of Motor Design Ltd, Shropshire, U.K.2Portunus is a registered trademark of the Cedrat Group, Meylan Cedex, France.

1.1. THERMAL ANALYSIS AND MANAGEMENT OF ELECTRIC

MACHINERY 3

Figure 1.1: LP thermal model of a sample electric machine. The abbreviations usedare reported in Table 1.1.

distributed fields are approximated as a number of single scalars. Consequently,the lumped parameters can be configured according to heat transfer fundamentals,dimensional data of the machine studied and the thermal properties of the materialsused. A simple LP thermal model of a sample machine can be found in Figure 1.1where conduction is represented by dark blue blocks and convection and radiationare represented by light blue blocks.

Table 1.1: Abbreviations used in Figure 1.1.

(F) Forward Hs Housing(R) Rear Plt PlateCd Conduction Sht ShaftCv Convection Sl-Wl Slot wallRd Radiation St-Br Stator boreBrn Bearing Yk Stator yokeEc End cap Tth Stator toothEs End space Wnd WindingE-Wnd End winding

As seen in Figure 1.1, the LP thermal model of a complete electric machinecomprises of the analytical thermal models of different parts of the machine [13,


28, 31, 34, 35, 41–43]. Given the heat sources, e.g. winding (copper losses), statorand rotor steel laminations (iron losses), the thermal network can be solved andthe temperature distribution in different parts of the machine can be estimated.

Implementing and running LP thermal models is a fairly easy and quick task.However, developing accurate analytical models for complex parts of electric ma-chines, e.g. the winding and the rotor, is challenging. As a result, thermal modelsdeveloped based on only analytical modeling methods may not be adequate toprecisely model thermal effects of electric machines with complex structures.

1.1.2 Thermal Modeling of Electric Machines: Numerical

Methods

Finite element analysis (FEA) and computational fluid dynamics (CFD) are com-monly used numerical methods for thermal analysis of electric machines [23]. Ina thermal finite element (FE) model, conduction in solid elements with specifiedconductivities can be modeled accurately but convection and radiation must be ap-proximated with boundary conditions based on empirical correlations [41, 44–48].Therefore, FEA can provide a fairly accurate picture of temperature distribution incomplex solid parts of the electric machine, e.g. the winding, provided the boundaryconditions are introduced accurately. The thermal FE model of the stationary partof a sample electric machine, implemented using the software JMAG3, is shownin Figure 1.2 [Paper IX]. As can be seen, the solid elements including the sta-tor steel laminations (grey), conductors (yellow), liner4 (dark blue), impregnation(light blue) and end winding ring (green) are implemented in the FE software andconvective heat transfers are modeled using the boundary conditions.

Thermal FE modeling has been applied to different electric machine topologies.A thermal FE model of induction motors is presented in [49]. In [50], thermal FEmodeling is applied to PM synchronous motors and a comparison between thermalFEA and LP modeling is presented. Also, a thermal FE model of high-speed PMmachines is developed in [51].

CFD analysis can be used to accurately model convective heat transfer and alsofluid flows of the coolant in the electric machines equipped with different kinds ofcooling systems [23, 52]. However, implementing and solving CFD and/or thermalFE models of electric machines can be very time-consuming [23, 53, 54]. There-fore, using these numerical methods for example in design optimization procedures,where many design iterations are commonly required, is not recommended.

3JMAG is a registered trademark of the JSOL Corporation, Tokyo, Japan.4Liner is a paper material insulating the stator laminations from the stator winding. It is also

called slot insulation.

1.2. OBJECTIVES AND SCOPE OF THESIS 5

88.5

90.2

91.9

93.6

95.3

97.0

a)

b)

Figure 1.2: Thermal FEA of the stationary part of a sample electric machine: a)FE model; b) FEA results.

1.2 Objectives and Scope of Thesis

The task of this PhD project5 was to develop a deepened knowledge in thermaldesign of electric machinery, and to find out how thermal characteristics of materialsand physical design work together in order to optimize performance or to minimizecost, volume and weight of the electric machines and their cooling systems to agiven performance. The research focuses on existing challenges and problems inthe ongoing product development for vehicle applications. It should be noted,however, that the outcome of this research can also be applied to electric machines

5The research presented in this thesis is part of a nationally financed research project namedTERMOOS, Thermal Design of Electric Machines for Hybrid Drive Lines. The project partnersare BAE Systems Hägglunds, KTH, BorgWarner TorqTransfer System AB (Haldex Traction AB),Lunds Tekniska Högskola (LTH), Bevi AB, Dahréntråd AB, MagComp, Surahammars Bruk ABand Vacuumschmelze AB.


designed for the other applications, e.g. aircraft or electric railway traction.As a first step, thermal modeling methods for liquid-cooled electric machines

have been considered. In this regard, the two most commonly used liquid coolants,water and oil, were studied. Since water cannot be in direct contact of the innerparts of electric machines, e.g. the stator and rotor steel laminations, it is normallyapplied to electric machines using a housing water jacket. Using oil as coolantprovides the possibility of direct cooling of the inner parts of the machine wherehot-spot temperatures rise, e.g. the end winding. As pointed out in Section 1.1.1and Section 1.1.2, each particular modeling approach has its own strengths andweaknesses. On the one hand, the results obtained from the analytical approachdo not have a high accuracy and are also very much dependent on the assumptionsmade to simplify the complex structure of machine parts, but on the other hand,numerical methods are very time-consuming.

In water-cooled electric machines, the coolant is trapped in the housing jacketand flows in controlled paths. Therefore, the heat transfer to the coolant can bemodeled with a good accuracy using available empirical correlations expressing fluidflow and heat transfer in the coolant pipes with different cross sectional geometriesand lengths. However, the hot-spot temperature estimation in complex machinestructures including the stator slots and the end winding bodies is challenging.

In [31, 55], the stator winding is modeled by considering the impregnation andconductors in a stator slot as a homogeneous body with an equivalent thermal con-ductivity. In [42], two analytical expressions for determining the equivalent thermalconductivity of a stator slot are evaluated using thermal FEA and it is shown thatthe analytical techniques may risk largely underestimating the hot-spot windingtemperature when the fill factor is low (below 0.3). In addition, the temperaturevariation in the axial direction, which can be substantial (hot-spot temperaturesoften appear in the end windings), is not considered. The commercial softwareMotor-CAD uses a LP thermal modeling approach to determine the temperaturedistribution in electric machines. In Motor-CAD, the winding is modeled as a num-ber of layers of copper, wire insulation, and inter-conductor insulation [34, 35, 56].However, the achieved accuracy of the results depends on several input parameters(provided by the software user) which may have a substantial influence on the esti-mated end winding temperature and the hot-spot temperature of the active part ofthe winding. It should be noted that the temperature of the winding has a directinfluence on the lifetime of the used insulation materials, and an accurate predictionof the temperature distribution in the winding enables the lifetime of the machineto be estimated [57]. Furthermore, without an accurate knowledge of the windingtemperature distribution, an effective cooling system design is not possible.

In oil cooled machines, the fluid flow distribution which is in direct contact ofthe inner parts of the machine, is relatively difficult to predict. However, preciseinformation of the fluid velocity is necessary for an accurate temperature predictionand available analytical expressions are generally not able to model fluid flow incomplex structures of oil cooled electric machines. One approach is to develop CFDmodels of the complete machine as presented in [58, 59]. However, as previously

1.3. OUTLINE OF THESIS 7

mentioned, this method is very time-consuming and hardly suitable for example ina design process where a considerable number of iterations is needed.

In this thesis, the aforementioned modeling issues are circumvented by combin-ing analytical and numerical methods to take advantages of both approaches. Thisenables accurate and relatively simple thermal models provided that the modelingis done with care.

As highlighted in Section 1.1, the attempts made to reduce the hot-spot tem-perature in critical parts of the electric machines can be classified in the followingtwo tracks:

1. High-performance and complex cooling system designs that effectively extractlosses from the critical parts of the electric machines6.

2. Improvements in the electromagnetic design to reduce the produced losses indifferent parts of the machine.

Another parameter that plays a vital role in the thermal behavior of electric ma-chinery, is thermal and loss characteristics of the materials used. Another objectiveof this PhD project was to investigate thermal effects of using different materialsin the electric machines. The materials considered are the impregnation materialfilling the stator slots and end winding bodies and the steel lamination materialsused to manufacture the stator and rotor.

Needless to say, losses as the input of the thermal models play an important rolein thermal management of electric machines. Since PM machines are commonlyused for high-performance applications, an accurate knowledge of losses producedin this kind of electric machine is necessary. Therefore, induced eddy currents inthe rotor magnets of PM machines is also studied in this work.

1.3 Outline of Thesis

This PhD thesis is organized in the form of compilation thesis, so that the chaptersare kept brief and the scientific contributions are further presented in the appendedpapers. The chapters provide the scientific background, introduce key concepts andpresent important simulation and experimental results.

The thesis is outlined as follows.Chapter 2 focuses on thermal modeling. In the first part of this chapter, a ther-

mal modeling approach for a water-cooled permanent-magnet assisted synchronousreluctance machine (PMaSRM) is presented. The thermal models of two criticalparts of the machine, the winding and the rotor, are detailed and thermal modelsof the remaining parts are reviewed. In the second part of the chapter, a thermalmodel for directly cooled electric machines is derived. The main focus is put on thethermal model of the winding and the complex cooling system structure.

6The research done at LTH, as one of the main TERMOOS project partners, was focused ondeveloping effective cooling systems.


In Chapter 3, thermal effects of the slot impregnation materials and steellaminations are investigated. To study thermal effects of impregnation materials,both simulation studies and corresponding experimental tests are carried out andthe results are presented. However, the study on the thermal impact of differentsteel lamination qualities relies solely on simulation results using the experimentallyverified thermal model presented in the Chapter 2.

Chapter 4 presents three FE-based approaches to model eddy currents andtheir associated losses in permanent magnet segments of PM machines. In the firstpart of this chapter, a classic approach based on a 3D FE electromagnetic model ofthe complete machine, is presented. In the second and third parts of this chapter,with the aim to reduce the total computation time, 2D electromagnetic FEA iscombined with analytical and partial 3D FE models and the results are comparedwith the complete 3D FEA of the machine.

Chapter 5 summarizes the conclusions of the work and provides some proposalsfor further research.

1.4 Scientific Contributions

The main contribution presented in this thesis is summarized in the following list.

• Thermal effects in a water-cooled PMaSRM is studied and modeled. A newapproach to model heat transfer in the winding, including the stator slots andend windings, is presented [Paper I]. In this method, advantages of both LPmodeling and FEA are exploited. Particularly, partial FEA is used to developa LP thermal model of the winding. The proposed partial FE model of thewinding is simple and can be implemented with relative ease. To model thetemperature variation from the innermost part of the slot to the slot wall,a LP thermal model is proposed which is based on dividing the slot into anumber of elliptical layers. This approach is attractive due to its simplicityand the fact that it closely models the actual temperature distribution fordifferent stator slot geometries. Additionally, as presented in [Paper II], twoLP thermal models of the rotor of a PMaSRM are implemented and compared.

• A thermal model for directly cooled electric machines is presented. Focus isput on critical parts of the machine including the stator slots, the end windingbody and the cooling system. Due to the non-linear distribution of the coolantin the cooling channels and on the outer surface of the end winding bodies,and also because of the existence of gathered oil at the bottom of the machine,the temperature distribution in different stator slots varies circumferentially.Therefore, a multi-segment structure is proposed that divides the machineinto a number of angular sections, in the circumferential direction. Moreover,a partial CFD analysis is adopted to estimate the coolant distribution in thecooling channels and on the outer surfaces of the end windings. As presentedin [Paper III], the proposed approach enables thermal modeling of directly

1.5. LIST OF PUBLICATIONS 9

cooled electric machines with good accuracy with only a limited need fortime consuming CFD simulations.

• Thermal impact of using two available impregnation materials and also analternative material with a higher thermal conductivity is studied. The twoavailable impregnation materials are varnish and Epoxylite7 with thermalconductivities of ≈ 0.25 W/mK and ≈ 0.85 W/mK, respectively. The al-ternative impregnation material studied in this research is a silicone basedthermally conductive material (SbTCM8) with the thermal conductivity of3.2 W/mK. As presented in [Paper IV], three identical machines using theaforementioned materials are manufactured and tested at different currentamplitudes and cooling conditions.

