manufacturing effects on iron losses in electrical machines847378/fulltext01.pdf · manufacturing e...

DEGREE PROJECT, IN , SECOND LEVELELECTRIC POWER ENGINEERING

STOCKHOLM, SWEDEN 2015

Manufacturing Effects on Iron Lossesin Electrical Machines

KONSTANTINOS BOURCHAS

KTH ROYAL INSTITUTE OF TECHNOLOGY

ELECTRICAL ENGINEERING


Konstantinos Bourchas

Master of Science Thesis in Electrical Machines and Drivesat the School of Electrical Engineering

Royal Institute of TechnologyStockholm, Sweden, June 2015

Supervisors: Dr. Alexander Stening(ABB LV Motors)Dr. Freddy Gyllensten (ABB LV Motors)

Examiner: Docent Juliette Soulard (KTH)

XR-EE-E2C 2015:006

ii

Manufacturing Effects on Iron Losses in Electrical Machines.KONSTANTINOS BOURCHAS

Copyright c©2015 by Konstantinos Bourchas.All rights reserved.

School of Electrical EngineeringDepartment of Energy ConversionRoyal Institute of TechnologySE-100 44 StockholmSweden

Contents

List of Symbols vii

List of Abbreviations ix

Abstract xi

Sammafattning xiii

Acknowledgments xvi

1 Introduction 11.1 Background . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11.2 Thesis Scope . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21.3 Outline . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2

2 Ferromagnetic Materials 32.1 Iron Losses . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3

2.1.1 Hysteresis Losses . . . . . . . . . . . . . . . . . . . . . . . . . . . . 32.1.2 Eddy Current Losses . . . . . . . . . . . . . . . . . . . . . . . . . . 4

2.2 Iron Loss Models . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52.2.1 Models based on the Steinmetz Equation . . . . . . . . . . . . . . . 52.2.2 Separation Models . . . . . . . . . . . . . . . . . . . . . . . . . . . 62.2.3 Hysteresis Models . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

2.3 Characterization of Magnetic Properties of Electrical Steels . . . . . . . . . 72.4 Magnetic Measurements by means of the Epstein Frame . . . . . . . . . . 82.5 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11

3 Manufacturing Effects on Iron Losses in Electrical Machines 133.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133.2 Forming the Core Laminations . . . . . . . . . . . . . . . . . . . . . . . . . 13

3.2.1 Mechanical Cutting . . . . . . . . . . . . . . . . . . . . . . . . . . . 143.2.1.1 Affected Area due to Mechanical Cutting . . . . . . . . . 143.2.1.2 Effect of Mechanical Cutting on Hysteresis and Eddy Cur-

rent Losses . . . . . . . . . . . . . . . . . . . . . . . . . . 143.2.1.3 Effect of Mechanical Cutting on the Magnetizing Current 153.2.1.4 Si-Content . . . . . . . . . . . . . . . . . . . . . . . . . . . 153.2.1.5 Cutting Perpendicular or Parallel to the Rolling Direction 15

3.2.2 Laser Cutting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16

iii

iv CONTENTS

3.2.2.1 CO2 and Fiber Laser . . . . . . . . . . . . . . . . . . . . . 163.2.2.2 Laser Settings . . . . . . . . . . . . . . . . . . . . . . . . . 163.2.2.3 Spatial Distribution of Degradation . . . . . . . . . . . . . 16

3.2.3 Comparison between Mechanical and Laser Cutting . . . . . . . . . 163.2.4 Abrasive Water Jet . . . . . . . . . . . . . . . . . . . . . . . . . . . 17

3.3 Core Assembly . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 173.3.1 Pressing during Stacking . . . . . . . . . . . . . . . . . . . . . . . . 173.3.2 Welding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 183.3.3 Cleating . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 183.3.4 Gluing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18

3.4 Motor Assembly . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 193.4.1 Shaft Insertion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 193.4.2 Pressing into Frame . . . . . . . . . . . . . . . . . . . . . . . . . . . 193.4.3 Rotor Machining . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19

3.5 Manufacturing Mitigations . . . . . . . . . . . . . . . . . . . . . . . . . . . 203.5.1 Annealing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 203.5.2 Tuning of Laser Settings . . . . . . . . . . . . . . . . . . . . . . . . 203.5.3 Maintenance of Punching Machine . . . . . . . . . . . . . . . . . . 21

3.6 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

4 Measurements 234.1 Introduction to the Experiments . . . . . . . . . . . . . . . . . . . . . . . . 23

4.1.1 Motivation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 234.1.2 Test Setup . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 244.1.3 Repeatability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 244.1.4 Tested Grades . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26

4.2 Mechanical Cutting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 274.2.1 M400-50A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28

4.2.1.1 Cutting Effect on Iron Losses and Permeability . . . . . . 284.2.1.2 Iron Loss Separation . . . . . . . . . . . . . . . . . . . . . 30

4.2.2 M270-50A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 324.2.2.1 Cutting Effect on Iron Losses and Permeability . . . . . . 324.2.2.2 Iron Loss Separation . . . . . . . . . . . . . . . . . . . . . 34

4.2.3 NO20 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 364.2.3.1 Cutting Effect on Iron Losses and Permeability . . . . . . 364.2.3.2 Iron Loss Separation . . . . . . . . . . . . . . . . . . . . . 37

4.2.4 Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 394.2.5 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 40

4.3 Laser Cutting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 414.3.1 Comparison among different Laser Settings . . . . . . . . . . . . . . 41

4.3.1.1 Selection of Laser Settings and Laser Machine . . . . . . . 414.3.1.2 Degradation of M400-50A due to Laser Cutting with Var-

ious Settings . . . . . . . . . . . . . . . . . . . . . . . . . 424.3.2 Cutting effect due to laser . . . . . . . . . . . . . . . . . . . . . . . 45

4.3.2.1 Cutting Effect due to Best Laser Setting (Set 8 ) . . . . . 464.3.2.2 Cutting Effect due to Worst Laser Setting (Set 2 ) . . . . . 49

4.3.3 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51

CONTENTS v

4.4 Comparison between Mechanical and Laser Cutting . . . . . . . . . . . . . 524.4.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 524.4.2 M400-50A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 524.4.3 M270-50A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 534.4.4 NO20 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 534.4.5 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54

4.5 Welding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 554.5.1 Measurement Results . . . . . . . . . . . . . . . . . . . . . . . . . . 554.5.2 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56

5 Simulations 575.1 Separation in Yoke and Teeth Regions . . . . . . . . . . . . . . . . . . . . 575.2 Model for Permeability at High Flux Densities . . . . . . . . . . . . . . . . 595.3 Simulations of an Induction Motor . . . . . . . . . . . . . . . . . . . . . . 605.4 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 64

6 Conclusions and Future Work 656.1 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 656.2 Future Work . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 66

Appendix A Guillotine Cutting 69A.1 M400-50A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 69A.2 M270-50A . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 72A.3 NO20 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75

Appendix B Laser Cutting 79

Bibliography 81

vi CONTENTS

List of Symbols

β exponential coefficient [-]µ0 vacuum permeability [(T· m)/A]µr relative permeability [-]A cross sectional area [m2]B flux density [T]Br remanent flux density [T]f frequency [Hz]H magnetizing field [A/m]Hc coercive field [A/m]I1 current in the primary winding [A]kec eddy current loss coefficient

[W

kg·(T·Hz)2]

kexc excess loss coefficient[

Wkg·(T ·Hz)1.5

]khyst hysteresis loss coefficient

[W

kg·T 2·Hz

]l length of single Epstein strip [m]lm effective path length of flux [m]m total mass of test specimen [kg]ma active mass of test specimen [kg]Ms saturation magnetization [A/m]N1 number of turns of primary winding of Epstein frame [-]N2 number of turns of secondary winding of Epstein frame [-]pec eddy current loss density [W/kg]pexc excess loss density [W/kg]pFe iron loss density [W/kg]physt hysteresis loss density [W/kg]Pc total losses of test sample [W]Pm measured power [W]Ps specific iron losses [W]Ri total resistance of the instruments that are connected to the secondary winding [Ohms]

vii

viii CONTENTS

List of Abbreviations

ELE Exponential Law Extrapolation

FEM Finite Element Method

HAZ Heat Affected Zone

LASER Light Amplification by Stimulated Emission of Radiation

MMF Magnetomotive Force

RD Parallel to the rolling direction

SST Single Sheet Test

TD Perpendicular to the rolling direction

ix

x CONTENTS

Abstract

In this master thesis, the magnetic properties of SiFe laminations after cutting and weld-ing are studied. The permeability and the iron loss density are investigated since they arecritical characteristics for the performance of electrical machines. The magnetic measure-ments are conducted on an Epstein frame for sinusoidal variations of the magnetic fluxdensity at frequencies of 50, 100 and 200 Hz, according to IEC 404-2. Mechanical cuttingwith guillotine and cutting by means of fiber and CO2 laser are performed. The influenceof the fiber laser settings is also investigated. Especially the assisting gas pressure andthe power, speed and frequency of the laser beam are considered.

