Manufacturing Effects on Iron Losses in Electrical Machines847378/... · 2015-08-20 · Manufacturing E ects on Iron Losses in Electrical Machines Konstantinos Bourchas Master of

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

2


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

[W

kg·(T ·Hz)1.5]

khyst hysteresis loss coefficient[

Wkg·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-plat paverkasav skarning och svetsning. Permeabilitet och jarnforlustdensitet undersoks eftersom dear kritiska variabler for elektriska maskiners prestanda. De magnetiska matningarnagenomfordes pa en Epstein ram med en flodesfrekvens pa 50, 100 och 200 Hz, enligtIEC 404-2. Effekterna av mekanisk skarning med giljotin samt skarning med fiber- ochCO2-laser studerades. Inverkan av olika fiberlaserinstallningar undersoktes ocksa genomatt variera gastrycket, skarhastigheten samt frekvensen och effekten av laserstralen.

For att oka skareffekten inkluderades Epsteinremsor med ytterligare 1, 2 och 3 langsgaendeskarsnitt. Det visas att mekanisk skarning har en mindre paverkan pa de magnetiskaegenskaperna hos materialet an vad laserskarning har. Matningar pa plat med 1.8% Sivisar att da prov med tre extra langsgaende giljotinklipp anvands kan permeabilitetenreduceras med upp till 63% och jarnforlusterna kan oka med upp till 35%. Motsvaranderesultat for laserskurna platar visar en permeabilitetsreduktion pa upp till 65% ochen jarnforlustokning pa upp till 65%. Ur studien av de tva studerade skarprocessernaframkommer aven att tunnare plat paverkas mindre negativt an tjockare plat. Ett antalolika installningar har provats for att utreda hur olika parametrar paverkar effekternaav laserskarning. Studien indikerar att skarning med hog effekt och hog hastighet gerden minsta paverkan pa materialets magnetiska egenskaper. Vilket aven har en positivinverkan pa produktiviteten vid laserskarning.

Epsteinprover har aven utforts for att undersoka vilka effekter som introduceras daSiFe-plat svetsas. Provstyckena bestod av remsor med en, tre, fem och 10 svetspunkter.Experimenten visar en jarnforlustokning med upp till 50% samt en permeabilitetsreduk-tion upp till 62% da platarna svetsats samman tva och tva.

En modell for att studera effekterna av de forandrade materialegenskaperna vid skarningpa en induktionsmotor utvecklas och implementeras i en FEM-baserad mjukvara. Resul-taten tyder pa en jarnforlustokning med 30% da skareffekten orsakad av giljotin beaktas.Vid simulering av laserskuren plat kan denna okning vara sa stor som 50%. Det framkom-mer aven att laserskarningen kan reducera effektfaktorn sa mycket som 2.6%.

Nyckelord - elektrisk maskin, induktion motor, jarnforluster, relativ per-meabilitet, giljotin, fiberlaser, CO2-laser, laser installningar, skareffekt, svet-sning, elektromagnetisk plat

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 Vasteras,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 Vasteras, 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 = kSEfαBβ (2.1)

where pFe are the specific iron losses (W/kg), f is the frequency and B 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 Bβf (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

∫ T

0

(dB

dt

)2

dt (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 = khystfB2 + kecf

2B2 (2.4)

Pry and Bean[19]

pFe = physt + ηapec =

khystfB2 + ηexckecf

2B2(2.5)

Bertotti[4]

pFe = physt + pec + pexc

= khystfB2 + kecf

2B2 + kexcf1.5B1.5

(2.6)

Jacobs[20]

pFe = k′hystfB2 + (kec + a1B

a2)B2f 2 (2.7)

where k′hyst = khyst(1 + Bmin

Bmax(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 (Figure2.4).

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 =N1

N2

Pm −(1.111|U2|)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.

• |U2| 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) =N1

lmI1(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) = − 1

N2A

∫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

x× 100 (4.1)

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

s =

√∑(x− x)2

n− 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 Hc

due 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 · 10−3

kec 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.



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 · 10−3

kec 11.98 · 10−5 12.33 · 10−5 13.18 · 10−5 13.97 · 10−5


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)



Similarly to the case of M400-50A and M270-50A, the hysteresis losses are affectedmore than the eddy current losses. This happens because the magnetic structure isaffected more than the resistivity of the material [35].