• The thermal and electromagnetic effects of using different steel laminationqualities in electric machines are investigated. As presented in [Paper V],varying the thickness and amount of alloy contents of seven commonly usedsteel laminations are studied on a PMaSRM. The lamination thickness andamount of alloy contents are the main parameters characterizing the thermal,mechanical and electromagnetic properties of the steel laminations materials.

• FE-based approaches for computing eddy current losses in permanent magnetsegments of PM machines are presented and compared.

1.5 List of Publications

The following is a list of publications in which the author has contributed to duringthis PhD project.

I S. Nategh, O. Wallmark, M. Leksell, and S. Zhao, “Thermal analysis of aPMaSRM using partial FEA and lumped parameter modeling,” IEEE Trans-actions on Energy Conversion, vol. 27, no. 2, pp. 477-488, June 2012.

II S. Nategh, O. Wallmark, and M. Leksell, “Thermal analysis of permanent-magnet synchronous reluctance machines,” in Proc. 14th European Conferenceon Power Electronics and Applications (EPE), Aug. 30-Sept. 1, 2011.

III S. Nategh, Z. Huang, O. Wallmark, M. Leksell, and A. Krings, “Thermalmodeling of directly cooled electric machines using lumped parameter andlimited CFD analysis,” submitted to the IEEE Transactions on Energy Con-version.

7Epoxylite is a registered trademark by Elantas PDG, INC. in St. Louis, MO 63147 USA.8Thermoset SC-320 Thermally Conductive Silicone Encapsulant, LORD Corporation, NC

27511-7923 USA.


IV S. Nategh, A. Krings, O. Wallmark, and M. Leksell, “Evaluation of impreg-nation materials for thermal management of liquid-cooled electric machines,”submitted to the IEEE Transactions on Industrial Electronics.

V S. Nategh, A. Krings, Z. Huang, O. Wallmark, M. Leksell, and M. Lindenmo,“Evaluation of stator and rotor lamination materials for thermal managementof a PMaSRM,” in Proc. XXth International Conference on Electrical Ma-chines (ICEM), 2-5 September, 2012.

VI S. Nategh, V. Lassila, Z. Huang, O. Wallmark, and M. Leksell, “Evaluationof alternative stator and rotor lamination materials for thermal managementof permanent magnet machines”, in Proc. IEEE International Magnetics Con-ference (INTERMAG), Taipei, 25-29 April, 2011.

VII A. Krings, S. Nategh, O. Wallmark, and J. Soulard, “Influence of the weldingprocess on the performance of slot-less PM motors with SiFi and NiFI statorlamination,” accepted for publication in IEEE Transactions on Industry Ap-plications.

VIII V. T. Buyukdegirmenci, S. Nategh, M. P. Magill, and P. T. Krein, “A fastand flexible analytical approach for thermal modeling of a linear stator struc-ture,” in Proc. IEEE International Electric Machines and Drives Conference(IEMDC), 12-15 May, 2013.

IX V. T. Buyukdegirmenci, M. P. Magill, S. Nategh, and P. T. Krein, “Devel-opment of closed-form solutions for fast thermal modeling of rotating electricmachinery,” in Proc. IEEE International Electric Machines and Drives Con-ference (IEMDC), 12-15 May, 2013.

X Z. Huang, S. Nategh, M. Alakula, V. Lassila, and Y. Jinliang “Direct oilcooling of traction motors in hybrid drives,” in Proc. IEEE InternationalElectric Vehicle Conference (IEVC), 4-8 March, 2012.

XI A. Krings, S. Nategh, O. Wallmark, and J. Soulard “Influence of the weld-ing process on the magnetic properties of a slot-less permanent magnet syn-chronous machine stator core,” in Proc. XXth International Conference onElectrical Machines (ICEM), 2-5 September, 2012.

XII A. Krings, S. Nategh, O. Wallmark, and J. Soulard “Local iron loss identi-fication by thermal measurements on an outer-rotor permanent magnet syn-chronous machine,” in Proc. 15th International Conference on Electrical Ma-chines and Systems (ICEMS), 21-24 October, 2012.

XIII A. Krings, S. Nategh, A. Stening, H. Grop, O. Wallmark, and J. Soulard“Measurement and modeling of iron losses in electrical machines,” invited pa-per, International Conference Magnetism and Metallurgy (WMM), 20-22 June,2012.

1.5. LIST OF PUBLICATIONS 11

1.5.1 Contributions of Individual Authors

The major parts of [Paper I]-[Paper VI], including the fundamental concepts, dis-cussions, simulation studies, experimental tests, and the manuscript are providedby the author of this thesis. The co-authors have contributed with helpful com-ments on the scientific contents, results and papers outline, as well as proofreading.It should be mentioned that in [Paper III], Zhe Huang from LTH has contributedwith the CFD simulations. In [Paper V], Magnus Lindenmo from Cogent PowerLtd. provided the required data of the studied steel laminations.

[Paper VIII] and [Paper IX] are the output of the research carried out duringa short-time study visit at the University of Illinois at Urbana-Champaign. Theauthor contributed actively in all parts of this research.

In [Paper X], the author contributed in developing the main idea of this paper,and also made contributions to Section I, Section III and Section IV of the paper.

In [Paper XIII], the author has written Section IV and contributed in the othersections. The author has made minor contributions in [Paper VII], [Paper XI] and[Paper XII].

Chapter 2

Thermal Modeling of

Liquid-Cooled Electric Machines

In this chapter, thermal modeling methods for liquid-cooled electric machines arepresented. The proposed modeling approaches are evaluated on a water-cooled PMa-SRM designed for hybrid electric vehicles and on two directly-oil-cooled inductionmachines manufactured using different impregnation materials filling the stator slotsand end winding bodies.

2.1 Introduction

As discussed in Chapter 1, LP modeling, which is an analytical approach, and nu-merical methods (FEM and CFD) are the major approaches used to model thermaleffects in the electric machinery. In this chapter, advantages of both LP model-ing and numerical methods are exploited in order to model thermal behavior ofdirectly- and indirectly-liquid-cooled electric machines. For indirect liquid cooling,a water-cooled PMaSRM designed for hybrid electric vehicle applications is mod-eled. Additionally, a direct cooling system implemented on two induction machinesis set up and analyzed.

In the water-cooled PMaSRM, the coolant is trapped in the channels insidethe housing and flows in the paths designed. Thereby, the cooling system can bethermally modeled using available empirical expressions modeling the convectiveheat transfer in the cooling channels. In this work, particular focus is put on thewinding and the rotor which are the critical parts from a thermal point of view[Paper I]. A thermal model for the winding is proposed that divides the stator slotinto a number of elliptical copper and impregnation layers. Additionally, resultsfrom a proposed simplified thermal FE model representing only a single slot of thestator and its corresponding end winding part are used to determine the parametersof the developed winding analytical model.

13

14CHAPTER 2. THERMAL MODELING OF LIQUID-COOLED ELECTRIC

MACHINES

Figure 2.1: Topology of the PMaSRM in consideration.

In directly cooled machines, the coolant is in direct contact with the inner partsof the electric machine, e.g. the stator back and the end winding. Such a coolingsystem potentially enables an effective cooling of both the stator back and the endwinding body. In the directly cooled machine considered in this work, the coolantdistribution in the cooling channels and on the end winding surfaces are difficultto predict analytically. Consequently, CFD simulations are needed to model thecoolant flow accurately. The results obtained from a simplified CFD simulationmodel are then used in the resulting LP thermal model. Additionally, to model theactual temperature distribution in the different stator slots and the end windingregions, a multi-segment structure is proposed that divides the machine into anumber of angular sections.

The developed thermal modeling technique for the winding of the water-cooledPMaSRM is a general approach that can be applied to the winding of other sorts ofelectric machines. Also, the proposed thermal modeling approach for the directly-oil-cooled induction machines can be applied to different types of directly cooledelectric machinery, e.g. synchronous reluctance and permanent magnet machines.

2.2 Thermal Modeling of Water-Cooled PM Machines

The geometry and general data of the studied PMaSRM are reported in Figure 2.1and Table 2.1, respectively [60, 61].

2.2.1 LP Thermal Model of the Winding

As mentioned in Chapter 1, the temperature of the winding plays an importantrole in the reliability and durability of the manufactured electric machines. Each10C increase in the winding temperature results in approximately a 50% decreasein the life time of the insulation material used and, consequently, the life time of the

2.2. THERMAL MODELING OF WATER-COOLED PM MACHINES 15

Table 2.1: PMaSRM data.

Rated power 19 kWRated speed 1500 rpmDC link voltage 400 VPole pairs 2 –Slots per pole per phase 2 –Cooling type Water jacket –Stator outer diameter 180 mmRotor diameter 112 mmAir-gap length 0.35 mmShaft diameter 52 mmActive length 210 mmMagnet segment axial thickness 10 mm

machine. Additionally, the temperature of the winding directly affects the efficiencyof the machine. Also, the fact that the hot-spot temperature normally occurs in thewinding highlights the importance of an accurate winding temperature prediction,especially in applications where a high power/torque density is demanded.

Among the proposed thermal models for the winding, the approaches basedon homogenization or the application of a multi-layer structure have attracted alot of interest. In [42], Idoughi et al. shows that modeling the stator slots asa homogenous body may risk underestimating the winding hot-spot temperatureespecially in machines with a low fill factor. The multi-layer configuration for thewinding suggested by Staton et al. in [35] 1 is able to estimate the temperaturedistribution in the winding provided that the model parameters, e.g. the numberof layers and connection point between the active part of the winding and the endwinding, are selected properly.

In this thesis, the multi-layer modeling approach for the winding is chosen andto decide on the critical parameters of the thermal model including those mentionedabove, thermal FEA is employed. The proposed partial FE model of the windingis simple and can be implemented with relative ease. The developed analyticalthermal model is based on dividing the slot into a number of elliptical layers andthe parameters of the proposed thermal model, representing the active part ofwinding and end winding, are calculated using an iterative method. The approachis attractive due to its simplicity and the fact that it closely models the actualtemperature distribution for several common stator slot geometries. Additionally,a FE modeling of the complete machine is not needed.

1The winding thermal model used in the software Motor-CAD is based on this winding modelstructure.


MACHINES

Partial Thermal FEA

In order to derive an accurate multi-layer LP thermal model of the winding, arealistic picture of the temperature distribution in the active part and in the endwinding is required. Since the heat is transferred mainly in the axial direction(along the z-axis) in the end winding and in the corresponding radial and tangentialdirections in the active winding, a 2D thermal FEA is not suitable and 3D thermalFEA has to be adopted. The modeling steps followed can be described as follows:

1. Model a single slot in 2D including conductors, impregnation, wedges, slotdivider, liners and stator steel laminations.

2. Extrude all parts in the axial direction by a height equal to the active lengthof the machine.

3. Extrude all parts, but the stator laminations, axially by a height equal to thedistance that the liners and wedges are extended out of the active part of themachine.

4. Extrude the conductors and impregnation by a height approximately equal tothe distance from the wedges and liners to where the winding starts to bend.

5. Model the end winding as a solid body representing the copper and impreg-nation.

6. Model the convective heat transfer from the end winding surfaces to the endspace air and from the stator surfaces to the air-gap and through housingto the ambient using appropriate heat transfer boundary conditions. Thethermal model used to model the heat transfer to the air-gap is reviewed inAppendix A of [Paper I].

7. Implement the losses in the winding and stator as heat sources.

Figure 2.2 shows the 3D thermal FE model used to model the winding. The modelis implemented using the FE software JMAG.

Iterative Approach

As shown in Figure 2.3, to model the heat transfer in the active part of the winding(stator slots), the isothermal surfaces in the stator slots are modeled as a number ofconcentric ellipses. The inner parts of the slot consist of copper and impregnationlayers, indicated by yellow and light blue colors, and the outer parts include anelliptical layer corresponding to the liner and a layer modeling the air pocketsbetween the liner and steel laminations. The average thickness of the air layerdepends on the quality of the manufacturing process. In this work, the assumedthickness of the air layer was 0.05 mm.