In order to increase the cutting effect, the specimens include Epstein strips with 1,2 and 3 additional cutting edges along their length. It is found that mechanical cuttingdegrades the magnetic properties of the material less than laser cutting. For 1.8% Silaminations, mechanical cutting causes up to 35% higher iron loss density and 63% lowerpermeability, compared to standard Epstein strips (30 mm wide). The correspondingdegradation for laser cut laminations is 65% iron loss density increase and 65% per-meability drop. Material of lower thickness but with the same Si-content shows lowermagnetic deterioration. Additionally, laser cutting with high-power/high-speed charac-teristics leads to the best magnetic characteristics among 15 laser settings. High speedsettings have positive impact on productivity, since the cutting time decreases.

The influence of welding is investigated by means of Epstein measurements. The testspecimens include strips with 1, 3, 5 and 10 welding points. Experiments show an ironloss increase up to 50% with a corresponding 62% reduction in the permeability.

A model that incorporates the cutting effect is developed and implemented in a FEM-based motor design software. Simulations are made for a reference induction motor.The results indicate a 30% increase in the iron losses compared to a model that doesnot consider the cutting effect. In case of laser cut core laminations, this increase reaches50%. The degradation profile considers also the deteriorated magnetizing properties. Thisleads to increased nominal current up to 1.7% for mechanically cut laminations and 3.4%for laser cut laminations. This corresponds to a 1.4% and 2.6% reduced power factor,respectively.

Keywords – electrical machine, induction motor, iron losses, relative per-meability, guillotine, fiber laser, CO2 laser, laser settings, cutting effect,welding, electrical steel.

xi

xii CONTENTS

Sammanfattning

I detta examensarbete studeras hur de magnetiska egenskaperna hos SiFe-pl̊at p̊averkasav skärning och svetsning. Permeabilitet och järnförlustdensitet undersöks eftersom deär kritiska variabler för elektriska maskiners prestanda. De magnetiska mätningarnagenomfördes p̊a en Epstein ram med en flödesfrekvens p̊a 50, 100 och 200 Hz, enligtIEC 404-2. Effekterna av mekanisk skärning med giljotin samt skärning med fiber- ochCO2-laser studerades. Inverkan av olika fiberlaserinställningar undersöktes ocks̊a genomatt variera gastrycket, skärhastigheten samt frekvensen och effekten av laserstr̊alen.

För att öka skäreffekten inkluderades Epsteinremsor med ytterligare 1, 2 och 3 längsg̊aendeskärsnitt. Det visas att mekanisk skärning har en mindre p̊averkan p̊a de magnetiskaegenskaperna hos materialet än vad laserskärning har. Mätningar p̊a pl̊at med 1.8% Sivisar att d̊a prov med tre extra längsg̊aende giljotinklipp används kan permeabilitetenreduceras med upp till 63% och järnförlusterna kan öka med upp till 35%. Motsvaranderesultat för laserskurna pl̊atar visar en permeabilitetsreduktion p̊a upp till 65% ochen järnförlustökning p̊a upp till 65%. Ur studien av de tv̊a studerade skärprocessernaframkommer även att tunnare pl̊at p̊averkas mindre negativt än tjockare pl̊at. Ett antalolika inställningar har provats för att utreda hur olika parametrar p̊averkar effekternaav laserskärning. Studien indikerar att skärning med hög effekt och hög hastighet gerden minsta p̊averkan p̊a materialets magnetiska egenskaper. Vilket även har en positivinverkan p̊a produktiviteten vid laserskärning.

Epsteinprover har även utförts för att undersöka vilka effekter som introduceras d̊aSiFe-pl̊at svetsas. Provstyckena bestod av remsor med en, tre, fem och 10 svetspunkter.Experimenten visar en järnförlustökning med upp till 50% samt en permeabilitetsreduk-tion upp till 62% d̊a pl̊atarna svetsats samman tv̊a och tv̊a.

En modell för att studera effekterna av de förändrade materialegenskaperna vid skärningp̊a en induktionsmotor utvecklas och implementeras i en FEM-baserad mjukvara. Resul-taten tyder p̊a en järnförlustökning med 30% d̊a skäreffekten orsakad av giljotin beaktas.Vid simulering av laserskuren pl̊at kan denna ökning vara s̊a stor som 50%. Det framkom-mer även att laserskärningen kan reducera effektfaktorn s̊a mycket som 2.6%.

Nyckelord - elektrisk maskin, induktion motor, järnförluster, relativ per-meabilitet, giljotin, fiberlaser, CO2-laser, laser inställningar, skäreffekt, svet-sning, elektromagnetisk pl̊at

xiii

xiv CONTENTS

Acknowledgments

This thesis concludes my MSc in Electric Power Engineering at KTH, Royal Instituteof Technology in Stockholm. The thesis was conducted at ABB LV Motors at Väster̊as,Sweden.

First of all, I am grateful to my main supervisor at ABB, Dr. Alexander Stening, forhis support during my master thesis project. His regular feedback helped me to improvethe content and the text of the thesis. I truly appreciate the discussions we had thathelped me to develop myself as motor designer. His understanding and feedback madethis master thesis an experience of a lifetime.

Secondly, I would like to thank my manager at ABB, Dr. Freddy Gyllensten. It hasbeen my honor working with him. He gave me very valuable feedback and many ideaswhich improved the scientific significance of the current master thesis. He was alwayseager to share with me his experience on motor design, something that helped me a lotto deeply understand many aspects of this field.

Moreover, many thanks go to my examiner at KTH, Docent Juliette Soulard for herfeedback on the text of my report. Additionally, in her course, I gained the initial highlevel knowledge on motor design. She also helped me to finish the latter stages of themaster thesis (presentation at KTH) as soon as possible and I am really thankful to herfor that.

Furthermore, I would like to thank Dr. Arvid Broddefalk and Magnus Lindenmofrom Surahammars Bruks AB. Their help and support throughout the project was veryvaluable. Many thanks go also to Mats Dahlen from Gerdins AB for the productivecooperation we had during my project.

Next, I would like to thank my colleagues at the group of Technology Development ofABB LV Motors, Lic.Tech. Rathna Chitroju, Lic.Tech. Kashif Khan and Dr. Dan Fors.The working environment was great and they helped me a lot with the discussions thatwe had.

I would also like to thank Dr. Andreas Krings from ABB Corporate Research. Hissuggestions in the early stages of the project helped me a lot and gave me good scientificdirections.

I am also thankful to Lic. Tech. Mats Leksell for the cooperation that we had duringmy time in the Eco Marathon team of KTH and during my time as research assistantat E2C lab. He gave me the opportunity to develop my skills as engineer and to gainvaluable hands-on experience. I will always be grateful for that.

I am also very thankful to my close friend and KTH classmate Alexandra Kapidoufor her support and all the great moments that we spent at Väster̊as, during our masterthesis elaboration. I would also like to thank my dear friend and KTH classmate TinRabuzin. We spent a lot of quality time discussing about our future and our dreams. I

xv

xvi CONTENTS

am also thankful to my friend Nikolaos Apostolopoulos who has also supported me asbrother during the last two years and he gave me real help to take important decisions.

Additionally, I would like to thank all of my friends in Stockholm, who helped me tohave a wonderful time in Sweden during the last two years. I will not say all of the namesbecause I am afraid that I will forget someone. I hope that we will continue be in touchfor the rest of our lives.

I am also grateful to my family. First to my parents Georgios Bourchas and ElliGeralidou for their endless support throughout my life. Without them, I would not beable to fulfill any of my dreams. I will always be thankful to them because they made mewhat I am today. Secondly, my two sisters Lina and Kally have offered me endless supportthrough their love and I am grateful for that. I would also like to apologize to them forthe time that I have not spent with them because of my studies abroad. Finally, I wishto express my deepest gratitude to Konstantina Nikolaou for her love and understandingnot only during my thesis but also during those two years of my master studies.

Konstantinos BourchasStockholm, Sweden

June 2015

Chapter 1

Introduction

This chapter describes the background and the scope of this master thesis. Moreover, anoutline of the thesis is presented.

1.1 Background

The last 20 years, the climate change has raised concerns worldwide [1]. This is the reasonwhy many regulatory authorities have established legislations regarding the decrease ofthe energy consumption, aiming at the reduction of the CO2 emissions. The EuropeanUnion has established a policy to combat the environmental pollution and the climatechange. One of the goals of this policy is a 20% increase in the energy efficiency by 2020and 27% by 2030 [1].

Electric motors account approximately for 65% of the energy use in industry [2]. Thatmeans that any increase of the efficiency of these motors can potentially lead to majorenergy savings. IEC 60034-30-1:2014 is a standard that regulates the efficiency levels ofthe industrial induction motors around the world [3]. Figure 1.1 illustrates the range ofthe efficiency classes as defined by IEC 60034-30-1:2014.

Figure 1.1: Efficiency classes of low voltage, 4-pole induction motors according to IEC60034-30-1:2014 [3].

The design of high-efficient motors requires accurate motor models. The iron loss

1

2 CHAPTER 1. INTRODUCTION

models are often considered as the main source of error for the prediction of the motorefficiency. The estimation of the iron losses in the stator and the rotor of the motordepends on the analytical description of the physical phenomena that cause these losses[4]. Furthermore, the processes during the production of an electrical motor lead todeterioration of the magnetic properties of the core materials. Therefore, they should betaken into consideration [5]. According to [6], the major source of steel degradation iscutting. This master thesis focuses on the effect of different cutting techniques on thestator and rotor material’s magnetic properties. The effect of welding is also investigated.