30 mm 15 mm 10 mm 7.5 mmkhyst 16.69 · 10−3 17.47 · 10−3 18.53 · 10−3 19.75 · 10−3

kec 3.49 · 10−5 3.63 · 10−5 3.78 · 10−5 3.92 · 10−5


Table 4.12: Loss coefficients of NO20 for different strip widths cut by guillotine. Referencein the comparison is the standard 30 mm Epstein strip.


4.2.4 Comparison

In this section, the three laminated materials are compared in terms of their degradationdue to mechanical cutting. Table 4.13 gives the iron loss increase due to one additionalcutting edge at three frequencies.

Grade 50 Hz 100 Hz 200 HzM400-50A 10.54% 10.38% 8.98%M270-50A 8.06% 7.10% 6.04%

NO20 4.49% 4.47% 4.43%

Table 4.13: Increase of iron loss density for the 15 mm wide strips relative to the 30 mmwide strips at 1 T.

As frequency increases, the iron loss density increases with lower gradient, as shownin Figure 4.32.

The increase of the iron loss density of the three laminated materials is frequencydependent. This shows that mechanical cutting affects not only the hysteresis losses, butalso the eddy current losses. NO20, the thinnest lamination, has the lowest dependencyon the frequency. The degradation of M270-50A has the greatest dependency on thefrequency. One reason for this is that M270-50A has higher Silicon content than M400-50A. Therefore, at higher frequencies, the increase of eddy current losses is more significantfor M400-50A than for M270-50A. This means that cutting, which mainly affects thehysteresis losses, is more pronounced in the case of M270-50A.

Figures 4.33, 4.34 and 4.35 illustrate the trend of the iron losses of the three laminatedmaterials for different widths and inductions. NO20 shows the lowest degradation. Thismeans that as the lamination thickness decreases, the degradation of the material is lower.

50 100 150 2000.7

0.75

0.8

0.85

0.9

0.95

1

Frequency (Hz)

Dev

iatio

n of

iron

loss

den

sity

(pu

)

M400−50AM270−50ANO20

Figure 4.32: Normalized increase of the iron loss density between a standard Epsteinstrip (30 mm) and a 15 mm wide strip at 50, 100 and 200 Hz at 1 T. Base value is thecorresponding increase at 50 Hz.


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.61.02

1.04

1.06

1.08

1.1

1.12

1.14

1.16

1.18

B (T)

Iron

loss

den

sity

(pu

)

15 mm M400−50A15 mm M270−50A15 mm NO20

Figure 4.33: Iron loss density of 15 mm widelaminations as a function of B at 50 Hz, withthe standard Epstein strip (30 mm wide) asreference.

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

B (T)

Iron

loss

den

sity

(pu

)

10 mm M400−50A10 mm M270−50A10 mm NO20

Figure 4.34: Iron loss density of 10 mm widelaminations as a function of B at 50 Hz, withthe standard Epstein strip (30 mm wide) asreference.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.61.1

1.2

1.3

1.4

1.5

1.6

1.7

1.8

1.9

B (T)

Iron

loss

den

sity

(pu

)

7.5 mm M400−50A7.5 mm M270−50A7.5 mm NO20

Figure 4.35: Iron loss density of 7.5 mm widelaminations as a function of B at 50 Hz, withthe standard Epstein strip (30 mm wide) asreference.

Similar results are found in [27]. Among the laminations with the same thickness, theone with high Si-content (M270-50A) shows higher degradation at low induction. Thisresult is different than what is found in [27, 39]. In those references it was found thathigher Si-content leads to larger degradation, independently on the induction. In thisstudy, however, it is found that this holds only for induction lower than 0.5 T.

4.2.5 Summary

In this section, the effect of mechanical cutting is investigated. The most importantresults of this experimental study are the following.

• The magnetic degradation of the SiFe laminations increases as their width decreases.

4.3. LASER CUTTING 41

That means that the iron loss density increases and the relative permeability drops.

• NO20 shows the lowest degradation. As the lamination thickness decreases, thedegradation of the material gets lower.

• The lamination with high Si-content (M270-50A) shows the largest degradation atlow induction.