Based on the elliptical configuration of the stator slot presented in Figure 2.3, theLP thermal model of the winding including the active part and the end winding can


Step 1

Step 2-4

Step 5

Figure 2.2: 3D thermal FE winding modeling steps. Light green: end winding ring,dark green: slot wedge, yellow: conductors, dark blue: slot liner, light blue: slotimpregnation.

now be developed. As shown in Figure 2.4, the thermal model comprises of n layersmodeling the heat transfer in the axial direction of the winding’s active part andm copper layers surrounded by the impregnation in the radial direction. Since thethermal conductivity of copper is significantly higher than the thermal conductivityof the impregnation material, the thermal resistances of the copper layers in theradial direction can be neglected. However, the heat transfer in the axial directionthrough the impregnation material is disregarded and only the axial heat transferthrough the copper conductors is included. As shown in Figure 2.4, the heat transfer


MACHINES

Copper

Air pockets

Impregnation

Liner

Figure 2.3: Proposed elliptical model of the active winding. The colors are inagreement with the colors used in Figure 2.2.

between the end winding and stator slots is also taken into consideration. Sincethe hot-spot temperature is normally located in the end winding, particularly inmachines with a low number of poles, including the end winding in the thermalmodel of the winding is necessary to estimate the hot-spot temperature of thewinding with a high accuracy.

To limit the complexity of the model, both m and n should be selected as smallas possible while the accuracy is kept in a reasonable range. The selections of m andn are in this work based on results from the thermal FEA described in Section 2.2.1.

Remark: More information about the optimum values for n and m, and alsothe connection point between the active part of the winding and the end windingis provided in Section III-B3 in [Paper I].

As shown in Figure 2.3, the minor and major radii of the m ellipses are denotedby ri and r′

i where i=1, 2, . . . , m. In addition, the ellipses are selected so that

r′

1 − r1 = r′

2 − r2 = · · · = r′

m − rm. (2.1)

From (2.1), the minor radius of the ith copper layer can be expressed as

ri = rm − r′

m + r′

i. (2.2)


Figure 2.4: Proposed LP thermal model of the winding with n axial layers and melliptical copper layers in the slot. The selected colors are in agreement with thecolors used in Figure 2.2 and Figure 2.3.

The area enclosed by the ith ellipse is given by πrir′

i and the (constant) windingstrand density σStr in the slot can be expressed as

σStr =N1

πr1r′

1

=N1 + N2

πr2r′

2

= · · · =

∑mj=1 Nj

πrmr′

m

=NStr

πrmr′

m

, (2.3)


MACHINES

where NStr is the total number of strands in a slot. Outside each ellipse makingup a copper layer is an impregnation layer with the constant thickness timp (seeFigure 2.3). The major radius of the outermost ellipse r′

m is approximated fromthe slot height hSlot as

hSlot = 2 (r′

m + tLnr + tAG + tImp) , (2.4)

where tLnr is the thickness of the slot liner and tAG the equivalent air-gap lengthbetween the liner and the slot wall. Now, by combining (2.2) and (2.3), it is foundthat the major radius of the ith copper layer ri is related to rm and r′

m as

∑ij=1 Nj

π (rm − r′

m + r′

i) r′

i

= σStr. (2.5)

To determine ri and r′

i, the value of r′

m is first looped in an iterative manner. Foreach r′

m, rm is also looped iteratively. Then, (2.2) and (2.5) are used to computeapproximations of ri and r′

i. The (known) value of the slot area ASlot is nowused to determine the thickness of the impregnation layer tImp since ASlot can beapproximated from

ASlot = π (rm + tLnr + tAG + tImp) · (r′

m + tLnr + tAG + tImp) (2.6)

and tAG and tLnr are known. With knowledge of the winding fill factor, ASlot, theslot wedge area, and liner thickness, a known value of the total impregnation areain the slot AImp can be determined. For the elliptical model described above, thetotal impregnation area is approximated as

AImp =

m∑

i=1

π (r′

i + tImp) · (ri + tImp) − πr′

iri

=m

∑

i=1

π(

ritImp + r′

itImp + t2Imp

)

. (2.7)

When the calculated value of AImp corresponds to the (known) slot impregnationarea, the iteration of rm is stopped and hSlot is computed using (2.4). When thecalculated value of hSlot corresponds to the (known) value of the slot height, theiteration of r′

m is also stopped.The flow chart for the above process is shown in Figure 2.5. Here, L1 and L2 arethe first guesses for r′

m and rm, respectively.

As mentioned before, due to the high thermal conductivity of copper, the ther-mal resistance of each elliptical copper layer is neglected and only the thermalresistances of the insulation/impregnation layers in the radial direction are consid-ered. The thermal resistance in the jth axial layer of the insulation/impregnation


Figure 2.5: Iterative approach for determining winding model parameters.

between copper layer i and copper layer i + 1 can be approximated as

RW,j,i =n

2πLAλImpNSlot

ln[

4(ri + tImp) + 2∆r

+ 4√

(ri + tImp)2 + ∆r(ri + tImp) + ∆r2/2

]

− ln

[

4ri + 2∆r + 4√

r2i + ∆rri + ∆r2/2

]

, (2.8)

where LA is the active machine length, NSlot the total number of slots, and ∆r =r′

1 − r1 =r′

2 − r2 = . . .=r′

m − rm.Remark: Equation (2.8) that represents the thermal resistance of an elliptical

cylinder, is derived in the Appendix B of [Paper I].


MACHINES

In the axial direction, the thermal resistance of the ith copper layer RAW,i iscalculated as

RAW,i =LA

λCuAStrNin, (2.9)

where AStr is the cross sectional area of a conductor strand.The thermal resistances between the forward (F) and rear (R) nodes of the end

winding and the active winding REW,F and REW,R are calculated as

REW,F =LEW,a,F

λCuNSlotAStrNStr

(2.10)

and

REW,R =LEW,a,R

λCuNSlotAStrNStr

, (2.11)

where LEW,a,F and LEW,a,R are the lengths of the conductors that connect thestator slot to the end winding ring in the forward (drive) and rear (non-drive) endwinding coils, respectively.

It should be noted that in the water-cooled machine studied, heat transfer ismostly through the active part of the winding to the coolant flowing in coolingducts. Therefore, the convective heat transfer from the end winding to the endspace air does not play an important role in the thermal behavior of the electricmachine particularly at low rotor speeds. Consequently, in the developed thermalmodel of the winding focus is put on the active part of the winding and part of theend winding the connects the winding active part to the end winding ring. In thenext case study presented in Section 2.3, the heat transfer from the end windingto the coolant flowing on the end winding surface has a significant influence onthe winding temperature distribution. Thereby, thermal model of the end windingis further discussed and developed. The same argument can be made for the aircooled electric machines where the convective heat transfer to the end space air cansignificantly affect temperature distribution in the electric machine.

2.2.2 LP Thermal Model of the Rotor

Due to the risk of permanent demagnetization of the rotor magnets if too hightemperatures are reached, an accurate knowledge of the temperature distributionin the rotor of PM machines is necessary. In addition, the temperature of thepermanent magnets has a direct influence on the performance of the machine sincetheir remanent flux density decreases with temperature. For the rotor, the heatflow in the axial direction is considered and the two-dimensional heat transfers inthe x- and y-directions are assumed to be independent. Such a three-dimensionalLP thermal rotor model can be simplified further by assuming that the directionof the heat flow in the rotor is radial only [33].


Remark: Different configurations for the LP thermal model of the rotor arediscussed in [Paper II].

Such an approximate radial geometry and the corresponding LP thermal modelare illustrated in Figure 2.6 where the thermal resistance of each (radially conduc-tive) block consists of several cylindrical sections with a specific angular span. Thethermal resistance of a cylindrical section RCyl with thermal conductivity λ and anangular span δSpan can be expressed as

RCyl =ln(rEx/rIn)

λLCylδSpan

, (2.12)

where rIn and rEx are the inner and outer radii of the cylinder and LCyl =LA/n isthe axial length of the cylindrical section [33, 45].

The eddy current losses in each magnet and the iron losses in the rotor lami-nations are both injected into the nodes representing the magnets as illustrated inFigure 2.6.

Block 1

Block 2

Block 3

Magnet 1

Magnet 2

Magnet 3

Block 4

Magnet 4

Block 5

Air gap

Stator Bore

Block 1

Block 2

Block 3Block 4

Block 5Magnet 1

Magnet 2

Magnet 3

Magnet 4

Figure 2.6: Radial LP thermal rotor model. Loss sources are indicated as browndots.

2.2.3 Complete LP Thermal Model

Using the LP thermal models of the different parts of the PMaSRM described in theprevious sections and below, the complete LP thermal model of the machine is firstimplemented, and then solved using the software Portunus. To model the convectiveheat transfer from the stator housing to the coolant, the analytical formulationsin [47, 62] and [63] are used. Also, the convective heat transfers from the end


MACHINES

winding, shaft, and rotor to the end space are modeled analytically [43]. Exceptfor the stator winding and the rotor, the LP thermal models of the remaining partsof the PMaSRM can be found in [28–31,55].

The complete thermal model, including the housing water jacket, is illustratedin Figure 2.7 where the thermal resistances derived using the references listed aboveare illustrated as gray blocks. The yellow and green blocks represent the active partof the winding and part of the end winding, respectively. The corresponding mod-els are presented in Section 2.2.1. The blue blocks represent the thermal model ofthe rotor which is reviewed in Section 2.2.2. In addition to the thermal resistancesthat can be calculated using analytical and numerical methods, there are a fewthermal resistances that are difficult to estimate without experimental data avail-able for model calibration. These resistances are visualized as orange blocks andrepresent the contact resistance between the stator back and stator housing, thecontact resistance between the rotor yoke and the shaft, and the thermal resistancerepresenting the heat transfer between the shaft and bearing. The exact values ofthese thermal resistances depend on several factors, e.g. the manufacturing process,the size of machine and the used materials. In this thesis, these resistances are firstpre-estimated according to [34] and then calibrated using data from experimentaltests.

It should be noted, however, that the contact resistance between the rotor yokeand the shaft as well as the thermal resistance representing the heat transfer be-tween the shaft and the bearings do not play an important role in the hot-spottemperature determination. Also, as presented in [34, 56], the interface gap be-tween stator steel laminations and housing varies mainly from 0.03 mm to 0.07 mmfor the electric machines with different sizes and the materials used. The estimatedvalue for the equivalent thickness of the interface gap between the stator steel lam-inations and housing for the studied machine is 0.048 mm which is in the specifiedrange. Therefore, even without any calibration and just using the average valuesfor this parameter results in estimating temperature distribution in the studiedmachine with a reasonably high accuracy.

Table 2.2: Abbreviations used in Figure 2.7.

(F) Forward Hs Housing(R) Rear Plt PlateCd Conduction Sht ShaftCv Convection Sl-Wl Slot wallRd Radiation St-Br Stator boreBrn Bearing Yk Stator yokeEc End cap Tth Stator toothEs End space Wnd WindingE-Wnd End winding


Am

bie

nt

Co

ola

nt

Hs

Hs(

R)

Hs(

F)

Ec(

R)

Ec(

F)

Es(

R)

Sh

t(R

)

Es(

R)

Es(

F)

Sh

t(F

)

Es(

F)

Brn

(F)

Brn

(R)

Sh

t

Plt

Yk

Tth St-

Br

Sl-

Wl

(n) W

nd

E-W

nd

(F)

E-W

nd

(R)

Hs

Sh

t

Yk

Tth

St-

Br

Sl-

Wl

(1)

Wn

d

Rd

-Cv

Cd

Cd

Cd

Cv

Cd

Cd

Cv

Cd

Cd

-Cv

Rd-C

vC

dC

dC

d

Cd

Cd

Cv

Cd

Cd

Cv

Cd

Cv-C

d

Rd

-Cv

Cd

Cd

Cd

Cd

Cd

Cd

Cd

Cv

-Cd

Cd

Cd

Cd

Cd

Cd

Cd

Cv

-Cd

Ax

ial

lay

er n

Ax

ial

lay

er1

Rd

-Cv

Rd-C

vR

d-C

v

Cv

Cv

Co

ola

nt

Cv

Co

ola

nt

Co

ola

nt

Cv

Cv

Es(

F)

Cv

Cv

Cv

Es(

R)

Act

ive

par

t of

mac

hin

e

Cv

Cv

Es(

F)

Ro

tor

Ro

tor

Cd

Cd

Figure 2.7: Complete LP thermal model of the PMaSRM. The abbreviations usedare reported in Table 2.2. Loss sources are indicated as brown dots. Bold wordsrepresent temperature sensor locations in the experimental setup.