1.2 Thesis Scope

The main objective of this thesis is to investigate the change of the magnetizing and ironloss characteristics of electrical steel due to cutting by means of guillotine and laser.

The project is divided in different stages as follows:

• Measurement and analysis of the effect of mechanical and laser cutting.

• Comparison between the two cutting techniques.

• Measurement and analysis of welding effect.

• Development of finite element method (FEM) model that incorporates the experi-mental results.

1.3 Outline

This master thesis consists of three parts. The first part is a literature review on theprocesses that affect the iron losses of the electrical steel. The second part includesexperimental results which concern the cutting and the welding effect. Finally, the thirdpart of the thesis presents a FEM model that incorporates the experimental results. Thethesis report is separated in six chapters with the following content.Chapter 1 presents the background and the scope of the project.Chapter 2 gives an overview of the physics, the models and the evaluation of the magneticproperties of the ferromagnetic materials.Chapter 3 discusses the main manufacturing steps that influence the properties of themagnetic materials.Chapter 4 presents the experiments which are conducted on the Epstein frame regardingthe magnetic properties of SiFe laminations after cutting and welding.Chapter 5 introduces a FEM model that incorporates the cutting effect.In Chapter 6, conclusions are drawn and suggestions for future investigations are made.

Chapter 2

Ferromagnetic Materials

Ferromagnetic materials consist of ferromagnetic domains which are small areas, wherethe magnetic dipoles are parallel to each other [7]. A basic characteristic of ferromagneticmaterials is the hysteresis. In this chapter, the loss mechanisms are presented. Addition-ally, the main iron loss models are given and the recommended methods for the evaluationof the magnetic properties of the ferromagnetic materials are discussed.

2.1 Iron Losses

The iron losses are also referred as core losses. They are created by the varying magneticfield in the iron parts of the machine. The two basic components of the iron losses arethe hysteresis and the eddy current losses. Both of these components result in the samephysical phenomenon which is Joule heating.

2.1.1 Hysteresis Losses

The hysteresis losses are mostly dependent on the microstructure of the magnetic material[8]. The electrical steel consists of uniformly magnetized regions, called domains. Whenno external field is applied, the statistical sum of the magnetization of all the domains iszero [9]. The neighboring domains are magnetized in an opposite direction and the borderthat separates two such domains is called domain wall. The domain wall is actually anenergy zone through which the magnetization gradually changes direction [4].

When an external field is applied, the domain wall moves in the direction of the field.As a result, the area of the domain whose magnetization is aligned to the field grows atthe expense of the area of the neighboring domain which has opposite magnetization [10].However, non-magnetic impurities (like carbon and sulfur) can be found in the electricalsteel. These impurities act as pinning sites and they hinder the domain wall motion[4, 9]. In this case, the domain wall overcomes the pinning sites by being subjected toincreased external field. After a certain field, the domain wall rapidly overcomes (jumps)the pinning site. This phenomenon is called Barkhausen jump and through this rapidmovement, eddy currents are induced. These eddy currents cause Joule losses which inthis case also called hysteresis losses [4]. The Barkhausen effect is one of the main reasonsof the hysteresis loop behavior of the magnetic material.

For low values of external field H, the domain walls do not overcome the pinningsites. Therefore there are no Barkhausen jumps. At this region, the magnetization is

3

4 CHAPTER 2. FERROMAGNETIC MATERIALS

reversible, which means that if the magnetizing field is removed, then the magnetizationof the material returns to zero. The slope of the BH curve in this region is expressedby the initial susceptibility [11]. For higher magnetizing fields, the magnetization of thematerial is no longer reversible [11]. If the flux density in the material reaches saturationand the external field is removed, the material sustains a remanent magnetization whichis expressed by Br. In order to demagnetize the material, an opposing coercive magneticfield Hc should be applied. This behavior of the ferromagnetic material is illustrated bythe hysteresis loop as depicted in Figure 2.1.

Figure 2.1: Initial BH curve and hysteresis loop of ferromagnetic materials.

2.1.2 Eddy Current Losses

The variation of the magnetic flux over time induces an electrical field in the magneticcore, which causes the flow of eddy currents. According to Lenz law, these currents tendto oppose the field that produced them. This current flow results in Joule losses, alsocalled eddy current losses [12, 13]. The most effective method to reduce the eddy currentlosses is to divide the core into thin sheet laminations as depicted in Figure 2.2.

2.2. IRON LOSS MODELS 5

Figure 2.2: Solid and laminated iron core to reduce the flow of the eddy currents (redlines). The vector of flux density B is perpendicular to the surface of the core.

2.2 Iron Loss Models

The estimation of the iron losses is one of the most challenging aspects in the design andanalysis of an electrical machine. There are many different analytical approaches whichestimate the iron losses for different induction levels and frequencies. In this section, themost common iron loss models are presented.

2.2.1 Models based on the Steinmetz Equation

Steinmetz was the first who developed an analytical approach to predict the iron lossesin 1892 [14, 15] . Equation 2.1 is called Steinmetz Equation and expresses the iron lossesof the material. This equation is valid only for sinusoidal flux density waveforms.

pFe = kSEfαB̂β (2.1)

where pFe are the specific iron losses (W/kg), f is the frequency and B̂ is the peak valueof the flux density. The coefficients kSE, α and β are obtained through fitting in theexperimental results.

Based on the Steinmetz’s initial empirical equation, several iron loss models havebeen developed. The Modified Steinmetz Equation [16] is such model and can be used forarbitrary flux density waveforms. The Modified Steinmetz Equation is given in formula2.2.

pFe = kSEfα−1eq B̂

βf (2.2)

where feq is an equivalent frequency which depends on the rate of change of the fluxdensity and is expressed as:

feq =2

∆B2π2

∫ T0

(dB

dt

)2dt (2.3)

Other approaches based on the Steinmetz Equation can be found in [17].


2.2.2 Separation Models

Another approach is splitting the iron losses in two or three terms. These terms correspondto the hysteresis, the eddy current and the excess losses. Table 2.1 summarizes the modelsthat are based on the separation approach.

Jordan[18]

pFe = physt + pec = khystfB̂2 + kecf

2B̂2 (2.4)

Pry and Bean[19]

pFe = physt + ηapec =

khystfB̂2 + ηexckecf

2B̂2(2.5)

Bertotti[4]

pFe = physt + pec + pexc

= khystfB̂2 + kecf

2B̂2 + kexcf1.5B̂1.5

(2.6)

Jacobs[20]

pFe = k′hystfB̂

2 + (kec + a1B̂a2)B̂2f 2 (2.7)

where k′hyst = khyst(1 +BminBmax

(r − 1))

Table 2.1: Separation models for iron loss estimation.

Jordan in [18] separates the iron losses in hysteresis and eddy current losses. Bothterms depend on the amplitude of the flux density. However, the hysteresis term dependson f (static losses), while the eddy current losses depend on f 2 (dynamic losses) [9].

Even though Equation 2.4 holds for NiFe laminations, it is not accurate for SiFe [9].This is the reason why, Pry and Bean in [19] introduced a correction factor ηexc to minimizethe discrepancy between the measured and the predicted eddy current losses.

Bertotti in [4] gave a physical meaning to this discrepancy by introducing a third term,which is the excess (or anomalous) losses. This term corresponds to the mesoscopic scalein the magnetization process and depends on the eddy currents due to the domain wallmotion, assuming that the hysteresis losses and the Barkhausen effect are disregarded [4].

Jacobs in [20], evolved Bertotti’s model in order to take into account the rotationallosses (through the constant r) and the high order losses (through the constant α2), whichis caused by the magnetic saturation. The separation models are valid for a frequencyrange, where the skin effect is negligible [8].

2.2.3 Hysteresis Models

Apart from the iron loss models that are based on the Steinmezt Equation and the sepa-ration concept, there are also models considering the hysteresis behavior of the magnetic

2.3. CHARACTERIZATIONOFMAGNETIC PROPERTIES OF ELECTRICAL STEELS7

material. Such models are more complicated, but they have higher accuracy. The Preisachmodel is one of these models [21]. More details about the various iron loss models can befound in [17].

2.3 Characterization of Magnetic Properties of Elec-

trical Steels

In this section, the basic methods of characterization of the magnetic properties of theelectrical steels are presented. Figure 2.3 gives an overview of these methods.

Figure 2.3: Overview of magnetic material characterization methods.

The characterization of magnetic materials is obtained through magnetic measure-ments. The magnetic measurements of electrical steels concern the determination of themagnetizing and iron loss characteristics of the material. The Epstein frame measure-ments and the Single Sheet Test (SST) are two methods for the characterization of stripshaped laminations. The geometry of the samples is simple and their dimensions aredetermined by standards. The catalogue data of the electrical steel manufacturers are ob-tained using Epstein frame measurements. The major drawback of these methods is thatthe geometry of the test specimens is not representative of the actual motor application.

Another method for the magnetic characterization is the measurements on a ring coretopology [22]. The main advantage of this method, compared to the Epstein and SST, isthat the geometry is representative of the stator yoke of electrical machines. Further tomention, this topology offers a closed flux path without any airgaps [9]. An alternativeapproach based on the ring core topology is conducting measurements on an actual statorcore. The principles of operation are the same as in the ring core topology. However, thestator teeth cause fringing effect, which means that the flux is not uniformly distributed.These effects can be corrected through models as presented in [23, 24]. More methodsand detailed description of magnetic measurements can be found in [25].