• The degradation of the iron loss density drops as the induction increases. Themaximum degradation of the relative permeability is found at 1.3 T, correspondingto the saturation knee.

• Mechanical cutting affects both hysteresis and eddy current losses. That can bemodeled though increased values of khyst and kec in the separation iron loss model.The increase of the hysteresis loss component is more significant.

4.3 Laser Cutting

Laser cutting deteriorates the magnetic properties of the electrical steel by the inductionof thermal stresses, which in turn change the texture of the material [39, 49, 56]. In thischapter, the investigation of the magnetic degradation due to this cutting technique ispresented.

4.3.1 Comparison among different Laser Settings

The first step in this study is to investigate the trend of the iron losses and the permeabilitywhen the magnetic material is cut with different laser settings or different types of laser.

First, experiments were conducted for M400-50A and NO20 laminations. Laser cuttingwith settings Set 7, Set 8 and CO2 was performed (see Table 4.14). After these firstmeasurements, it was decided that more laser settings should be tested so that a betteroverview of the influence of the laser settings could be gained. The next experiments wereconducted only on M400-50A laminations. The test specimens are Epstein strips with 2additional cutting edges, i.e 3 adjacent strips of 10 mm width. The reason for this, is toincrease the cutting effect, keeping the number of required strips in a reasonable range.In this section, only the results of M400-50A laminations are presented. The influence ofdifferent laser settings on NO20 are presented in the Appendix (see Figures B.1, B.2).

4.3.1.1 Selection of Laser Settings and Laser Machine

In [39], two laser settings of CO2 laser with different power and speed are studied. In[49], eight settings with varying power, speed and assisting gas pressure are tested onCO2 laser as well. Two different laser types (CO2 and Yb:YAG) are compared in [56].In this project, cutting with fiber and CO2 laser is performed. Particularly, fourteensettings of fiber laser and one setting of CO2 laser are tested. However, only eight chosensettings of fiber laser could cut the strips. Each setting consists of four parameters: thepower, cutting speed, frequency and pressure of the assisting gas. Thus, compared to theliterature, this study takes into consideration both the laser type and the different settings.Additionally, not only the power, speed and gas pressure, but the influence of frequency


is also investigated. Table 4.14 summarizes the settings with which the strips were cutsuccessfully. The experimental results indicated the influence of different parametersto magnetic materials. Thus, the values of parameters (power, speed, frequency andpressure) that would result in the best magnetic properties were estimated. However, asetting with these best parameters was not tested. This is due to the fact that all thetested settings were selected before the conduction of the experiments.

Settings Power (pu) Cuttingspeed (pu)

Freq. (pu) Nitrogen gaspressure (pu)

Set 1 0.4 1 0.4 0.5Set 2 1 0.02 0.4 0.5Set 3 0.4 1 0.16 0.5Set 4 0.4 1 0.28 0.5Set 5 0.4 1 0.4 1Set 6 0.4 1 0.01 0.5Set 7 0.08 0.1 0.4 0.5Set 8 1 1 0.4 0.5CO2 0.3 0.14 1 0.65

Table 4.14: Laser settings of M400-50A

The assisting gas is nitrogen. This gas type is preferable compared to oxygen, to avoidoxidation [49]. The values in Table 4.14 are given in per-unit system. The base value ofeach parameter corresponds to the maximum tested value.

4.3.1.2 Degradation of M400-50A due to Laser Cutting with Various Settings

Figures 4.36-4.37 illustrate the permeability and the iron loss density of M400-50A, aftercutting with the settings shown in Table 4.14.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6600

800

1000

1200

1400

1600

1800

2000

2200

2400

2600

B(T)

Rel

ativ

e pe

rmea

bilit

y

Set 1Set 2Set 3Set 4Set 5Set 6Set 7Set 8CO

2

Figure 4.36: Relative permeability of 10mmM400-50A at 50 Hz.

0 0.5 1 1.50

1

2

3

4

5

6

B(T)

Iron

loss

den

sity

(W

/kg)

Set 1Set 2Set 3Set 4Set 5Set 6Set 7Set 8CO

2

Figure 4.37: Iron losses of 10mm M400-50Aat 50 Hz.