MACHINES

2.2.4 Experimental Evaluation

In order to verify the developed thermal model, the prototype PMaSRM describedin Table 2.1 and Figure 2.1 has been equipped with a number of temperaturesensors. All temperatures are measured using a temperature logging system. Thetemperature sensors mounted in the rotor are connected to the measurement setupusing a slip-ring unit mounted on the rotor shaft.

A comparison between the experimental and LP model results can be foundin Figure 2.8. In this test, the speed is kept at 150 rpm (5 Hz) and the machineproduces 90 Nm of torque. The ambient temperature is 25C and the flow rateand temperature of the inlet water are 25.0 cm3/s and 25.5C, respectively. Themachine is also tested experimentally at the same load (90 Nm) and 1500 rpm. Theambient temperature is 25.9C and the flow rate and temperature of the inlet waterare 27.2 cm3/s and 26.2C, respectively. The results are reported in Figure 2.9.

At 150 rpm and 90 Nm load, due to the low electrical frequency (5 Hz), theiron losses are neglected and the copper losses in the winding represent the singlesource of losses. Since the resistivity of copper is temperature dependent, the cal-culated copper losses are corrected according to the resulting winding temperaturedistribution using iteration.

At 1500 rpm and 90 Nm, the iron losses are estimated by measuring the outputand input power and then subtracting the calculated copper losses where proximitylosses have been computed using the analytical method outlined in [64] and 2DElectromagnetic FEA. The eddy current losses in the permanent magnets are com-puted using 3D electromagnetic FEA (see Chapter 4). Additionally, the frictionlosses are negligible at this speed [65, 66].

Remark: A detailed description of the experimental setup implemented and alsoa comparison between the experimental and the thermal model results at steadystate can be found in Section V in [Paper I]. Additionally, the PM eddy currentloss computation methods are discussed in Chapter 4.


0 20 40 60 80 1000

20

40

60

80

100

120

140

Experiment

Thermal model

End winding

Max. Magnet

Bearing

Time (min)

Tem

per

atu

re(

C)

Figure 2.8: Comparison between experimental and LP modeling results: 90 Nm,150 rpm.


MACHINES

0 20 40 60 80 1000

20

40

60

80

100

120

140

Experiment

Thermal model

End winding

Max. Magnet

Bearing

Time (min)

Tem

per

atu

re(

C)

Figure 2.9: Comparison between experimental and LP modeling results: 90 Nm,1500 rpm.

2.3. THERMAL MODELING OF DIRECTLY COOLED ELECTRIC

MACHINES 29

Oil outletOil outlet

Housing oilchannel

Oil inlet

a)b)

c)

Stator oil channels

123

4

5

6

78

910

11 12 13 1415

1617

18

19

2021

222324

12 mm

Housing oil channel

Stator oil channels

End Winding

Stator back

d)

Figure 2.10: a) Machine outside view; b) Inside view; c) Stator lamination design;d) Structure of the cooling system (the arrows indicate fluid flow).

2.3 Thermal Modeling of Directly Cooled Electric

Machines

The cooling system studied provides direct cooling of the stator back and on theouter surface of the end winding body. The cooling system comprises of a housing oilchannel in the circumferential direction and 24 stator oil channels in the stator backoriented axially. The inlet nipple is mounted close to the top part of the machine.The housing oil channel has a height and width of 2.5 and 25 mm, respectively. Theheight and width of each stator oil channel are 2 mm and 12 mm. After passing thestator channels, the oil drops off on the outer surface of the end winding body andis then collected at the bottom of the housing interior. The collected oil is emptiedusing two outlets mounted in both end shields.

The cooling system described above is presented in more detail in Figure 2.10.


MACHINES

2.3.1 Thermal Modeling Approach

In this thesis, focus is put on thermal modeling of the critical parts of a directlycooled machine including the stator slots, the end winding body and the coolingsystem. Since the coolant flow is not uniformly distributed in the cooling channels,the heat transfer to the coolant varies from channel to channel in the circumferentialdirection. The same statement can be made for the heat transfer to the coolantflowing on the end winding surfaces. Also, the gathered oil at the bottom of themachine at higher inlet coolant flow rates enhances the heat transfer from the endwinding to the coolant at the lower parts of the machine. The above mentionedstates of fluid lead to a varying temperature distribution in the slots and also in thedifferent outer surfaces of the end winding. Therefore, a multi-segment structureis proposed for the thermal model that divides the stator, slots, end windings andcooling system into a number of angular segments, as shown in Figure 2.15 at theend of this chapter.

In addition, to accurately model thermal effects of the cooling system designed,a detailed information of the oil distribution in different parts of the machine isneeded. In this work, CFD simulations are used to compute the coolant flow dis-tribution.

Thermal Model of the Stator Winding

In order to model the thermal effects in the stator winding, a realistic picture of theheat transfer directions in the active and end winding parts of the stator winding isneeded. A cross-sectional view of the end winding arrangement extending outsidea stator slot is illustrated in Figure 2.11. As seen, the end winding arrangementincludes the end winding ring, where the conductors are oriented mainly in thecircumferential direction, and the conductors that connect the end winding ring tothe active part of the winding inside the stator slots.

Based on the coolant flow distribution in the cooling channels and the endwinding surface, as discussed above, the heat transfer inside the end winding ar-rangement and in the stator slots follows the following paths:

1. From the stator slots, through the steel laminations and to the oil flowing inthe cooling channels.

2. From the end winding arrangement to the oil dropped from the end of thecooling channels.

3. From the end winding arrangement into the stator slots.

4. From the end winding arrangement to the oil gathered at the bottom of themachine.

To model the heat transfer in the active part of the stator winding correspondingto the first item, the thermal modeling approach presented in Section 2.2.1 is used.


MACHINES 31

Act

ive

par

t of

mac

hin

e

End w

indin

g r

ing

Node

LE

W

LE

W,b

LE

W,a

Figure 2.11: Cross-sectional view of the end winding arrangement.

The axial temperature variation is neglected here and the below expression is usedto model the conductive heat transfer in the impregnation layer between copperlayer i and copper layer i + 1.

RW,i =1

2πLAλImp

ln[

4 (ri + tImp) + 2∆r

+ 4

√

(ri + tImp)2 + ∆r (ri + tImp) + ∆r2/2

]

− ln[

4ri + 2∆r + 4√

ri2 + ∆rri + ∆r2/2

]

(2.13)

In this work, the heat transfer from the end winding nodes to the stator slots(through the distance LEW,b in Figure 2.11) is represented by REw,Dr,1 and REw,NDr,1

corresponding to the drive and non-drive ends of the machines. In this part of themachine, the produced heat passes first through the impregnation material in theend winding ring and then through the conductors that connect the end windingring to each stator slot. The thermal resistance that represents the heat transferin the conductors between the end winding ring and each stator slot (through thedistance LEW,a in Figure 2.11) on the drive side of the machine can be expressedas

REW,Dr,1a =LEW,a

λCuAStrNStr

. (2.14)

Here, LEW,a is the length of the conductors that connect the stator slot to theend winding ring (see Figure 2.11), λCu is the thermal conductivity of copper, AStr


MACHINES

is the conductor cross-sectional area, and NStr is the number of conductors in eachslot.

The thermal resistance representing the heat transfer from the end winding nodeto the conductors that connect the end winding ring to each stator slot (throughthe distance LEW,b − LEW,a in Figure 2.11) can be expressed as

REW,Dr,1b =LEW,b − LEW,a

λEWAEW

. (2.15)

Here, λEW is the equivalent thermal conductivity of the end winding ring; con-sidered as a homogenous body and approximated as [42]:

λEW = λImp

(1 + FEW) λCu + (1 − FEW) λImp

(1 − FEW) λCu + (1 + FEW) λImp

(2.16)

where FEW is the average end winding fill factor. Also, AEW is the end windingcross-sectional area divided by the number of slots NSlot. Hence, AEW can beapproximated as

AEW =π

(

D2EW,Out − D2

EW,In

)

4NSlot

(2.17)

where DEW,In and DEW,Out are the inner and outer diameters of the end windingring. Finally, REW,Dr,1 can now be approximated as

REW,Dr,1 = REW,Dr,1a + REW,Dr,1b (2.18)

The axially directed heat transfer from the end winding node to the axial endof the end winding arrangement (through the distance LEW−LEW,b in Figure 2.11)is represented by REw,Dr,2 and REw,NDr,2, and is obtained as

REW,Dr,2 =LEW − LEW,b

λEWAEW

(2.19)

where LEW is the axial length of the end winding. Analogous expressions for thenon-drive side of the electric machines can be easily derived.

The heat transfer between the end winding nodes in each angular segment isrepresented by REW,Bt which can be expressed as

REW,Bt =π (DEW,Out + DEW,In)

NSlotλCuFEW (DEW,Out − DEW,In) (LEW − LEW,a). (2.20)

Remark: The derivation of (2.20) is detailed in Appendix B of [Paper 3].The proposed thermal model of angular segment n of the winding is illus-

trated in Figure 2.12. As seen, the above mentioned thermal resistances model-ing the conductive heat transfer inside the end winding are represented as darkgrey blocks. Also, the C concentric elliptical impregnation layers in the stator slot,


MACHINES 33

Figure 2.12: Thermal model of angular segment n of the winding. The arrowsindicate the nodes in where loss sources are introduced.

RW,1, RW,2, . . . RW,C , are represented as light grey blocks. Finally, the convectiveheat transfer from the end winding surface to the oil ROil,EW,n is represented inlight orange. Since the housing oil channel is centered axially (see Figure 2.10), itis assumed that the oil distribution is identical on the drive and non-drive sides ofthe machines.

Heat Transfer to Coolant: CFD Simulations

To model the heat transfer to the coolant, an accurate estimation of the fluid flowinside the cooling channels and on the end winding surfaces is needed. In this work,in order to determine the flow rate in the oil channels, as well as the fluid velocityon the outer surface of the end windings, two sets of CFD simulations have beencarried out using the software Fluent2. These CFD simulations are based on theassumption that the level of the gathered oil at the bottom of the machines can beneglected. Hence, they are valid only up to a certain inlet flow rate.

First, the cooling channels including the housing and stator oil channels areimplemented in the CFD software. The single inlet of the housing oil channel andthe 48 outlets of the stator oil channels are modeled using appropriate boundaryconditions. Figure 2.13 shows the predicted flow rate distribution in the oil channelsfor two different inlet flow rates. On the one hand, due to the large cross-sectionalarea and the low height/width ratio of the channels, the coolant is very unevenly

2Fluent is a registered trademark by Ansys, Inc. in the United States.


MACHINES

5.3e-6

5.2e-1

1.3e-1

2.6e-1

3.9e-1

Figure 2.13: Predicted fluid velocity in the housing and stator oil channels (m/s).The inlet oil flow rate is 3.5 lit/min.

distributed between the different oil channels, so that heat transfer to the coolantoccurs mainly in the channels located close to the top part of the machine. On theother hand, the gathered oil at the bottom of the machine cools the lower machineparts. To compute the fluid velocity on the end winding surfaces, the obtained fluidflow rates in the stator oil channels are used as inlet boundary conditions in theCFD simulation model of the end winding. Figure 2.14 shows the fluid velocity closeto the outer surface of the end winding body at an inlet flow rate of 3.5 lit/min.These results show that for this and lower flow rates, the coolant flow on the endwinding body is essentially circumferential.

The computed fluid velocity is used to model the heat transfer from the statorback and end winding to the coolant flowing in the channels and on the end windingsurface.

The below formulation from [67] is employed to obtain the Nusselt number Nu

in a rectangular channel for a laminar flow:

Nu = 7.49 − 17.02 (HCh/WCh) + 22.43(HCh/WCh)2

− 9.94 (HCh/WCh)3

+(0.065RePrDH/LCh)

1 +[

0.04(RePrDH/LCh)2/3] . (2.21)

Here, Re and Pr are the dimensionless Reynolds and Prandtl numbers, respectively,HCh/WCh is the height/width ratio of the channel, DH is the hydraulic diameterof the channel, and LCh is the length of each cooling channel.


MACHINES 35

3.0e-5

1.4e-1

2.9e-2

6.5e-2

1.1e-1

Figure 2.14: Predicted fluid velocity on the outer surface of the end winding body(m/s).

To model the heat transfer from the end winding surface to the oil, the followingequation from [45] is used to obtain the Nusselt number:

Nu = 0.664Re1/2Pr

1/3 (2.22)

Eq. (2.22) is developed for a laminar flow on a flat surface. Since in the multi-layer model, the end winding cylindrical surface is divided into a number of smallersurfaces, each surface corresponding to a specific layer is approximated as a flatsurface.