2.4 Magnetic Measurements by means of the Epstein

Frame

Measurements with the 25 cm Epstein frame is the standardized method to characterizethe magnetic properties of electrical steel and follows the IEC 404-2 standard [26]. Thesamples consist of rectangular strips. Therefore they are easy to cut and measure. Ad-ditionally, the effect of the cutting direction is canceled, since strips that are cut in therolling direction (RD) and strips are cut transversally to the rolling direction (TD) aremeasured simultaneously in the Epstein frame. The main characteristics of the Epsteinmeasurements, following the IEC 404-2 standards, are summarized below:

• The Epstein frame consists of 4 coils, as depicted in Figure 2.4. The strips undertest are inserted in these coils. Each of these coils includes a primary (excitation)and a secondary (measurement) winding. The primary windings are connected inseries, as illustrated in Figure 2.5. The same applies for the secondary windings.

• The Epstein strips form a square which has double-lapped joints. This way, each ofthe four branches has the same length and cross sectional area.

• The strip width shall be 30 mm while the strip length shall be in the range of 280mm-320 mm.

• The number of strips shall be a multiple of 4.• lm is the effective magnetic path of the Epstein frame and equal to 0.94 m (Figure

2.4).

Figure 2.4: The 25 cm Epstein frame [26].

2.4. MAGNETIC MEASUREMENTS BY MEANS OF THE EPSTEIN FRAME 9

Figure 2.5: Connection of windings in the Epstein frame.

The waveforms of the magnetizing current and the output voltage are illustrated inFigures 2.6-2.7. The magnetizing current is regulated so that the voltage in the secondaryis sinusoidal. This way, the condition of sinusoidal flux density is satisfied, as IEC 404-2defines.

Figure 2.6: Example of the current in the pri-mary winding of the Epstein frame.

Figure 2.7: Voltage in the secondary windingof the Epstein frame.

The total losses of the test specimen (i.e all the strips) are determined by Equation2.8.

Pc =N1N2

Pm −(1.111|Ū2|)2

Ri(2.8)


Where,

• Pc are the total losses of the test sample.

• N1 is the total number of turns of the primary winding.

• N2 is the total number of turns of the secondary winding.

• Pm is the measured power.

• |Ū2| is the average value of the rectified voltage that is induced in the secondarywinding.

• Ri is the total resistance of the instruments that are connected to the secondarywinding.

The specific iron losses are then determined by Equation 2.9.

Ps =Pcma

(2.9)

Where ma is the active mass of the test specimen and is defined as shown in Equation2.10.

ma =m · lm

4l(2.10)

Where, l is the length of one Epstein strip and m is the total mass of the test specimen(includes all strips of the Epstein frame).

The calculation of the magnetizing characteristics of the tested material lies on thedetermination of the values of the magnetizing field H and the corresponding inducedflux density B. The magnetizing field H is obtained through the current in the primarywinding I1 and is given by Equation 2.11.

H(t) =N1lmI1(t) (2.11)

The induced flux can be obtained directly by means of a fluxmeter or by digitallyintegrating the voltage of the secondary winding, as given in Equation 2.12 [9].

B(t) = − 1N2A

∫u2(t)dt (2.12)

Where A is the cross sectional area of the test specimen and is given by Equation 2.13.

A =m

4 · l · ρm(2.13)

with ρm being the conventional density as determined by IEC 404-13.

2.5. SUMMARY 11

2.5 Summary

In this chapter the physical phenomena behind the losses of the ferromagnetic materialswere shortly presented. Furthermore, analytical models to estimate the iron losses werediscussed. Historically, Steinmetz was the first to develop a mathematical model thatdescribes the iron losses as a function of the induction and the frequency. Nowadays,there are many approaches towards the iron loss estimation. The selection of an iron lossmodel is a trade-off between complexity of implementation and accuracy of estimation.

Additionally, the major methods of magnetic measurements were presented. SST andEpstein frame require samples with simple geometry. These are also the two methods usedby the electrical steel manufacturers. On the other hand, the ring core measurements usespecimens that are more representative of the motor geometry. The method by means ofthe Epstein frame was thoroughly discussed, since it is used in the current thesis project.

Chapter 3


The accurate estimation of the iron losses depends not only on the use of a sophisticatediron loss model, but also on the incorporation of the manufacturing effects that deterioratethe magnetic properties of the electrical steel. In this chapter, the major processes thatcause degradation of the magnetic material, according to the literature, are presented.

3.1 Introduction

The production of an electrical machine consists of different manufacturing processes.Each of these production steps induce mechanical and thermal stresses to the magneticmaterial used in the active parts of the motor. These stresses change the magnetic andthe electrical properties of the material. In this chapter, the major manufacturing effectsare presented.

Figure 3.1: Overview of manufacturing effects on iron losses in electrical machines.

3.2 Forming the Core Laminations

The electrical steel laminations used in the stator and rotor cores of electrical machinesare typically obtained by cutting through punching or laser. In the case of Epstein orSST strips, guillotine cutting is mainly used. The use of this tool emerges from thefact that the guillotine cutting is closer to the punching process which is used in massproduction of stator and rotor cores. The standards covering the Epstein measurements

13

14 CHAPTER 3. MANUFACTURING EFFECTS ON IRON LOSSES IN EM

suggest guillotine as the cutting technique for the test specimens [26]. Laser cutting ismainly used for the production of stator and rotor cores of prototype motors or small scaleproduction, since the adjustment of the punching tool into a new design has relativelyhigh cost.

3.2.1 Mechanical Cutting

The major factor for the deterioration of the magnetic properties of the ferromagneticmaterials is the cutting process [27, 28, 29]. This degradation is caused by the inducedmechanical stresses [30] during cutting. These stresses lead to an increase in the materials’specific iron losses and a drop of the relative permeability [31].

3.2.1.1 Affected Area due to Mechanical Cutting

In [32], SST measurements indicate that there is a degradation of the magnetic proper-ties of SiFe laminations in an area which can be greater than 10 mm from the cut edge.According to [33], the magnetically deteriorated area of high Si-content laminations canbe found up to 15 mm from the cut edge, while the respective distance for low Si-contentis 10 mm. Similar results are obtained in [34], where experiments on concentric ring coresare analyzed. According to these measurements, punching can create a degradation zoneup to 10-20 mm from the punched edge, while the results in [35], where same configurationis used, confirm that the degradation zone can extend up to 10 mm from the edge. In[30, 36, 37, 27, 38, 39, 40, 41, 42, 43] the lamination strips are cut in thinner pieces sothat the cutting length is increased. In [34, 35], similar experiments were conducted usingconcentric toroidal cores instead of strips. These results indicate that the degradation ofthe material depends on the punched width. In [33, 44] search coils along the lamina-tion strips are used in order to obtain the flux density at several distances from the cutedge. These measurements result in the determination of the material degradation as afunction of the absolute distance from the cut edge. Additionally, microhardness tests in[45] indicate a strain deformation up to 0.5 mm from the cutting edge, which results indegradation of the magnetic properties.

3.2.1.2 Effect of Mechanical Cutting on Hysteresis and Eddy Current Losses

The magnetic degradation due to mechanical cutting by means of guillotine or punching,causes increase of the iron losses and decrease of the relative permeability. According to[41, 46], the cutting process mainly affects the hysteresis component of the iron losses.The increase in the total iron losses can thereby be modeled by increasing the value ofthe hysteresis loss coefficient.

However, the plastic deformation after cutting also affects the eddy current losses.The reason for this is the degradation of the insulation which leads to lower apparentresistivity. Since the mechanical deformation due to punching is very limited (up to tensof micrometers) [30], the deterioration of the insulation is insignificant and the increasein the eddy current losses is very low.

3.2. FORMING THE CORE LAMINATIONS 15

3.2.1.3 Effect of Mechanical Cutting on the Magnetizing Current

The deterioration of the magnetizing properties practically means that the core needsmore magnetizing field in order to develop a certain induction level. In [47], magneticmeasurements on grid geometry are presented. This geometry consists of stacked sta-tor laminations, where only slots are punched. According to the results, almost 10%additional magnetizing current is needed to maintain the same flux under a pole.

In [34], measurements on concentric ring core specimens indicate an increase in themagnetizing current with an increased cutting length.

3.2.1.4 Si-Content

The process followed in order to produce a high silicon electrical steel is more expensiveand meticulous than with low silicon content. The manufacturing procedures in thecase of high Si-content lead to low impurities in the electrical steel and larger grain size.The silicon in the electrical steel increases the resistivity and therefore decreases the eddycurrent losses. The larger grains improve the hysteresis characteristics of the steel, leadingto lower hysteresis losses.

In [27], SST measurements are conducted in high, medium and low Si-alloyed grades.It is shown that for the same induction levels, the exciting field increases with increasingSi-content. The cutting effect is more significant for higher Si-content steels. Accordingto [6, 27], the content of Si in the steel laminations plays a major role in the degradationof the material. More specifically, an increased Si content in steel laminations, leads to ahigher increase of the exciting field and iron losses for a specific cutting length [27].

In [33], SST measurements on SiFe alloys indicate that the magnetic deteriorationin the case of high Si-content material expands up to 15 mm from the cut edge, whilethe corresponding area for low Si-content expands less than 10 mm from the cut edge.However, the authors mention that the most influential factor concerning the extent ofdeterioration is the grain size and not the Si-content.