Set 8, the laser setting with highest power/highest speed, results in the best magneticcharacteristics. Due to the high speed, the productivity of the core laminations can


increase while keeping the magnetic degradation in the lowest possible level (lowest amongthe tested settings).

On the other hand, Set 2 (highest power/lowest speed), introduces the maximumdegradation. Fiber laser is superior to CO2 laser regarding the magnetic properties ofthe cut material. The only setting of fiber that results in worse magnetic properties thanCO2 laser, is Set 2. Table 4.15 summarizes the differences between Set 8 and Set 2 whichconstitute the best and the worst settings that were tested.

0.1 T 0.5 T 1 T 1.5 T max degr.Deviation of µr -11% -45% -50% -17% -50% @ 0.8TDeviation of pFe +21% +36% +26% +15% +36% @0.5T

Table 4.15: Deviation of relative permeability and iron loss density between Set 8 andSet 2 laser cut M400-50A at 50 Hz. Results from Set 8 settings are used as reference.

Influence of assisting gas pressureIncrease of the gas pressure from 0.5 to 1 pu (Set 1 - Set 5 ) leads to higher relativepermeability and lower iron loss density. At 1 T and 50 Hz, the permeability increase is10% while the iron loss density drops by 4%.

Influence of laser beam frequencyThe influence of the laser beam frequency is investigated by keeping the other parametersconstant (Set 1, Set 3, Set 4, Set 6 ). Figure 4.38 illustrates the trend of the relativepermeability and the iron loss density of M400-50A under these settings.

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.41840

1860

1880

1900

1920

1940

f (pu)

Rel

ativ

e pe

rmea

bilit

y

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.42.32

2.34

2.36

2.38

2.4

2.42

p Fe (

W/k

g)

µr

pFe

Figure 4.38: Dependency of relative permeability and iron loss density on laser beamfrequency. Values are considered at 1 T and 50 Hz.

The best magnetic properties among Set 1, Set 3, Set 4 and Set 6 are presented at afrequency of 0.28 pu. At a frequency of 0.4 pu, the permeability and the iron loss densityshow their maximum degradation. The result of this study is that increasing frequency


up to a certain value leads to better magnetic properties. After this value, the materialis heavily deteriorated. In this case, this critical frequency lies between 0.28 and 0.4 pu.

Influence of cutting speedThe conducted experiments show that the degradation of the material increases as thecutting speed drops. One reason for this could be that the material is thermally penetratedfor longer time.

Influence of laser powerComparison between Set 1 and Set 8 shows that the material has higher magnetic degra-dation at lower laser power. This finding is different than what is found in [56], wherelower power leads to higher permeability. One reason for this could be that the geometryof the lamination affects the degradation, as shown in [49]. In [56], 5 mm wide strips areused, while in the current investigation 10 mm wide strips are tested.

Combinations of laser power and cutting speedThe dependency of the iron loss density on the laser power and cutting speed is illustratedin Figure 4.39. Set 1, Set 2, Set 7 and Set 8 are considered, where gas pressure andlaser beam frequency are kept constant (gas pressure of 0.5 pu and frequency of 0.4 pu,respectively). A certain trend of the iron loss density with varying power and speed cannotbe identified. The reason for this could be that the four points are not representative ofthe entire range of laser powers and cutting speeds.

Failing laser settingsThe laser settings that were tested, but could not cut the Epstein strips, are summarizedin Table 4.16. A corresponding photograph can be seen in Figure 4.40. With Set 14, thestrips could be cut, but the edge of the strip was melt due to the high thermal penetration.Additionally, the burr height in this case was very high.

The effect of the laser parameters on the material degradation, according to the ex-

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 12

2.1

2.2

2.3

2.4

2.5

2.6

2.7

2.8

Iron

loss

den

sity

(W

/kg)

Speed (pu)

Power= 0.4 pu

Power= 0.08 pu

Power= 1 pu

Power= 1 pu

Figure 4.39: Iron loss density at 1 T and 50 Hz for varying cutting speed and laser power.


Setting Power (pu) Cuttingspeed(pu)

Freq.(pu)

Nitrogengas pressure(pu)

Set 9 0.2 1 0.4 0.5Set 10 0.3 1 0.4 0.5Set 11 0.35 1 0.4 0.5Set 12 0.37 1 0.4 0.5Set 13 0.4 1 0.28 0.5Set 14 1 0.002 0.4 0.5

Table 4.16: Laser settings of M400-50A

Figure 4.40: Strips that have not been cut due to inappropriate laser beam settings.

periments, is shown in Table 4.17.