If the temperature difference on the surface of the end winding is negligible,e.g. for a very low inlet flow rate, that the level of the collected oil at the buttonof the machine is very low, it can be assumed that all the oil comes out from theaxial channels at the top part of the machine. As a result, the following equationfrom [45] can be used to find the Nusselt number for an external flow to a cylinder(end winding), and the CFD simulation for the end winding part is not needed.

Nu = 0.3 +0.62Re

1/2Pr1/3

[

1 + (0.4/Pr)2/3

]1/4

[

1 + (Re

282000)5/8

]4/5

. (2.23)

Using the Nusselt numbers, the convective heat transfer coefficients, and as aresult, the corresponding thermal resistances can be easily obtained. It should bepointed out that the Reynolds number parameter (Re) used in (2.21)-(2.23) is afunction of the fluid velocity.


MACHINES

Remark: The formulas used to model the convective heat transfer to the coolantare illustrated in Section III-C in [Paper 3].

For the part of the machine that is submerged in the oil at the bottom, the fluidvelocity can be simply found using the inlet coolant flow rate which is equal to theoutlet coolant flow rate at the steady state flow, and the geometry of the end space.The level of the oil is found using the Bernoulli equation.

2.3.2 Complete LP Thermal Model

As discussed at the beginning of this chapter, the developed thermal model fordirectly cooled machines has a multi-segment structure. The complete thermalmodel is presented in Figure 2.15 where the stator laminations, stator winding,housing and cooling system are divided into n angular segments. The thermalresistances of the stator (RStat and RStat,Bt) are visualized as light green blocks. Thelight grey (RW) and the dark grey blocks (REW,Dr, REW,NDr, and REW,Bt) representthe stator slots and the end winding, respectively. Note that (RW) includes (RW,1),(RW,2), . . . (RW,C) corresponding to C layers of the active parts of winding (seeFigure 2.12). Also, (REW,Dr) and (REW,NDr) are divided into (REW,Dr,1, REW,Dr,2)and (REW,NDr,1, REW,NDr,2) as described in the Section 2.3.1. The dark greenblocks (RHous) represent the thermal model of the housing and the blue blocks(RAir,Hs,1, RAir,Hs,2,..., RAir,Hs,n) represent the convective heat transfer from thehousing fins to the ambient. The light orange blocks (ROil,EW,1 and ROil,EW,2)represent the convective heat transfer from the end winding to the coolant (oil) andthe dark orange blocks (ROil,St,1, ROil,St,2,..., ROil,St,n) represent the heat transferfrom the stator back to the oil flowing in the channels.

LP thermal models of the remaining parts of the machine, e.g. rotor, air-gap,frame, end shields, and the convective heat transfer from the inner solid parts ofthe machine to the air can be found in [28, 30, 43, 55].

2.3.3 Experimental Evaluation

In order to evaluate the developed thermal model, the proposed approach is testedon two identical machines impregnated using Epoxylite and varnish which are themost commonly used materials filling the stator slots and end winding bodies.Additionally, a cooling loop including the container, pump, flow meter, radiatorand fan is provided to introduce the coolant (oil) with a controllable flow rate tothe machine’s inlet nipple.

A comparison between the measured and predicted end winding temperaturedistribution in the machines impregnated using varnish and Epoxylite is presentedin Figure 2.16 and Figure 2.17, respectively. In Figure 2.16 and Figure 2.17, Sen-sor 1 is located in the middle of the end winding part corresponding to slot 22 (seeFigure 2.10), and the sensors 2,3,..., 6 are distributed evenly every 60 degrees clock-wise. In these tests, the inlet flow rate is regulated using the pump at 3.0 lit/minand 3.5 lit/min. Also the applied current to the winding varies from 3.7 A to 4.1 A.


MACHINES 37

Figure 2.15: The developed LP thermal model of the directly-oil-cooled electricmachines.

Remark: A detailed description of the experimental setup implemented andalso comparison between the experimental and the thermal model results at a widerange of cooling conditions and current amplitudes can be found in Section IV in[Paper III].

The experimental results demonstrate that the thermal model is able to estimatehot-spot temperature in the end winding and also the temperature distribution indifferent winding parts with a good accuracy (the maximum observed relative erroris 6%).


MACHINES

1 2 3 4 5 6100

110

120

130

140

150

T

emp

eratu

re(C

)

Sensor number (-)

4.1 A, 3.5 lit/min3.7 A, 3.0 lit/min

Figure 2.16: A comparison between measured and estimated temperature distribu-tion in the end winding body for the machine manufactured using varnish. Thesolid and the dashed lines represent the experimental and simulation results, re-spectively.

2.4 Summary of Chapter

In this chapter, thermal modeling methods for liquid-cooled electric machines werediscussed. In the first part of this chapter, thermal effects of a water-cooledPMaSRM were studied and a thermal model was developed that is able to esti-mate the temperature distribution with a good accuracy in both the rotor and thewinding. The approach proposed for thermal modeling of the winding is based onpartial FEA of a single slot. The suggested FE model of the winding is relativelysimple and can be implemented and run in short time.

In the second part of this chapter, a thermal modeling approach for directlycooled electric machines was proposed. The required CFD simulations in this ap-proach are simple and only basic skills in CFD modeling and heat transfer areneeded.

In the proposed solutions in this chapter, major advantages of LP modeling andnumerical methods are exploited by avoiding time-consuming FE or CFD analysisof a complete machine. The main message in this chapter to the reader is thatmodeling the heat transfer even in the machines with complex structures (geometryand/or cooling system) is possible provided that the thermal modeling method is

2.4. SUMMARY OF CHAPTER 39

1 2 3 4 5 685

90

95

100

105

110

115

T

emp

eratu

re(C

)

Sensor number (-)

4.1 A, 3.5 lit/min3.7 A, 3.0 lit/min

Figure 2.17: A comparison between measured and estimated temperature distri-bution in the end winding body for the machine manufactured using Epoxylite.The solid and the dashed lines represent the experimental and simulation results,respectively.

chosen with care.

Chapter 3

Thermal Effects – Material

Selection

In this chapter, thermal effects of different materials used in electric machines areinvestigated. The materials considered are the impregnation material, filling thestator slots and the end winding bodies, and the steel laminations used to build thestator and rotor parts of the machine.

3.1 Introduction

In addition to electromagnetic and cooling system design, another parameter thatsignificantly affects the thermal behavior of the electric machinery is the thermalproperties of the materials used.

In the first part of this chapter, thermal effects of two available and one alterna-tive impregnation materials used to fill the stator slots and the end winding bodiesare compared. The available impregnation materials are epoxy and varnish. Var-nish has been used in electric machinery applications as insulation material for morethan 70 years [68]. Since varnish is cheap and can be easily applied, the majority ofstandard electric machinery are still being impregnated using this material. Duringthe last twenty years, epoxy has been introduced to electric machines, especiallyin high-performance applications. This material has a higher thermal conductivityin comparison to varnish which enhances the heat transfer in the active part ofthe winding and the end winding bodies. Therefore, a reduction in the hot-spottemperature of the winding can be realized. However, there is still a need forimpregnation materials with higher thermal conductivities. In this regard, a sili-cone based thermally conductive material (SbTCM1) is investigated as a windingimpregnation material in this thesis.

1Thermoset SC-320 Thermally Conductive Silicone Encapsulant, LORD Corporation, NC27511-7923 USA.

41

42 CHAPTER 3. THERMAL EFFECTS – MATERIAL SELECTION

In order to investigate the thermal effects of the aforementioned materials, ap-propriate cooling should be provided for both the active part of the stator windingand the end winding where these materials are applied. Therefore, the direct cool-ing method presented in Chapter 2 is chosen. Three identical induction machinesusing varnish, Epoxylite and SbTCM as winding impregnation materials have beenmanufactured and tested. The thermal modeling method for directly cooled ma-chines, presented in Chapter 2 is adopted to carry out a simulation study on thestudied materials.

In the second part of this chapter, the thermal effects of using different steellaminations in the stator and rotor of electric machines are evaluated. In this regard,two key parameters characterizing electromagnetic, loss and thermal properties ofthe steel lamination materials, namely the lamination thickness and the amount ofalloy contents, are investigated in the form of a simulation study.

3.2 Thermal Effects of Using Different Impregnation

Materials

In addition to thermal conductivity, the thermal impact of different impregnationmaterials is also dependent on the quality of the impregnation process, the coolingconditions and the loss levels. A winding impregnation process with high qualityresults in less air pockets in the impregnation body and a higher rate of heat trans-fer. Moreover, thermal effects of the impregnation material in machines operatingunder different cooling conditions and levels of loss production, are not identical.In this section, first a comparison between the thermal, electrical and mechanicalproperties of the studied materials is presented and then thermal effects, consideringthe above mentioned parameters, are studied.

3.2.1 Impregnation Materials Studied

The thermal, electrical and mechanical properties of the impregnation materialsconsidered are compared in Table 3.1. The thermal conductivity has influence onthe hot-spot temperature of the winding. The dielectric strength and volume resis-tivity are important since the impregnation materials should also act as an electricinsulator materials. It should be noted, however, that a high voltage differencebetween the conductors in one slot is not expected and paper insulation (liners)are normally used to separate the stator laminations from the conductors in theslot. In the case that two coils are put in one slot, the paper dividers are used toseparate the conductors of the coils. For materials applied using vacuum bars, alower viscosity leads to a convenient vacuum impregnation. The values reported inTable 3.1 for the viscosities of Epoxylite and SbTCM, are at 50C and 25C, re-spectively. These temperatures are reached during the impregnation process. Theimpregnation process using SbTCM is carried out at room temperature. However,

3.2. THERMAL EFFECTS OF USING DIFFERENT IMPREGNATION

MATERIALS 43

preheating Epoxylite to 50C results in a significantly lower viscosity that simplifiesthe impregnation process.

The working temperatures of Epoxylite and varnish as standard impregnationmaterials can be from 110C to 180C according to their classes [57]. Long-termtests (> 1000 hr) on SbTCM at 150C is carried out and no change in the electrical,mechanical and thermal properties of this material is observed.

Table 3.1: Comparison between the studied impregnation materials.

Materials Varnish Epoxylite SbTCMThermal conductivity (W/mk) 0.18 – 0.25 0.80 – 1.10 3.20Dielectric strength (kV/mm) ≈80 ≈20 ≈10Volume resistivity at 25C (Ω·cm) > 1015 > 1014 > 1014

Viscosity (Pa·s) - 3.5 (at 50C) 25 (at 25C)

3.2.2 Impregnation Process Using SbTCM

The machines manufactured using varnish and Epoxylite follow the standard pro-cedures of impregnation. The impregnation using SbTCM is similar to Epoxyliteand the material is applied to the stator slots and the end winding using a vacuumbar. The manufacturing process is shown in Figure 3.1.

First the material and the hardener are mixed at a 1:1 ratio by weight andvolume. A vacuum impregnation was then carried out using a vacuum bar at anaverage pressure of 80 kPa. Following the vacuum impregnation, the stationerypart of the machine was put in an oven for 90 minutes where temperature wasadjusted to 125C. Finally, the mould was removed from the end winding body.In Figure 3.1h), the qualities of impregnation on the first and second tries arecompared. At the first attempt, the impregnation process followed the vacuumimpregnation procedure using Epoxylite. However, at the second attempt, theimpregnation process was modified, e.g. the inner surface of the mould was notcovered by the grease.

3.2.3 Results

To investigate the influence of the impregnation process quality on the thermal im-pact of the impregnation material, a simulation study is carried out. The simulationstudy is based on the developed LP thermal model for directly cooled machines pre-sented in Chapter 2. The influence of the other parameters, e.g. cooling conditionsand loss levels, are studied through simulation results as well as correspondingexperimental tests.


a)

c)

e)

g)

b)

d)

f)

h)

Figure 3.1: Manufacturing process: a) Mixing hardener with resin; b-d) Impregna-tion process in vacuum bar; e) Curing in an industrial oven; f-g) Removing mouldfrom the end winding; h) The final products.

3.2. THERMAL EFFECTS OF USING DIFFERENT IMPREGNATION

MATERIALS 45

0.4 0.5 0.6 0.7 0.8 0.980

90

100

110

120

130

140

150

160

H

ot

spot

tem

per

atu

re(C

)

Impregnation goodness (-)

VarnishEpoxySbTCM

Figure 3.2: A comparison between the hot-spot temperatures of the machines manu-factured using different impregnation materials for different impregnation goodnessvalues.