In [39], three SiFe grades with different Si-content are tested by means of SST mea-surements. Through the experiments, the stress tensor for different directions inside thematerial is determined. The authors conclude that the stresses after mechanical cuttingare higher for high silicon laminations.

3.2.1.5 Cutting Perpendicular or Parallel to the Rolling Direction

Cutting of electrical steel laminations can be performed in parallel or transversally tothe rolling direction of the mother coil. The cutting direction has a large impact on themagnetic characteristics of the electrical steels.

In [41], the authors conduct Epstein measurements on strips that are cut parallel(RD) and perpendicular (TD) to the rolling direction. The results show that cutting haslower impact for strips that are parallel cut. Similar experimental results are presented in[27, 37]. Machine manufacturers usually assemble the stator and rotor cores by stackingcore laminations that are twisted 90o or with lower angle. This way, the anisotropy ofmagnetic properties is canceled out.

In case of segmented stator cores, there is flexibility in the orientation of the cuttingedges of the teeth and yoke. In [41], the authors suggest that the stator segments should


be cut so that the teeth, which present the highest induction, are oriented RD while theyoke which normally has lower level of induction, can be cut TD.

3.2.2 Laser Cutting

Laser stands for Light Amplification by Stimulated Emission of Radiation. Laser cuttingis a non-contact method of cutting and it is mainly used during the manufacturing ofprototypes or small-scale production motors. Laser cutting causes irreversible damage tothe magnetic characteristics of the electrical steel due to the high temperatures that aredeveloped during cutting [31].

3.2.2.1 CO2 and Fiber Laser

The laser machines are classified depending on the source of the optical gain. Two majortypes are the CO2 and the fiber laser. The CO2 laser has been commercially availablesince the 1970s and it belongs to the category of gas lasers. That means that the source ofoptical gain is a gas, usually carbon dioxide. The fiber laser is an improved version of theNd:YAG (Neodymium-doped Yttrium Aluminium Garnet) laser that has existed sincethe 1980s. The optical gain medium of a fiber laser is an optic fiber which is doped withrare earth elements. The advantage of the fiber laser is the considerably higher cuttingspeed (as high as three times) than a corresponding CO2 laser, in the case of laminationsthat are less than 4 mm thick. Moreover, the running cost of a fiber laser is up to 50%lower than that of a respective CO2 laser. Its maintenance is less expensive as well [48].

3.2.2.2 Laser Settings

The performance of the laser cutting technique is regulated by parameters like the typeof laser, the power, the cutting speed, the beam spot size, the type of assisting gas andthe gas pressure. Regulating these settings leads to different magnetic properties of themagnetic material. More information based on literature can be found in section 3.5.2.

3.2.2.3 Spatial Distribution of Degradation

SST measurements with different power, speed and gas pressure settings in [49] indicateno significant loss variations when specimens with large length are cut. This revealed thatthe magnetic degradation due to laser cutting is dependent on the geometrical shape ofthe cut sample. Similar results are obtained in [40], where the degradation is measured inthe whole width of the strips and a relation to the geometry of the samples is recognized.

Another characteristic of the laser cut laminations that reveals the nature of the spatialdistribution of the magnetic degradation is that the heat affected zone (HAZ) is dependenton the thermal history of the cutting process which means that the largest degradation isevident in the region that is cut first [40, 49].

3.2.3 Comparison between Mechanical and Laser Cutting

Punching induces shearing forces at the cut edges causing plastic deformation while lasercauses thermal stresses at the edges [40, 49].

3.3. CORE ASSEMBLY 17

In [31], Epstein measurements on 2% SiFe steel indicate 6% higher losses for lasercut laminations compared to punched ones. In [39], SST measurements on 0.31% and2.98% Si laminations indicate that laser cutting gives better results than punching. Thathappens when small samples are concerned, while X-ray analysis reveals that laser causeshigher internal stresses than mechanical cutting.

SST measurements in [49], reveal that the losses of 3% Si laminations after laser cuttingare up to 15% higher than the corresponding losses after mechanical cutting. Additionally,the same experiments show that in the case of laser cut samples, higher field strength isrequired to reach a certain level of induction. According to the same paper, the lasercutting technique provides limited possibilities of improvement of the material’s magneticproperties due to the induced thermal stresses.

The spatial distribution of the magnetic deterioration is also different for the twocutting techniques. Experiments in [40] highlight that the degradation of the mechanicalcut strips appears close to the cut edges, while the corresponding degradation for lasercut strips is evident in the total width of the strip.

Finally, when mechanical and laser cutting methods are compared, it should alwaysbe taken into consideration that the laser cutting performance is not dependent on time,while the quality of the mechanical cutting degrades with time. This is the reason whythe punching and guillotine tools require maintenance when the sharpening of the cuttingblade is discussed.

3.2.4 Abrasive Water Jet

Another method of lamination cutting which could be considered as an alternative to thelaser cutting is the abrasive waterjet cutting. SST measurements in [50] show that thiscutting technique causes very low deterioration in the magnetic properties of electricalsteels and compared to mechanical and laser cutting gives the best results. Although thisis the best cutting method, regarding the magnetic results, this technique is not widelyused due to the low speed (800 mm/min for 0.5 mm thick laminations) [50].

3.3 Core Assembly

In this section, the methods used for stacking the core laminations are presented. Afterpressing, the three main techniques for holding the stack together are welding, cleatingand gluing.

3.3.1 Pressing during Stacking

The next manufacturing step after cutting the core laminations is the pressing. Thisprocess affects both eddy current and hysteresis losses. The damage of the insulationcoating affects the eddy current losses, while the forces applied may deform the materialand therefore the magnetic properties degrade and the hysteresis losses increase [41].Measurements in [41] show that an unpressed lamination stator core has 185% lowerlosses than a pressed one. Additionally, in [6] and [51], ring core measurements indicatean increase of 400% in the change of the specific iron losses (∆pFe) between two coresthat are pressed with 1 MPa and 8 MPa respectively.


3.3.2 Welding

During the welding process, the lamination stack of the core is assembled through weldingseams in the direction of the active length of the machine. During this process, mechanicaland thermal stresses are induced and degrade the magnetic properties of the material [51].Additionally, welding causes short circuits between the laminations which decrease theeffective resistivity of the core and therefore increases eddy current losses [35].

In [51], magnetic measurements on ring core topology are performed and the resultindicate that as the number of welding seams increases there is an increase in the ironlosses and a drop in the permeability.

In [52], the welding effect is investigated on a toroidal core topology with 8 weldingseams and NO20 laminations. The results are compared to a non welded, taped core andthe outcome is that the magnetic properties of the material are significantly degraded andthe specific iron losses increased.

Finally, in [41] the authors investigate the welding effect on a stator core topology. Asreference a taped stator core is used and the studied core has 12 welding seams. It is shownthat the additional losses are decreasing with the increase of frequency and induction level,which means that the loss increase in this case is caused by the degradation of the magneticproperties of the material. It is also worth mentioning that the same study concludes thatthere is an increase in iron losses of 0.5-1% per welding seam, when stator yoke carriesmaximum flux density.

3.3.3 Cleating

Cleating is a method used for holding the lamination stack together. In this technique,metal strips are placed into slots in the periphery of the stator core. These strips are calledcleats. Once the laminations are pressed together, the two ends of the cleats are bent overso that they create a holding tab [53]. It is believed that cleating causes lower degradationthan welding. This is due to the fact that cleating does not induce any thermal stressesand it does not cause short circuits between the laminations.

3.3.4 Gluing

An alternative method of holding the lamination stack pressed together is gluing, alsocalled sticking. Gluing is mostly used in applications where light weighted stator coresare required and there is no extra material available for welding or cleating [53].

The use of this technique is described in [51], where a toroidal core is assembled. Anadhesive varnish is applied to the core laminations. Afterwards, the stack is assembledthrough a heating process. The gluing has very low or negligible effect on the magneticproperties of the material, since the varnish has non magnetic content. Therefore, anypossible degradation due to gluing is because of the thermal treatment. The results of theexperiments in [51] indicate that welding with 2 seams increases the iron losses by nearly80% compared to a glued core, while the corresponding difference of a welding with 6seams is 400%.

However, this technique has the drawback of being expensive, limiting its usage mostlyto special applications with low-weight requirements [53].

3.4. MOTOR ASSEMBLY 19

3.4 Motor Assembly

After manufacturing the stator and rotor cores, the assembly of the motor takes place.In this step of the process, the motor takes its final form.

3.4.1 Shaft Insertion

In inner rotor designs, the shaft of the electrical motors is the part that transmits thetorque from the rotor to the load.

The shaft should not move relatively to the rotor core. By heating the rotor core, itexpands and the shaft can be then inserted. As a next step, the rotor is rapidly cooleddown and the shaft is then embedded into the rotor core. This process degrades themagnetic properties of the rotor core laminations due to the thermal stresses as well asdue to the mechanical stresses when the core shrinks and applies a compressive force tothe shaft. These mechanical stresses as shown in [45], affect the magnetic properties ofthe material. Therefore the hysteresis losses are affected.