Parameter Permeability Iron lossesCutting speed Laser power Gas pressure Frequency (untila critical value)

Table 4.17: Effect of cutting speed, gas pressure and frequency on degradation of themagnetic properties of M400-50A.

4.3.2 Cutting effect due to laser

The cutting effect due to laser is investigated for M270-50A laminations. Similar to thecase of the mechanical cutting, the strips were cut with one, two and three additional cutedges as depicted in Figure 4.7. The two laser settings that were used are Set 8 and Set 2


which constitute the settings with the best and the worst result concerning the magneticproperties of the cut material. This investigation is significant, because there is not anysimilar study found in literature (cutting effect is determined only for mechanically cutlaminations).

4.3.2.1 Cutting Effect due to Best Laser Setting (Set 8 )

The degradation of M270-50A due to cutting with Set 8 is given in Figures 4.41 and4.42. The increase of the iron loss density is more significant for low induction. This issimilar to mechanical cutting. The iron loss density of a 7.5 mm wide strip at 0.5 T is65% higher than the respective 30 mm wide strip. At 1.5 T, the corresponding increaseis 22%. According to Figure 4.42, the permeability drops as the strip width decreases.

Figures 4.43 and 4.44 illustrate the deviation of the iron loss density and the relativepermeability at different flux densities. Reference is the standard Epstein strip (30 mmwide).

5 10 15 20 25 301

1.1

1.2

1.3

1.4

1.5

1.6

1.7

1.8

Strip width (mm)

Iron

loss

den

sity

(pu

)

0.5 T1 T1.5 T

Figure 4.41: Deviation of the iron loss densityof M270-50A as a function of the strip widthfor laser cutting with Set 8 and frequency of50 Hz, with the standard Epstein strip (30mm wide) as reference.

5 10 15 20 25 30

0.4

0.5

0.6

0.7

0.8

0.9

1

Strip width (mm)

Per

mea

bilit

y (p

u)

0.5 T1 T1.5 T

Figure 4.42: Deviation of the relative perme-ability of M270-50A as a function of the stripwidth for laser cutting with Set 8 and fre-quency of 50 Hz, with the standard Epsteinstrip (30 mm wide) as reference.


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.81

1.1

1.2

1.3

1.4

1.5

1.6

1.7

1.8

1.9

B (T)

Iron

loss

den

sity

(pu

)

15 mm10 mm7.5 mm


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8

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 mm10 mm7.5 mm


The permeability is deteriorated mostly in the region 0.5-1 T, as summarized in Table4.18 and shown in Figure 4.44. This finding is different from the case of mechanicalcutting, where the permeability degrades mostly in the knee of saturation.


0.5 T 0.5 T - 1 T


Table 4.18: Degradation of iron losses and permeability due to laser cutting with Set 8of M270-50A at 50 Hz.

The hysteresis loop of the laser cut material is deformed compared to the mechanicallycut one, as shown in Figure 4.45. This deformation can be seen around the origin of theaxis, i.e for low values of induction and magnetizing field.


−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 4.45: Hysteresis loop of a 30 and a 7.5 mm wide M270-50A strip. Laser cuttingwith Set 8 is performed. Maximum induction is 1.5 T at 50 Hz.

Laser cutting with Set 8 changes the hysteresis loop. The deviations of the remanenceBr and the coercivity Hc are presented in Table 4.19.


Table 4.19: Deviation of remanent magnetization and coercive field strength of M270-50Aat 50Hz. Laser cutting with Set 8 is performed.

Separation of iron lossesSurface fitting of the iron loss density to separation model with two terms indicates howlaser cutting with Set 8 changes the loss distribution. The fitting model is given byEquation 2.4. The hysteresis and eddy current loss coefficients as resulted from thisfitting for 50, 100 and 200 Hz are summarized in Table 4.20.