Impregnation Goodness

Impregnation goodness2 is selected in order to determine the influence of the im-pregnation process quality on the thermal performance of different impregnationmaterials. A comparison between the hot-spot temperatures of the machines manu-factured using varnish, Epoxylite and SbTCM at a reasonable range of impregnationgoodness is presented in Figure 3.2. The current magnitude and inlet flow rate areassumed to be 4.1 A and 3.5 lit/min, respectively.

As seen in Figure 3.2, using impregnation materials with higher thermal con-ductivities leads to appreciable reductions in the hot-spot temperatures of electricmachines. Additionally, the electric machine impregnated using the SbTCM isless influenced by the quality of impregnation process compared to varnish andEpoxylite which is an advantage.

Remark: The complete simulation study is elaborated in Section III in [Pa-per IV].

2The impregnation goodness represents the volume of impregnation material over the totalvolume of the impregnation body including the air pockets and impregnation material.


Cooling Conditions and Loss Levels

Three identical machines using the studied impregnation materials have been man-ufactured and equipped with a number of PT100 temperature sensors in the activepart of the stator winding and the end winding bodies. In order to make sure thatthe sensors are located at the same positions in the three machines and temperaturedifferences are not due to the sensors displacements, the PT100 sensor locations arefirst investigated before applying the impregnation materials to the winding.

Remark: The full description of the temperature sensors investigation processcan be found in Section IV-B in [Paper IV].

The cooling loop designed for the directly-oil-cooled machines includes a pump,a fan, a radiator, a flow meter and an oil container. The experimental setup isshown in Figure 3.3.

Figure 3.3: Experimental setup. The metal end cap is replaced by an end cap madeof transparent material to enable visual observation of the level of oil gathered atthe bottom of the machine and also the oil distribution on the end winding surfaces.

The experimental tests are done at different current magnitudes and inlet flowrates. The minimum and maximum flow rates are set based on the level of thegathered oil at the bottom of the machine. At an inlet flow rate of 3.5 lit/min, thegathered oil is close to reach the lowermost part of the rotor and below 2.5 lit/min,there is no oil at the bottom of the machine.

The simulation and corresponding experimental results are presented for inletflow rates of 2.5 lit/min, 3.0 lit/min and 3.5 lit/min in Figure 3.4 – Figure 3.6,respectively. As seen, at an inlet flow rate of 2.5 lit/min, the hot-spot temperaturedifference between the machines impregnated using Epoxylite and SbTCM variesfrom 10C to 24C. Also, a temperature difference of 20C to 28C can be observed

3.3. THERMAL EFFECTS OF USING DIFFERENT STEEL LAMINATION

MATERIALS 47

between the machines manufactured using varnish and Epoxylite. Since the maxi-mum working temperature of the varnish used is 150C, it was not possible to testthe machine manufactured using varnish at current magnitudes higher than 3.7 A,at the flow rates of 2.5 lit/min and 3.0 lit/min. At higher flow rates, e.g. 3.0 lit/minand 3.5 lit/min, the thermal impact of using impregnation materials with higherthermal conductivities is diminished. However, the temperature differences are stillsubstantial.

3.2 3.4 3.6 3.8 4 4.2 4.4

60

70

80

90

100

110

120

130

140

Hot

spot

tem

per

atu

re(C

)

Current (A)

VarnishEpoxySbTCM

Figure 3.4: A comparison between the hot-spot temperatures of the machines man-ufactured using different impregnation materials, at an inlet coolant flow rate of2.5 lit/min. The solid lines and the dashed lines represent the experimental andsimulation results, respectively.

Remark: The full description of experimental setup, sensor locations and testconditions can be found in Section IV of [Paper IV].

3.3 Thermal Effects of Using Different Steel Lamination

Materials

The magnetic and thermal properties of electrical steel laminations are largely afunction of the amount of alloy contents [Paper V]. Materials with a high amountof alloy contents produce less iron losses in a specific frequency interval, and havea lower thermal conductivity. An increase in the amount of alloy contents inhibits


3.2 3.4 3.6 3.8 4 4.2 4.4

60

70

80

90

100

110

120

130

H

ot

spot

tem

per

atu

re(C

)

Current (A)

VarnishEpoxySbTCM


eddy currents and narrows the hysteresis loop of the material [Paper V]. Therefore,a reduction in the produced iron losses can be observed. However, laminations withhigher alloy contents have lower thermal conductivities that can result in difficultieswith heat dissipation from the stator and rotor laminations. Consequently, thehot-spot temperature in the critical parts of the machine increases. Furthermore,increasing the Si-content decreases the saturation point of the BH curve [69, 70].

Another important parameter of non-oriented steel laminations is the laminationthickness. Laminations with higher thicknesses produce more eddy current losses[70]-[Paper V].

As discussed, with changing the amount of alloy contents of the steel lami-nations, a trade-off between the produced iron losses and the thermal conductiv-ity is obvious. In this section, the thermal impact of laminations with differentthicknesses and amounts of alloy contents used in a PMaSRM designed for HEVapplications is investigated. The PMaSRM is detailed in Chapter 2.


MATERIALS 49

3.2 3.4 3.6 3.8 4 4.2 4.4

60

70

80

90

100

110

120

130

140

H

ot

spot

tem

per

atu

re(C

)

Current (A)

VarnishEpoxySbTCM


3.3.1 Comparison Between Different Steel Laminations

A comparison between laminations with different amounts of alloy contents and thesame thickness is presented in Table 3.2. The thickness of the selected materials is0.35 mm. As seen, with a 40% decrease in the amount of alloy contents, the samepercentage of increase in the produced iron losses can be observed. The change inthermal conductivity is a function of the temperature and at higher temperatures,the rate of change is reduced. Table 3.2 also shows that there is a strong connectionbetween the price and the amount of alloy contents.

Laminations with different thicknesses and identical amount of alloy contents(by % weight) are compared in Table 3.3. Since the amount of alloy contents doesnot change, the thermal conductivity of the materials is kept identical (32.0 W/mkat 100C). However, thicker steel laminations produce more eddy current losses.Also, the increases in lamination thickness reduce the manufacturing cost by up to50%.


Table 3.2: Comparison between laminations with different amounts of alloy contentsfor the same thickness (0.35 mm). The presented losses are measured at 50 Hz and1.5 T.

Material grade M235-35A M250-35A M300-35A M330-35A% Si+Al (pu) 1.00 0.91 0.74 0.58λ at 22C (W/mk) 19.9 20.6 22.8 26.2λ at 100C (W/mk) 29.1 31.0 32.0 34.4λ at 200C (W/mk) 32.4 34.4 35.4 36.8Loss (W/kg) 2.35 2.50 3.00 3.30Price (pu) 1.00 0.89 0.77 0.62

Table 3.3: Comparison between laminations with different thicknesses and the sameamount of alloy contents. Since the produced iron losses in NO18 at 50 Hz are small,the presented iron loss values are measured at 400 Hz and 1 T.

Material grade NO18 M300-35A M310-50A M350-65AThickness (mm) 0.18 0.35 0.50 0.65Loss (W/kg) 12.3 18.8 27.4 39.6Price (pu) 1.00 0.69 0.52 0.50

3.3.2 Iron Loss Estimation

The iron losses are estimated using the frequency-dependent iron loss characteristicsof the steel laminations, and the instantaneous magnetic flux density distributionobtained from 2D electromagnetic FEA of the studied PMaSRM.

Remark: The full description of the iron loss calculation method and the resultscan be found in Section II in [Paper V].

3.3.3 Thermal Impact of Using Laminations with Different

Qualities

Estimation of the temperature distribution in different parts of the PMaSRM con-sidered is enabled using the LP thermal model presented in Chapter 2 along withthe computed iron losses in the steel laminations and copper losses in the windingas inputs. Additionally, eddy current losses in the permanent magnet segments ofthe rotor are estimated using a 3D FEA of the electric machine (see Chapter 4 formore details).

A comparison between the resulting hot-spot temperatures of the winding andalso the temperature of the outermost magnet in the rotor can be found in Fig-ure 3.7 and Figure 3.8. The simulations are carried out for two operating points: 1)1500 rpm and 108 Nm which represents an operating condition close to the nominal


MATERIALS 51

point of operation, and 2) 3000 rpm and 72 Nm which represents operating in thefield weakening region.

235−35 250−35 300−35 330−35170

175

180

185

190

235−35 250−35 300−35 330−35

80

100

120

140

a)

Max.

win

din

gte

mp.

b)

Steel lamination

Steel lamination

Max.

roto

rte

mp.

1500 rpm

1500 rpm

3000 rpm

3000 rpm

Figure 3.7: Predicted temperatures (C) in the PMaSRM: a) Winding; b) Rotor.

As can be seen in Figure 3.7 and Figure 3.8, variations in the alloy contentsfor materials with the same thickness does not have a considerable impact on theresulting temperatures of the PMaSRM in consideration at the nominal operatingpoint (an increase of up to 2% in the temperature of the winding and rotor is found).Also, the increase in the temperature of the winding is minor in the field weakeningrange (3000 rpm and 72 Nm) and a 10% increase in the predicted temperatureof rotor is found. It should be noted that using materials with lower amounts ofalloy contents can result in a 40% reduction in the lamination cost. The change inthickness of the laminations has a significant influence on the thermal behavior ofthe studied PMaSRM and up to 75% increase in the temperature of the rotor isobtained in the field weakening region.


NO18 300−35 310−50 350−65170

180

190

200

NO18 300−35 310−50 350−65

80

100

120

140

160

180

a)

Max.

win

din

gte

mp.

b)

Steel lamination

Steel lamination

Max.

roto

rte

mp.

1500 rpm

1500 rpm

3000 rpm

3000 rpm

Figure 3.8: Resulting temperatures (C) in the PMaSRM: a) Winding; b) Rotor.


In this chapter, thermal effects of using different winding impregnation materialsand steel lamination qualities were evaluated. Two commonly used impregnationmaterials were first compared with a silicone based thermally conductive impregna-tion material. Significant reductions in the hot-spot temperature of winding wereachieved which is promising for thermal management of electric machines and alsothe resulting efficiency particularly in high-performance applications. However,more investigations especially on the mechanical properties of the studied SbTCMare needed to introduce this material as a standard impregnation material in electricmachinery.

In the second part of this chapter, a simulation study on the thermal effectsof using different steel laminations in the PMaSRM considered in previous chap-ters was carried out. Two parameters characterizing the thermal, mechanical andelectromagnetic properties of the steel laminations, the amount of alloy contentsand the lamination thickness, were studied. It was shown that using expensivematerials with higher amounts of alloy contents results in a minor reduction in the


temperature of machine critical parts in comparison to the materials with smallerthickness. To verify the simulations, corresponding experimental tests should bedone.

Since the thermal behavior of electric machines directly affects the electromag-netic performance and the efficiency, in the electric machine design process, theelectrical and thermal properties of the selected materials should be taken intoaccount. There is room for major improvements in the thermal performance ofelectric machinery by selecting the right available materials and also searching forinnovative materials with improved thermal properties that can be exploited incritical parts of the machine.

Chapter 4

Magnet Eddy Current Loss

Estimation

This chapter presents the FE-based approaches used to model eddy current lossesin the permanent magnet segments of the PM machines. In this regard, 2D elec-tromagnetic time-domain FEA is combined with analytical models and partial 3Dfrequency-domain FE models to reduce the required computation time for eddy cur-rent loss computation.

4.1 Introduction

PM machines containing rare-earth permanent magnets enable high power densitiesand a high efficiency. However, during recent years the price of rare-earth basedPM materials has increased substantially which limits the practical use of thesemachines to high-performance applications that can deal with the high materialcost. PMaSRMs designed using a lower amount of permanent magnets represent acost effective option for several high-performance applications including automotivetraction [71, 72].

As is well known, the field harmonics induce eddy current losses in the perma-nent magnet segments of the rotor. Since the rotor is a passive part from a thermalpoint of view, to prevent from thermal demagnetization of the magnets an accu-rate estimation of the rotor losses is necessary. In this chapter, different FE-basedapproaches used to estimate the eddy current losses in the permanent magnets arereviewed and discussed.