3.4.2 Pressing into Frame

Another manufacturing process is the insertion of the stator core into the frame. Thesetwo parts should be in good contact, since the frame assists in the motor cooling. Theprocess followed for this manufacturing step starts with the heating of the frame. Oncethe frame has expanded, the stator core is inserted. Afterwards, the assembled partsare cooled and the stator core is then embedded in the frame. During this process,compressive stresses from the frame to the stator core are induced. These mechanicalstresses deteriorate the magnetic properties of the ferromagnetic material. In [54], theiron losses of a surface mounted PM motor are measured before and after the insertion ina cast aluminum frame. The results indicate an increase of 10% in the iron losses betweenthe two cases. Additionally, according to [5], the degradation of the steel after insertingthe core in the frame is more significant for laminations of higher Si content.

3.4.3 Rotor Machining

The process of machining takes place mainly when induction motors are manufactured.After casting the aluminum into the rotor bars, there may be imperfections, like aluminumleftovers on the rotor’s surface. Machining removes these remnants of aluminum and it alsoensures that the rotor has the correct dimensions for the air gap requirement. Machiningis also used in other types of machines so that the air gap width is obtained with theexpected accuracy.

Machining damages the insulation of the outer part of the rotor laminations andcreates short-circuits among them. In [9, 23], magnetic measurements are performed ontwo identical CoFe stator cores. The one core is just stacked and compressed, while theother one is also glued and machined. The results indicate an increase in the eddy currentlosses for the machined core.


3.5 Manufacturing Mitigations

In this section, the processes that mitigate the material degradation due to the differentmanufacturing steps are presented. These processes either recover the magnetic propertiesof the material or they regulate the manufacturing process so that it degrades less themagnetic properties of the material. Stress relief annealing belongs to the first category,while the fine tuning of the laser settings and the maintenance of the punching machinebelong to the second one.

3.5.1 Annealing

Stress relief annealing is a technique, used to recover the magnetic properties of theferromagnetic materials after cutting.

In [35], two identical induction motors are tested, while only one of the stator cores isannealed. At rated voltage, an iron loss reduction of 15% is found. In [45], the authorsuse the annealing process to verify that the degradation of the magnetic properties dueto cutting arise from the plastic strain in the cut edges. This strain is removed throughannealing of 720oC for 2 hours and a decrease in the maximum permeability is found.

Annealing can also be performed after laser cutting. Particularly, in [31], laser cutlaminations are tested. The annealing is performed for four cases:

• Laminations just after cut

• Cut and annealed

• First annealed and then cut

• Annealed, cut and then annealed again

The hysteresis characteristic of the fourth case is the superior one with the lowestvalue of coercive field and the highest value of magnetization knee. In [23], magneticmeasurements on a CoFe stator core before and after annealing are performed. Theannealing temperature is 720oC and the duration is 2 hours. The un-annealed core has17 times higher coercive field Hc while the maximum value of induction B is 3.5 timeslower. This result highlights the necessity of annealing, when CoFe laminations are used.

In [37], annealing is applied in 1% SiFe Epstein strips for 1 hour at temperatures from450oC to 700oC. It is shown that annealing at 700oC decreases the iron losses 15 timesmore than annealing at 450oC.

The effectiveness of annealing depends on the temperature and the time. In [55], theannealing is performed on a SiFe stator core at approximately 800oC for 8 minutes. It isshown that for a flux density of 1.5 T, the iron losses of the sample decrease by 4.9%. Inthis case, annealing does not have large impact on the iron losses. One reason for thiscould be the short duration of annealing (8 minutes).

3.5.2 Tuning of Laser Settings

A method to decrease the manufacturing effect due to laser cutting is to fine tune thesettings of the laser in order to achieve lowest degradation of the electrical steel lamination.The type of the laser tool and the tuning of settings like the power, the cutting speed,

3.6. SUMMARY 21

the beam spot size, the type of assisting gas and the gas pressure have a large influenceon the deterioration of the cut magnetic material.

Measurements in [39], indicate that the pulsed mode laser with low speed providesbetter results than the continuous mode. The internal stresses after cutting with pulsedmode are higher. This can be explained by the fact that the internal stresses are translatedas effective pinning sites for the domain walls of the magnetic material. Therefore, thespeed of the domain walls is drastically decreased and the eddy current losses drop as well[39].

Finally, SST measurements in [56], between a CO2 and a fiber laser show that thespecific iron losses are almost the same when strips are cut both in RD and in TD. Anincrease of the energy input at constant power changes the relative permeability of thematerial. However, the relation of change is not linear. Best permeability characteristicscan be seen for 4kJ/m while the worst magnetic properties are evident at 24kJ/m, whichis the highest tested energy value.

3.5.3 Maintenance of Punching Machine

Similarly to the tuning of the laser settings, the maintenance of the punching machinehas a large impact on the deterioration that the cutting causes to the steel laminations.This maintenance concerns the regrinding (sharpening) of the cutting blade.

Schmidt in [37] compares the specific losses of Epstein strips for a sharp and a bluntcutting tool. When the strips are cut in the rolling direction, the blunt blade causesapproximately 7% higher iron losses than the corresponding sharpened tool. Similarstudy in [41] highlights that a newly sharpened punching tool causes up to 4% lowerlosses than the catalogue values, while a worn-out tool causes up to 6% higher losses thanthe respective catalogue values.

3.6 Summary

The manufacturing process introduces deterioration in the magnetic properties of theelectrical steel that is used in the stator and rotor core laminations. The material degra-dation consists of a reduced permeability and increased specific iron losses. Taking intoconsideration the material deterioration after the major manufacturing steps can lead tomore accurate estimation of the characteristics of the produced motor.

Chapter 4

Measurements

In this chapter, the conducted experiments regarding the cutting and welding effectsare presented. In Section 4.1, an overview of the experiments is presented. The testsetup and the repeatability of the measurements are also described. In Section 4.2, theresults from the Epstein measurements with mechanically cut laminations of M400-50A,M270-50A and NO20 are presented. Sections 4.2.1, 4.2.2 and 4.2.3, concern the resultsof the Epstein measurements on M400-50A, M270-50A and NO20, respectively, while inSection 4.2.4, a comparison of the results for the three grades is presented. Section 4.3concerns the measurements on laser cut laminations. Different laser settings are testedand the cutting effect on laser cut laminations of M270-50A is presented. In Section 4.5,experiments regarding the influence of welding are shown.

4.1 Introduction to the Experiments

The purpose of the experiments is the investigation of the cutting effect due to the twomajor cutting techniques, namely punching and laser cutting. Since the development ofa new punching tool is an expensive process, a guillotine cutting is used instead. Apartfrom the cutting effect, the influence of welding is also investigated.

This degree project was held at ABB LV Motors. The laminated materials, the guil-lotine and the measurement equipment were provided by Surahammars Bruks AB, a partof TATA Steel group, which is the second largest electrical steel manufacturer in Europe.The laser cutting was conducted at Gerdins AB, a company that specializes in compo-nents, cable systems and cutting technology. The welding was made at the factory ofABB LV Motors.

Two challenges in this project were the planning of the experiments and the timescheduling. Before starting the experimental procedure, a preliminary time plan wasmade. However, many changes to the initial plan were made, due to the consideration ofnew investigations.

4.1.1 Motivation

Studies on the cutting and welding effect have been done before, as presented in the liter-ature review (see Chapter 3). However, in most cases, a qualitative analysis is presented.The purpose of this project is to gain the absolute values of the material characteristics af-ter cutting and welding. Thus, the results can be implemented in the analytical and finite

23

24 CHAPTER 4. MEASUREMENTS

element models of electrical machines. Only few references investigate laser cutting fordifferent lamination widths and testing of a broad range of laser settings. Moreover, thesuggested method for the investigation of welding effect requires the standard equipmentfor magnetic material characterization (Epstein frame). The laminated materials that areunder investigation are selected because they are typical for electrical machines. To havea more complete investigation, laminations with the same Silicon content and differentthickness, as well as laminations with different Silicon content and the same thickness areselected.

4.1.2 Test Setup

The experiments were conducted with guillotine and laser cut laminations. For mechanicalcutting, a guillotine at Surahammars Bruks AB was used. Figure 4.1 illustrates thismachine. The guillotine is adjusted to cut standard Epstein strips (30 mm wide). Cuttingthinner strips was challenging and time consuming, because a non standardized methodof cutting should be adopted.

The laser cutting was performed at Gerdins AB by means of fiber and CO2 lasers.The laminated materials were sent from Surahammars Bruks AB to Gerdins AB. Figure4.2 depicts the fiber laser that is used at Gerdins AB.

After cutting the test specimens, their magnetic properties were measured in the Ep-stein frame provided by Surahammars Bruks AB. The frequencies used in the experimentsare 50, 100 and 200 Hz. The reason for this selection is the limitations of the measurementequipment that is used. The selected range of frequencies are representative for industrialline fed motors, where the electrical frequency is 50 or 60 Hz. A training was taken atthe factory of Surahammars Bruks AB, to learn cutting with guillotine and conductingmeasurements on the Epstein frame.

4.1.3 Repeatability

To ensure the validity and reliability of the experimental results, each measurement wasrepeated for three samples. The magnetic characteristics presented are the mean values

Figure 4.1: The guillotine at SurahammarsBruks AB.

Figure 4.2: The fiber laser machine atGerdins AB.

4.1. INTRODUCTION TO THE EXPERIMENTS 25

Figure 4.3: The Epstein frame used for themeasurements of the magnetic properties atSurahammars Bruks AB.

Figure 4.4: Zoom of the overlapping strips atthe one edge of the Epstein frame (red regionof Figure 4.3). Epstein strips with 2 addi-tional cutting edges are used in this configu-ration.

of these three measurements, unless otherwise stated. Thus the calculated values aremore representative of the mother coil properties. Figures 4.5-4.6 illustrate the relativestandard deviation of the permeability and iron loss density of the three samples in thecase of mechanically cut M400-50A 15 mm wide strips as a function of the flux density.The relative standard deviation expresses the variation of the measurements from themean value and is given by Equation 4.1.

RSD =s

x̄× 100 (4.1)

Where s and x̄ are the standard deviation and the mean value, respectively, of thesemeasurements. The standard deviation is expressed by Equation 4.2.

s =

√∑(x− x̄)2n− 1

(4.2)

Where x is the measured value and n is the number of values.The maximum values of RSD in the conducted measurements are shown in Table 4.1.

max. RSD of µr at 1 T max. RSD of pFe at 1 TSample 7.5 mm M400-50A 7.5 mm M400-50AValue 1.1 % 1.6 %

Table 4.1: Maximum max values of RSD.


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60

0.5

1

1.5

2

2.5

B (T)

RS

D o

f rel

ativ

e pe

rmea

bilty

(%

)

Figure 4.5: RSD of the relative permeabilityof mechanically cut 7.5 mm wide M00-50Astrips.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60

0.5

1

1.5

2

2.5

3

B (T)

RS

D o

f iro

n lo

ss d

ensi

ty (

%)

Figure 4.6: RSD of the iron loss densityof mechanically cut 7.5 mm wide M00-50Astrips.

4.1.4 Tested Grades

The laminated materials that are under investigation are M400-50A, M270-50A and NO20.M270-50A and NO20 contain more Silicon than M400-50A. M400-50A and M270-50A havea thickness of 0.5 mm, while NO20 is a 0.2 mm thick lamination. Table 4.2 summarizesthe basic characteristics of the tested materials.

Table 4.3 summarizes the experiments that were carried out during this project. Themagnetic measurements of non standard Epstein strips (less than 30 mm wide) requiredextra time and effort. The reason for this is that the sub-strips, which constitute astandard Epstein strip (see Figure 4.7), should be attached to each other with tape.Otherwise, the strips could not be inserted in the Epstein frame.

Si-content Resistivity Thickness

M400-50A 1.8% 42 µΩcm 0.5 mmM270-50A 3.2% 55 µΩcm 0.5 mm

NO20 3.2% 52 µΩcm 0.2 mm

Table 4.2: Overview of the tested material.

4.2. MECHANICAL CUTTING 27

M400-50A M270-50A NO20 Epstein mea-surements

Strips

Mechanical cutting for4 strip widths.

X X X 36 1680

Laser cutting for 4strip widths.

- X - 4 100

Laser cutting with 15different settings (9 ofthem measured).

X - - 16 768

Laser cutting with 3different settings.

- - X 9 648

Welding X - - 5 100

Total 70 3296

Table 4.3: Overview of experiments and number of strips.

4.2 Mechanical Cutting

According to IEC 404-2 which is the standard regarding the Epstein measurements [26],the total width of the strips under test should be 30 mm. In order to increase the cuttingeffect, the samples are cut along their length in 1/2, 1/3 and 1/4 widths. Therefore, theEpstein tests are conducted on strips whose width is 30, 15, 10 and 7.5 mm. Figure 4.7illustrates the four type of samples, as they were cut at Surahammars Bruks AB.

Figure 4.7: Schematic diagram of the 4 different configurations of the Epstein strips.Fromtop to the bottom: A standard Epstein strip (30 mm wide) and strips with one, two andthree additional cutting edges, respectively.


4.2.1 M400-50A

4.2.1.1 Cutting Effect on Iron Losses and Permeability

Figures 4.8 and 4.9 illustrate the variation of the iron losses and the relative permeabilityas a function of the strip width. For simplicity, three different induction levels at 50 Hzare selected. The quantities presented in those plots are normalized. The reference valuesare those that correspond to standard Epstein strips (30 mm wide).

In Figure 4.8, the increase of the iron losses is more significant at 0.5 T and reaches avalue of 1.35 pu at 7.5 mm. At 1.5 T, the corresponding value is 1.2 pu. The deviationof the iron losses due to cutting decreases as the induction level increases. In the caseof M400-50A, the Epstein strips of 7.5 mm width showed permanent plastic deformation(they were bent). Figure 4.9, depicts the trend of the relative permeability of M400-50Aas a function of the strip width. The permeability drops as the strip width decreases.

Figures 4.10, 4.11 illustrate the increase of the iron loss density and the reduction ofthe relative permeability as a function of the induction. The reference is the standardEpstein strip (30 mm wide).

5 10 15 20 25 301

1.05

1.1

1.15

1.2

1.25

1.3

1.35

1.4

Punched width (mm)

Nor

mal

ized

iron

loss

es

0.5 T1 T1.5 T

Figure 4.8: Normalized iron losses of M400-50A as a function of the strip width at 50 Hz.

5 10 15 20 25 300.4

0.5

0.6

0.7

0.8

0.9

1

Punched width (mm)

Nor

mal

ized

µr

0.5 T1 T1.5 T

Figure 4.9: Normalized relative permeabilityof M400-50A as a function of the strip widthat 50 Hz.


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.61

1.05

1.1

1.15

1.2

1.25

1.3

1.35

1.4

B (T)

Iron

loss

den

sity

(pu

)

15 mm M400−50A10 mm M400−50A7.5 mm M400−50A

Figure 4.10: Deviation of iron losses of M400-50A at 50 Hz, with the standard Epstein strip(30 mm wide) as reference.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6

0.4

0.5

0.6

0.7

0.8

0.9

1

B (T)

Rel

ativ

e pe

rmea

bilit

y (p

u)

15 mm M400−50A10 mm M400−50A7.5 mm M400−50A

Figure 4.11: Deviation of permeability ofM400-50A at 50 Hz, with the standard Ep-stein strip (30 mm wide) as reference.

The increase of the iron losses gets lower with increased induction. The largest de-viation in permeability between a 30 mm and a 7.5 mm wide strip is 63% at 1.3 T,corresponding to saturation knee. Table 4.4 summarizes the above results. The degra-dation of the material follows the same trend for the frequencies of 100 and 200 Hz (seeFigures A.1, A.2 and A.3 in the Appendix ).

Figure 4.12 depicts the hysteresis loop of M400-50A for strip widths of 30 and 7.5 mmat 50 Hz. The maximum applied field strength was selected to create a flux density of1.5 T. In Table 4.5, the deviations of the remanent flux density Br and coercive field Hcdue to mechanical cutting in Figure 4.12 are presented. Cutting modifies the hysteresisloop of the material. The coercive field increases, while the remanence decreases. Onereason for the drop of Br is the internal stresses in the material which causes magneticanisotropy [40]. The increase of Hc indicates an increase of the eddy current losses dueto cutting [23]. One reason for this, is the fact that cutting damages the insulation atthe edge of the material. The increase of Hc is also due to the microscopic stresses that

Iron losses Relative permeabilityInduction level ofmax. degradation

0.5 T 1.3 T

Max. degradation +35% -63%

Table 4.4: Degradation of iron losses and permeability due to mechanical cutting of M400-50A at 50 Hz.

30 mm 7.5 mm DeviationBr (T) 1.12 0.72 -36%Hc (A/m) 87.00 115.07 +32%

Table 4.5: Deviation of remanent magnetization and coercive field strength of M400-50Aat 50Hz.


−2500 −2000 −1500 −1000 −500 0 500 1000 1500 2000 2500−1.5

−1

−0.5

0

0.5

1

1.5

H (A/m)

B (

T)

30 mm7.5 mm

Figure 4.12: Hysteresis loop of M400-50A at 50 Hz.

hinder the motion of the domain walls [40].

4.2.1.2 Iron Loss Separation

The purpose of this section is to indicate how mechanical cutting changes the iron lossdistribution. The first step in this investigation is to separate the measured iron losses intohysteresis and eddy current loss components. This is achieved through surface fitting ofthe iron losses for given frequencies and induction levels (see Figure A.4 in the Appendix).The fitting equation is the separation model with two terms, which is expressed by Equa-tion 2.4. The reason for using two terms is that the iron loss model of the used FEMsoftware requires the determination of the hysteresis and eddy current loss coefficients.Moreover, the separation model with two terms is used to highlight the distribution ofiron losses to hysteresis and eddy current losses. This method is presented in [35].

Figures 4.13, 4.14 and 4.15 depict the fitting in the case of a standard Epstein strip (30mm wide) for 50, 100 and 200 Hz, respectively. The blue points indicate the experimentalvalues of iron losses while the red line corresponds to the separation model approach.


0 0.5 1 1.50

0.5

1

1.5

2

2.5

3

3.5

450 Hz

B (T)

Iron

loss

den

sity

(W

/kg)

measured valuesseparation model with 2 terms

Figure 4.13: Separation model with twoterms and measured values of iron loss den-sity at 50 Hz in the case of 30 mm wide strips.

0 0.5 1 1.50

1

2

3

4

5

6

7

8

9

10100 Hz

B (T)

Iron

loss

den

sity

(W

/kg)


Figure 4.14: Separation model with twoterms and measured values of iron loss densityat 100 Hz in the case of 30 mm wide strips.