First, the eddy current losses are computed using 3D electromagnetic FEA. ThePMaSRM presented in Chapter 2 is considered as a case study. This method is thencompared with a combination of a 2D time-domain FEA of the complete machineand a 3D frequency-domain FEA of only the rotor. Finally, a combination of 2Dtime-domain FEA of the PMaSRM and analytical modeling of eddy current losses

55

56 CHAPTER 4. MAGNET EDDY CURRENT LOSS ESTIMATION

is used to estimate the produced eddy current losses in the permanent magnets.The last two approaches aim to reduce the required eddy current computation time.

4.2 Complete 3D Electromagnetic FEM

Induced eddy currents in the permanent magnets of the rotor represent a 3D phe-nomenon that, in general, cannot be modeled using pure 2D calculations. However,running a 3D FEA of the complete machine is, in general, a very time-consumingprocess, especially when stator phase current harmonics are considered.

To reduce the induced eddy current losses, the rotor magnets are commonlyaxially divided into a number of segments. Provided that axial end effects canbe neglected and also under the assumption that the magnet segments are fullyisolated electrically from each other, the 3D FE model of the complete machinecan be reduced to a 3D model comprising of only the axial half of a single rotormagnet. The developed 3D FE model of the PMaSRM presented in Chapter 2is shown in Figure 4.1. As seen, the front side of the model is covered by an airlayer to model the insulation between the magnet segments. A symmetry boundarycondition is applied in the axial ends of the model. The FE software used in thisstudy is JMAG.

The FEA steps followed can be summarized as:

1. A quarter 3D model of the machine considered is created.

2. The computational mesh is generated. The mesh element size is determinedaccording to the skin depth of the induced eddy currents. The skin depth δcan be approximated as [57]

δ =1√

πfµσ(4.1)

where µ and σ are the magnetic permeability and electric conductivity of thepermanent magnet and f is the frequency of the applied magnetic field. Themesh element size is then chosen considerably smaller than δ.

3. The proper material properties and winding configuration are applied.

4. Since half of a magnet segment in the axial direction is modeled, the propersymmetry boundary condition is implemented.

5. Periodic boundary condition is implemented since a quarter of the machineis modeled.

6. The FEA is executed.

4.3. 2D AND PARTIAL 3D ELECTROMAGNETIC FEA 57

a)

b)

Figure 4.1: The developed 3D FE model used for eddy current losses computation:a) Front view; b) Behind view.

4.3 2D and Partial 3D Electromagnetic FEA

The 3D FE analysis described above is very time-consuming, especially if currentharmonics should be taken into account. In order to reduce the computation time,a 2D FEA combined with partial 3D frequency-domain FE modeling is consideredas well. Using this approach, the air-gap flux density is computed using a 2Dtransient FEA of the complete machine. Then, the 3D FE model of the rotoris implemented and the computed air-gap flux density from the 2D FE model isapplied to the rotor surface as a boundary condition. The 3D frequency-domainFEA (of the rotor only) is executed and the eddy current losses in the permanentmagnet segments are computed.

The 2D and partial 3D FE models of the studied PMaSRM are presented inFigure 4.2 and Figure 4.3, respectively.


Magnet 1

Magnet 2

Magnet 3

Magnet 4

Figure 4.2: The developed 2DFEM model used to compute the flux density in theair-gap.

Figure 4.3: The developed partial 3DFEM model used to compute the inducededdy current losses in the magnets.

4.4. 2D ELECTROMAGNETIC FEA AND ANALYTICAL MODELING 59

4.4 2D Electromagnetic FEA and Analytical Modeling

In this section, the partial 3D FEM presented in Section 4.3 is compared with ananalytical approach. Here, the computed rotor flux density from the 2D transientFEA is exported to Matlab and the eddy current losses are then estimated fordifferent harmonic components of the flux density waveform. The eddy currentlosses PEddy can then be expressed as [73, 74]

PEddy =bt

2δσ|H |2

sinha

δ− sin

a

δ

cosha

δ− cos

a

δ

− 4a4t

π5

|γ2H |2σ

×∞

∑

n=0

(

ξn2 − 2βni

2)

βnrξn3 sinh βnrb

(2n + 1)5|βn|6 (cosh βnrb + cos βnib)

− 4a4t

π5

|γ2H |2σ

×∞

∑

n=0

(

ξn2 − 2βnr

2)

βniξn3 sin βnrb

(2n + 1)5|βn|6 (cosh βnrb + cos βnib)(4.2)

where a, b and t are the magnet width, length and thickness, respectively as illus-trated in Figure 4.4. The skin depth δ is obtained using (4.1) and the additionalparameters ξn, γ and βn are found as

ξn = (2n + 1)π

a(4.3)

γ =1 + j

δ(4.4)

βn =

√

λn2 + γ2 = βnr + jβni. (4.5)

Eq. 4.2 is derived by solving Maxwell’s equations for time harmonic fields wherea magnetic field (H) is applied to a thin conductor (t >> a and b).

H

t

a

b

J

Figure 4.4: Assumed model geometry valid for the analytical expression of inducededdy currents presented in (4.2).


4.5 Evaluation

A comparison between the computed eddy current losses in the permanent magnetsusing the above described FE-based approaches is presented here for two differentoperating points. The results presented in Figure 4.5 is corresponding to a rotorspeed of 3000 rpm when the stator winding is supplied by a sinusoidal current withan amplitude of 11 A corresponding to id =−10.83 A and iq =1.91 A. The resultingtorque at this operating point is 72 Nm which is approximately 65% of the ratedtorque.

In Figure 4.6, the results are compared at 6000 rpm (200 Hz) and the currentmagnitude is kept at 11 A1 (id = −10.96 A and iq = 0.96 A). The resulting torqueis 39.5 Nm.

As seen in Figure 4.5 and Figure 4.6, the eddy current loss estimation approachbased on a combination of 2D FEM and 3D partial FEM seems to enable eddycurrent loss estimation with a good accuracy in the magnets located closer to theshaft (Magnet 1, 2 and 3). However, the induced eddy current losses in Magnet4, located close to the outer rotor surface, is underestimated which is likely dueto saturation in the rotor surface close to Magnet 4. Since the 3D partial FEAis carried out in a frequency domain, saturation in the steel laminations is notbasically taken into account that results in underestimating the applied magneticfield to Magnet 4 and, in turn, underestimating the produced eddy current losses.

The results obtained using a combination of 2D FEM and analytical modelingdeviate somewhat from the fully FE-based approaches considered. This can likelybe attributed to the fact that in this method, the flux variation in the permanentmagnet segments is not taken into account. It should be noted, however, thatthis method represents a rapid approach to estimate the eddy current losses inmachines where the flux density distribution is close to uniform in the permanentmagnet segments.

In order to compare the required time for eddy current loss computation usingdifferent approaches, the FE simulations are performed using a personal computerequipped with a Core 2 Duo (TM) 3.16 GHz CPU and a 4.0 Gb memory. Theruntime for the full 3D FE simulation is ≈20 hours. However, the required timefor the 3D frequency-domain FEA of the rotor and the 2D FEA of the completemachine are ≈21 and 12 minutes, respectively.


In this chapter, three FE-based approaches used to model induced eddy currentlosses in magnet segments of PM machines have been reviewed and compared inthe form of a small case study.

1The operating points presented in Figure 4.5 and Figure 4.6 are selected to compare differenteddy current loss estimation methods and thermal, mechanical and voltage constrains of thestudied PMaSRM are not taken into consideration


0

5

10

Magnet 1 Magnet 2 Magnet 3 Magnet 4Permanent Magnet

Eddy

curr

ent

loss

es(W

)

Full 3D2D+Partial 3D2D+Anly.

Figure 4.5: A comparison between the computed eddy current losses in the perma-nent magnet segments using different FE-based approaches at 3000 rpm.

0

20

40

60

Magnet 1 Magnet 2 Magnet 3 Magnet 4Permanent Magnet

Eddy

curr

ent

loss

es(W

)

Full 3D2D+Partial 3D2D+Anly.

Figure 4.6: A comparison between the computed eddy current losses in the perma-nent magnet segments using different FE-based approaches at 6000 rpm.

It was found that the reduction in computational time resulted in a lower ac-curacy. The approach based on FEA of the complete machine can be used inapplications where a high accuracy is needed, for example in a final design verifi-cation stage. However, in design procedures where a high number of iterations isrequired and the aim is mainly to compare different designs, the methods basedon the combinations of 2D FEA, partial 3D FEA and analytical models can besuggested.

This work can be continued with experimental verification of the computed eddycurrent losses in the permanent magnet segments.

Chapter 5

Concluding Remarks

The research presented in this thesis is summarized and conclusions are drawnin this chapter. Also, some suggestions and ideas for future work related to thedifferent topics included in this thesis are provided.

5.1 Summary

In this thesis, thermal effects in the electric machines have been studied. Thefocus has been put on machines designed for automotive traction. However, theresults can also be applied to machines designed for the other high-performanceapplications, e.g. aerospace or railway traction.

The thermal modeling approaches presented in this work show how an appropri-ate combination of analytical and numerical methods enables accurate modeling ofheat transfer in electric machinery. The developed thermal models are fairly simplesince the approaches are restricted to a limited need for numerical solutions, e.g.FE methods and CFD simulations.

In the PMaSRM considered the coolant is trapped in the housing jacket and thecoolant flow and heat transfer to the coolant can be modeled using the available an-alytical expressions. However, the hot-spot temperature estimation in the windingand the temperature distribution in the rotor structure cannot be addressed withequally simple means. Since modeling the heat transfer in the winding includingthe stator slots and the end winding bodies using available analytical models wasnot possible, a partial FE model containing the coils of a single slot of the windingwas developed. The FE modeling results were then used to determine certain keyparameters of the analytical models presented in this thesis. For the rotor, a simpleradial model is proposed and evaluated. The complete thermal LP model of thestudied PMaSRM was developed using the proposed thermal models of the windingand rotor. Experimental results verified that the developed thermal model enablesthe estimation of the temperature distribution in the machine with a high accuracy.It should be noted that the proposed algorithm to model the heat transfer in the

63

64 CHAPTER 5. CONCLUDING REMARKS

winding is a general method that can be applied to the winding of the other kindsof electric machines.

To model the thermal effects of the directly cooled machines considered, an ac-curate estimation of the coolant flow distribution is needed. Therefore, a relativelysimple set of CFD simulations are carried out to estimate the fluid distribution inthe channels and the end winding body surfaces. The results obtained from theCFD simulations were then used to model the heat transfer to the coolant andto develop a LP thermal model of the complete machine. The structure of thepresented LP thermal model is modified so that the heat transfer in the circumfer-ential direction is taken into consideration. The heat transfer in the circumferentialdirection is mainly due to the nonuniform distribution of the coolant on the outersurfaces of the end winding bodies and in the cooling channels, and also because ofthe gathered oil at the bottom of the machine. Corresponding experimental testsdemonstrated that the developed thermal model is able to predict the temperaturedistribution in the machine with a sufficient accuracy.

The above considered examples convey the important message that suitable andpurposeful combinations of numerical and analytical techniques enable the devel-opment of accurate thermal models for electric machines with complex structures.

In this thesis, the influence of using different impregnation materials filling thestator slots and the end winding bodies were studied. Significant reductions in thehot-spot temperatures of the windings impregnated using materials with higherthermal conductivities at different operating and cooling conditions were observed.

In addition, the thermal effects of using different steel laminations were inves-tigated. The influences of the lamination thickness and amount of alloy contentson the thermal behavior of the machine were studies separately. The simulationresults indicated that using lamination materials with different amounts of alloycontents leads to a 2%-10% variation in the hot-spot temperatures of rotor andwinding for the PMaSRM in consideration. There is a clear trade-off in thermalconductivity and loss characteristics of the steel lamination materials considered.Therefore, choosing a more expensive lamination does not necessarily result in asignificant improvement in the thermal behavior of the machine.

Finally, a review of available FE based approaches used to compute the eddycurrent losses in the permanent magnet segments of the rotor of a PMaSRM wascarried out and advantages and disadvantages of each approach were discussed.

5.2 Proposal for Future Work

Some ideas for future research in this field are outlined as follows.

In Chapter 3, the thermal effects of a silicone based thermally conductive ma-terial were studied. The results obtained are promising. However, more investi-gations, especially on the mechanical properties of the material, are needed. Ad-ditionally, the search for alternative impregnation materials with a high thermalconductivity should be continued.