0 0.5 1 1.50

5

10

15

20

25

30200 Hz

B (T)

Iron

loss

den

sity

(W

/kg)



The iron loss coefficients for 30, 15, 10 and 7.5 mm wide strips are summarized inTable 4.6.

Mechanical cutting has larger impact on hysteresis losses than on eddy current losses.

30 mm 15 mm 10 mm 7.5 mmkhyst 22.52 · 10−3 25.63 · 10−3 29.83 · 10−3 30.64 · 10−3kec 18.29 · 10−5 18.9 · 10−5 19.63 · 10−5 20.24 · 10−5

Deviation of khyst(%) 0 +14 +32 +36Deviation of kec(%) 0 +3 +7 +10

Table 4.6: Loss coefficients of M400-50A for different strip widths cut by guillotine. Ref-erence in the comparison is the standard 30 mm Epstein strip.


The reason for this, is the fact that cutting changes the magnetic structure of the steelmore than it changes its resistivity [35]. Therefore, the rise of the hysteresis losses is morenoticeable.

In [28, 38] it is suggested that only khyst should change when the cutting effect isincorporated in the motor design process. In this thesis, it is found that despite thelarge increase of khyst, kec also rises up to 10% for 7.5 mm wide strips. That means thatthe change of kec should be taken into consideration when the magnetic degradation isincorporated in the motor design process, as presented in see Chapter 5.

4.2.2 M270-50A


Similar investigation with M400-50A is conducted for M270-50A. Figures 4.16 and 4.17illustrate the degradation of the magnetic material as a function of the strip width. Fig-ures 4.18 and 4.19 and Table 4.7 give the deviation of the iron loss density and relativepermeability as a function of the induction.

5 10 15 20 25 301

1.05

1.1

1.15

1.2

1.25

1.3

1.35

1.4

1.45

Punched width (mm)

Nor

mal

ized

iron

loss

es

0.5 T1 T1.5 T

Figure 4.16: Normalized iron losses of M270-50A as a function of the strip width.

5 10 15 20 25 300.4

0.5

0.6

0.7

0.8

0.9

1

Punched width (mm)

Nor

mal

ized

µr

0.5 T1 T1.5 T

Figure 4.17: Normalized relative permeabilityof M270-50A as a function of the strip width.


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.61

1.1

1.2

1.3

1.4

1.5

1.6

1.7

1.8

B (T)

Iron

loss

den

sity

(pu

)

15 mm M270−50A10 mm M270−50A7.5 mm M270−50A

Figure 4.18: Deviation of iron losses of M270-50A at 50 Hz, with the standard Epstein strip(30 mm wide) as reference.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6

0.4

0.5

0.6

0.7

0.8

0.9

1

B (T)

Rel

ativ

e pe

rmea

bilit

y (p

u)

15 mm M270−50A10 mm M270−50A7.5 mm M270−50A

Figure 4.19: Deviation of permeability ofM270-50A at 50 Hz, with the standard Ep-stein strip (30 mm wide) as reference.

The deterioration of M270-50A follows the same trend as M400-50A. The trend of theiron losses and permeability is similar at 100 and 200 Hz (see Figures A.6, A.7 and A.8in Appendix).

The cutting effect on M270-50A can also be seen in Figure 4.20 through the hysteresisloops of a standard Epstein strip (30 mm) and a 7.5 mm wide strip.


0.5 T 1.3 T


Table 4.7: Degradation of iron losses and permeability due to mechanical cutting of M270-50A at 50 Hz.


30 mm 7.5 mm DeviationBr (T) 0.81 0.59 -27%Hc (A/m) 61.64 79.32 +29%

Table 4.8: Deviation of remanent magnetization and coercive field strength of M270-50Aat 50Hz.

−3000 −2000 −1000 0 1000 2000 3000−1.5

−1

−0.5

0

0.5

1

1.5

H (A/m)

B (

T)

30 mm7.5 mm

Figure 4.20: Hysteresis loop of M270-50A at 50 Hz.

Mechanical cutting changes the hysteresis loop, as in the case of M400-50A. Thedeviations of the remanence Br and the coercivity Hc are presented in Table 4.8.


Similar to M400-50A, the two-terms separation model is used for separating the total ironlosses in hysteresis and eddy current loss components. Figures 4.21, 4.22, 4.23 illustratethe fitting at 50, 100 and 200 Hz, respectively. More details about the surface fitting arepresented in the Appendix (see Figure A.9).


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60

0.5

1

1.5

2

2.5

350 Hz

B (T)

Iron

loss

den

sity

(W

/kg)



0 0.5 1 1.50

1

2

3

4

5

6

7100 Hz

B (T)

Iron

loss

den

sity

(W

/kg)



0 0.5 1 1.50

2

4

6

8

10

12

14

16

18

20200 Hz

B (T)

Iron

loss

den

sity

(W

/kg)



The hysteresis and eddy current loss coefficients of 30, 15, 10 and 7.5 mm wide M270-50A strips are summarized in Table 4.9. The trend of the two loss components is similarto M400-50A (see Section 4.2.1.2).

30 mm 15 mm 10 mm 7.5 mmkhyst 16.07 · 10−3 17.7 · 10−3 20.36 · 10−3 22.06 · 10−3kec 11.98 · 10−5 12.33 · 10−5 13.18 · 10−5 13.97 · 10−5

Deviation of khyst(%) 0 +10 +27 +37Deviation of kec(%) 0 +3 +10 +17

Table 4.9: Loss coefficients of M270-50A for different strip widths cut by guillotine. Ref-erence in the comparison is the standard 30 mm Epstein strip.


4.2.3 NO20


Similar to M400-50A and M270-50A, the influence of mechanical cutting is investigatedfor NO20, a 0.2 mm thick lamination. Figures 4.24 and 4.25 depict the trend of the ironloss density and the relative permeability of NO20 at 50 Hz as a function of the stripwidth. Figure 4.26 and 4.27 shows the deviation of the iron loss density and the relativepermeability as a function of the induction, respectively.

The deviations of the iron losses and the permeability due to mechanical cutting aresimilar to the cases of M400-50A and M270-50A (for analysis see Section 4.2.1.1).

The maximum values of degradation of NO20 are summarized in Table 4.10. Similarbehavior of the iron losses and the permeability can be observed at 100 and 200 Hz (seeFigures A.11,A.12 and A.13 in the Appendix).

5 10 15 20 25 301

1.02

1.04

1.06

1.08

1.1

1.12

1.14

1.16

1.18

1.2

Punched width (mm)

Nor

mal

ized

iron

loss

es

0.5 T1 T1.5 T

Figure 4.24: Normalized iron losses of NO20as a function of the strip width.

5 10 15 20 25 30

0.65

0.7

0.75

0.8

0.85

0.9

0.95

1

Punched width (mm)

Nor

mal

ized

µr

0.5 T1 T1.5 T

Figure 4.25: Normalized relative permeabilityof NO20 as a function of the strip width.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.61

1.05

1.1

1.15

1.2

1.25

1.3

1.35

1.4

B (T)

Iron

loss

den

sity

(pu

)

15 mm NO2010 mm NO207.5 mm NO20

Figure 4.26: Deviation of iron losses of NO20at 50 Hz, with the standard Epstein strip (30mm wide) as reference.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60.55

0.6

0.65

0.7

0.75

0.8

0.85

0.9

0.95

1

B (T)

Rel

ativ

e pe

rmea

bilit

y (p

u)

15 mm NO2010 mm NO207.5 mm NO20

Figure 4.27: Deviation of the permeability ofNO20 at 50 Hz, with the standard Epsteinstrip (30 mm wide) as reference.



0.5 T 1.3 T


Table 4.10: Degradation of iron losses and permeability due to mechanical cutting ofNO20 at 50 Hz.

−3000 −2000 −1000 0 1000 2000 3000−1.5

−1

−0.5

0

0.5

1

1.5

H (A/m)

B (

T)

30 mm7.5 mm

Figure 4.28: Hysteresis loop of NO20 at 50 Hz.

30 mm 7.5 mm DeviationBr (T) 1.05 0.78 -25.71%Hc (A/m) 43.82 50.89 +16.13%

Table 4.11: Deviation of remanent magnetization and coercive field strength of NO20 at50Hz.

The hysteresis loops of NO20 in case of a standard Epstein strip and a 7.5 mm widestrip are illustrated in Figure 4.28. The deviations of Br and Hc are given in Table 4.11.

Similarly with M400-50A and M270-50A, the cutting effect changes the hysteresisloops of NO20. Further explanation is given in section 4.2.1.


The fitting of the iron loss density for 50, 100 and 200 Hz is presented in Figures 4.29,4.30 and 4.31, respectively. Further information regarding the surface fitting is given inthe Appendix (see Figure A.14).


0 0.5 1 1.50

0.5

1

1.5

2

2.550 Hz

B (T)

Iron

loss

den

sity

(W

/kg)



0 0.5 1 1.50

0.5

1

1.5

2

2.5

3

3.5

4

4.5

5100 Hz

B (T)

Iron

loss

den

sity

(W

/kg)



0 0.5 1 1.50

2

4

6

8

10

12200 Hz

B (T)

Iron

loss

den

sity

(W

/kg)



manufacturing effects on iron losses in electrical machines847378/fulltext01.pdf · manufacturing e...

Documents