5.2. PROPOSAL FOR FUTURE WORK 65

A simulation study on the thermal impact of using different steel laminationmaterials on a PMaSRM was carried out. Unfortunately, experimental evaluationof the obtained simulation results is not included in this thesis. However, suchan experiment-based study is recommended. Furthermore, work on the thermalimpact of new materials with different thermal and electromagnetic characteristicsthat have shown the possibility of being used in the stator and rotor bodies of theelectric machines is suggested.

Available commercial and educational tools for modeling heat transfer in electricmachines are mainly based on pure LP modeling or numerical methods [23]. Asshown in this thesis, in order to estimate the temperature distribution in complexmachine structures accurately, advantages of both analytical and numerical meth-ods should be exploited. Some efforts are being made to provide tools that enablethis coupling between numerical and analytical approaches1. However, there is alack of a tool that provides the user with access to the LP network, the FE andCFD models. Consequently, in the opinion of the author, the development of a toolwith the above mentioned specifications would be a major step forward in thermalmodeling of electric machines.

In Chapter 4, different FE-based approaches available for computing eddy cur-rent losses in permanent magnets were reviewed and compared. Since the amountof eddy current losses in the permanent magnets used in the rotor is low in com-parison to the other sources of losses in electric machines, measuring the accurateamount of the produced eddy current losses is difficult. Hence, the development ofa practical method for eddy current loss measurement in the rotor of a PM machinewould be beneficial for electric machine community.

1In the new version of the software Motor-CAD that was released in April, 2013, the user hasaccess to both the LP and FE models.

List of Figures

1.1 LP thermal model of a sample electric machine. The abbreviations usedare reported in Table 1.1. . . . . . . . . . . . . . . . . . . . . . . . . . . 3

1.2 Thermal FEA of the stationary part of a sample electric machine: a)FE model; b) FEA results. . . . . . . . . . . . . . . . . . . . . . . . . . 5

2.1 Topology of the PMaSRM in consideration. . . . . . . . . . . . . . . . . 14

2.2 3D thermal FE winding modeling steps. Light green: end winding ring,dark green: slot wedge, yellow: conductors, dark blue: slot liner, lightblue: slot impregnation. . . . . . . . . . . . . . . . . . . . . . . . . . . . 17

2.3 Proposed elliptical model of the active winding. The colors are in agree-ment with the colors used in Figure 2.2. . . . . . . . . . . . . . . . . . . 18

2.4 Proposed LP thermal model of the winding with n axial layers and melliptical copper layers in the slot. The selected colors are in agreementwith the colors used in Figure 2.2 and Figure 2.3. . . . . . . . . . . . . . 19

2.5 Iterative approach for determining winding model parameters. . . . . . 21

2.6 Radial LP thermal rotor model. Loss sources are indicated as brown dots. 23

2.7 Complete LP thermal model of the PMaSRM. The abbreviations usedare reported in Table 2.2. Loss sources are indicated as brown dots.Bold words represent temperature sensor locations in the experimentalsetup. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25

2.8 Comparison between experimental and LP modeling results: 90 Nm,150 rpm. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27

2.9 Comparison between experimental and LP modeling results: 90 Nm,1500 rpm. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28

2.10 a) Machine outside view; b) Inside view; c) Stator lamination design; d)Structure of the cooling system (the arrows indicate fluid flow). . . . . . 29

2.11 Cross-sectional view of the end winding arrangement. . . . . . . . . . . 31

2.12 Thermal model of angular segment n of the winding. The arrows indi-cate the nodes in where loss sources are introduced. . . . . . . . . . . . 33

2.13 Predicted fluid velocity in the housing and stator oil channels (m/s).The inlet oil flow rate is 3.5 lit/min. . . . . . . . . . . . . . . . . . . . . 34

67

68 List of Figures

2.14 Predicted fluid velocity on the outer surface of the end winding body(m/s). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35

2.15 The developed LP thermal model of the directly-oil-cooled electric ma-chines. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37

2.16 A comparison between measured and estimated temperature distribu-tion in the end winding body for the machine manufactured using var-nish. The solid and the dashed lines represent the experimental andsimulation results, respectively. . . . . . . . . . . . . . . . . . . . . . . . 38

2.17 A comparison between measured and estimated temperature distribu-tion in the end winding body for the machine manufactured using Epoxylite.The solid and the dashed lines represent the experimental and simula-tion results, respectively. . . . . . . . . . . . . . . . . . . . . . . . . . . . 39

3.1 Manufacturing process: a) Mixing hardener with resin; b-d) Impreg-nation process in vacuum bar; e) Curing in an industrial oven; f-g)Removing mould from the end winding; h) The final products. . . . . . 44

3.2 A comparison between the hot-spot temperatures of the machines man-ufactured using different impregnation materials for different impregna-tion goodness values. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45

3.3 Experimental setup. The metal end cap is replaced by an end cap madeof transparent material to enable visual observation of the level of oilgathered at the bottom of the machine and also the oil distribution onthe end winding surfaces. . . . . . . . . . . . . . . . . . . . . . . . . . . 46

3.4 A comparison between the hot-spot temperatures of the machines man-ufactured using different impregnation materials, at an inlet coolant flowrate of 2.5 lit/min. The solid lines and the dashed lines represent theexperimental and simulation results, respectively. . . . . . . . . . . . . . 47



3.7 Predicted temperatures (C) in the PMaSRM: a) Winding; b) Rotor. . 513.8 Resulting temperatures (C) in the PMaSRM: a) Winding; b) Rotor. . 52

4.1 The developed 3D FE model used for eddy current losses computation:a) Front view; b) Behind view. . . . . . . . . . . . . . . . . . . . . . . . 57

4.2 The developed 2DFEM model used to compute the flux density in theair-gap. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58

4.3 The developed partial 3DFEM model used to compute the induced eddycurrent losses in the magnets. . . . . . . . . . . . . . . . . . . . . . . . . 58

List of Figures 69

4.4 Assumed model geometry valid for the analytical expression of inducededdy currents presented in (4.2). . . . . . . . . . . . . . . . . . . . . . . 59

4.5 A comparison between the computed eddy current losses in the perma-nent magnet segments using different FE-based approaches at 3000 rpm. 61

4.6 A comparison between the computed eddy current losses in the perma-nent magnet segments using different FE-based approaches at 6000 rpm. 61

List of Tables

1.1 Abbreviations used in Figure 1.1. . . . . . . . . . . . . . . . . . . . . . . 3

2.1 PMaSRM data. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 152.2 Abbreviations used in Figure 2.7. . . . . . . . . . . . . . . . . . . . . . . 24

3.1 Comparison between the studied impregnation materials. . . . . . . . . 433.2 Comparison between laminations with different amounts of alloy con-

tents for the same thickness (0.35 mm). The presented losses are mea-sured at 50 Hz and 1.5 T. . . . . . . . . . . . . . . . . . . . . . . . . . . 50

3.3 Comparison between laminations with different thicknesses and the sameamount of alloy contents. Since the produced iron losses in NO18 at50 Hz are small, the presented iron loss values are measured at 400 Hzand 1 T. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50

71

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Appendix A

Glossary of Symbols and

Abbreviations

SymbolsAEW End winding ring cross sectional area of one slot in the axial directionAImp Impregnation area in a slotASlot Slot areaAStr Conductor cross sectional areaa Magnet widthb Magnet lengthDEW,In Inner diameter of the end winding ringDEW,Out Outer diameter of the end winding ringDH Hydraulic diameterFEW Average end winding fill factorf FrequencyH Magnetic fieldHCh Stator channel heighthSlot Slot heightLA Active length of machineLCh Stator channel lengthLCyl Axial length of a cylindrical rotor sectionLEW Axial length of the end winding coilsLEW,a Length of the conductors that connect the active part of winding to

the end winding ringLEW,b Distance between the end winding node and the active part of

machineLEW,F Length of forward end winding coilsLEW,R Length of rear end winding coilsm Number of copper layers in a slotNi Number of winding strands in winding layer i

81

82 Appendix A. Glossary of Symbols and Abbreviations

NSlot Total number of slotsNStr Total number of winding strands in a slotNu Nusselt numbern Number of axial layersPEddy Eddy current lossesPr Prandtl numberRAir,Hs,n Thermal resistance between the housing and the ambient nodesRAW,i Thermal resistance in the axial direction of the ith copper layerRCyl Thermal resistance of a cylindrical rotor sectionREW,Bt Thermal resistance that represents heat transfer inside the end

winding ringREW,Dr Thermal resistance that represents heat transfer from the active part

of winding to the end winding (drive side)REW,F Thermal resistance between the forward end winding node and the

active windingREW,NDr Thermal resistance that represents heat transfer from the active part

of winding to the end winding (non-drive side)REW,R Thermal resistance between the rear end winding node and the active

windingRe Reynolds numberROil,EW,n Thermal resistance that models heat transfer from the end winding

to the coolant in the angular segment n of the thermal modelROil,St,n Thermal resistance that models heat transfer from the stator back to

the coolant in the angular segment n of the thermal modelRStat Thermal resistance that models heat transfer in the stator laminationsRStat,Bt Thermal resistance that models heat transfer in the stator laminations

between the thermal model angular segmentsRW,i Thermal resistance of the impregnation between the copper layer i

and i + 1RW,j,i Thermal resistance in the jth axial layer of the impregnation between

copper layer i and copper layer i + 1rEx Outer radius of a cylindrical rotor sectionrIn Inner radius of a cylindrical rotor sectionri Minor radius of elliptical copper layer ir′

i Major radius of elliptical copper layer it Magnet ticknesstAG Equivalent air-gap length between the slot liner and slot walltImp Impregnation layer thicknesstLnr Thickness of liner and slot insulationWCh Stator channel widthµ Magnetic permeability∆r Length difference between the major and minor radii of the elliptical

copper layer iδSpan Angular span of a cylindrical rotor section

Appendix A. Glossary of Symbols and Abbreviations 83

λ Thermal conductivityλCu Thermal conductivity of copperλEW Equivalent thermal conductivity of the end windingλImp Thermal conductivity of the winding impregnationσ Electric conductivityσStr Winding strand density in a slotδ Skin depth

AbbreviationsBrn BearingCd ConductionCFD Computational fluid dynamicsCv ConvectionEc End capEs End spaceE-Wnd End windingF Forward (drive) side of electric machinesFE Finite elementFEA Finite element analysisFEM Finite element methodHs HousingIPM Interior permanent magnetLP Lumped parameterPlt PlatePM Permanent megnetPMaSRM Permanent magnet assisted synchronous reluctance machineR Rear (non-drive) side of electric machinesRd RadiationSbTCM Silicone based thermally conductive materialSht ShaftSl-Wl Slot wallSt-Br Stator boreTth Stator teethYk Stator yokeWnd Winding

Appendix B

Selected Publications

Paper I

S. Nategh, O. Wallmark, M. Leksell, and S. Zhao, “Thermal Analysis ofa PMaSRM Using Partial FEA and Lumped Parameter Modeling,” IEEETransactions on Energy Conversion, vol. 27, no. 2, pp. 477-488, June2012.

© 2012 IEEE.

85

Paper II

S. Nategh, O. Wallmark, and M. Leksell, “Thermal analysis of permanent-magnet synchronous reluctance machines,” in Proc. 14th European Con-ference on Power Electronics and Applications (EPE), Aug. 30-Sept. 1,2011.

© 2011 IEEE.

99

Paper III

S. Nategh, Z. Huang, O. Wallmark, M. Leksell, and A. Krings, “ThermalModeling of Directly Cooled Electric Machines Using Lumped Parameterand Limited CFD Analysis,” submitted to the IEEE Transactions on En-ergy Conversion.

111

Paper IV

S. Nategh, A. Krings, O. Wallmark, and M. Leksell, “Evaluation of Im-pregnation Materials for Thermal Management of Liquid-Cooled ElectricMachines,” submitted to the IEEE Transactions on Industrial Electronics.

According to the TIE regulations, the maximum number of pages forthe submitted manuscript is limited to eight. However, the final versioncan be extended to a higher number of pages. Therefore, the shortenedversion of this paper is submitted to be reviewed and the full version isavailable here.

123

Paper V

S. Nategh, A. Krings, Z. Huang, O. Wallmark, M. Leksell, and M. Lin-denmo, “Evaluation of stator and rotor lamination materials for thermalmanagement of a PMaSRM,” in Proc. XXth International Conference onElectrical Machines (ICEM), 2-5 September, 2012.

© 2012 IEEE.

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