30 mm 15 mm 10 mm 7.5 mmkhyst 19.57 · 10−3 21.87 · 10−3 25.63 · 10−3 27.52 · 10−3

kec 11.78 · 10−5 12.52 · 10−5 13.51 · 10−5 14.42 · 10−5


Table 4.20: Loss coefficients of M270-50A for different strip widths when cut by laser withSet 8. Reference in the comparison is the standard 30 mm Epstein strip which is cut withthe same setting.

The hysteresis losses are affected more than the eddy current losses. This means thatthe magnetic structure of the material is affected more than its resistivity. khyst has an


increase of up to 36%, while the corresponding increase of kec is 21%. Similar to the caseof mechanical cutting, the increase of the iron loss coefficients can be used in the motormodels to incorporate the iron loss increase due to cutting, as described in Chapter 5.

4.3.2.2 Cutting Effect due to Worst Laser Setting (Set 2 )

Figures 4.46, 4.47 illustrate the degradation of M270-50A due to laser cutting with Set 2.The trend of the iron loss density and the permeability is similar to the case of cuttingwith Set 8.

5 10 15 20 25 301

1.1

1.2

1.3

1.4

1.5

1.6

1.7

Strip width (mm)

Iron

loss

den

sity

(pu

)

0.5 T1 T1.5 T


5 10 15 20 25 30

0.4

0.5

0.6

0.7

0.8

0.9

1

Strip width (mm)

Per

mea

bilit

y (p

u)

0.5 T1 T1.5 T


The iron loss density increases with reduced strip width. The degradation of the ironlosses is larger for low induction levels, as shown in Figures 4.46 and 4.48. At 0.5 T and50 Hz, the iron loss increase of a 7.5 mm wide strip is 53%. The corresponding increaseat 1.5 T is 23%. According to Figures 4.47 and 4.49, the permeability drops as the stripwidth decreases. The permeability of a 7.5 mm wide strip at 50 Hz is reduced by 67%compared to a standard 30 mm wide Epstein strip. Similar to Set 8, the permeabilityshows maximum degradation in the range 0.5-1 T.


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.81

1.1

1.2

1.3

1.4

1.5

1.6

1.7

B (T)

Iron

loss

den

sity

(pu

)

15 mm10 mm7.5 mm


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8

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 mm10 mm7.5 mm


Similar to Set 8, the hysteresis loops in the case of Set 2 are also deformed, as shownin Figure 4.50. The deviations of Br and Hc are presented in Table 4.21.

−4000 −3000 −2000 −1000 0 1000 2000 3000 4000−1.5

−1

−0.5

0

0.5

1

1.5

H (A/m)

B (

T)

30 mm7.5 mm

Figure 4.50: Hysteresis loop of a 30 and a 7.5 mm wide M270-50A strip. Laser cut withSet 2 is performed. Maximum induction is 1.5 T at 50 Hz.



Table 4.21: Deviation of remanent magnetization and coercive field strength of M270-50Aat 50Hz. Laser cutting with Set 2 is performed.

The drop of the Br and the increase of Hc are more significant when cutting with Set8, as shown in Tables 4.19 and 4.21.

Iron loss separationSimilar to Set 8, the iron loss coefficients of M270-50A in the case of laser cutting withSet 2 are summarized in Table 4.22. The increase of the hysteresis and eddy current losscoefficients is very close. One reason for that could be the fact that high-power/low-speedlaser settings damage the insulation more than a high-power/high-speed settings.

30 mm 15 mm 10 mm 7.5 mmkhyst 23.23 · 10−3 27.53 · 10−3 29.62 · 10−3 30.69 · 10−3

kec 12.78 · 10−5 14.46 · 10−5 16.06 · 10−5 17.02 · 10−5


Table 4.22: Loss coefficients of M270-50A for different strip widths when cut by laser withSet 2. Reference in the comparison is the standard 30 mm Epstein strip which is cut withthe same setting.

4.3.3 Summary

The main results of the experiments on laser cutting are summarized as follows.

• Fiber laser leads to better magnetic properties than CO2 laser.

• Increased speed of the laser beam leads to better magnetic characteristics. Thisresult can be used to increase the productivity of manufacturing of core laminations.

• Increased gas pressure has positive impact on the magnetic properties of the mate-rial.

• Increased frequency results in lower degradation. That changes after a critical fre-quency. Then the magnetic material shows steep deterioration.

• Laser cutting causes drop of Br and rise of Hc. The best laser setting (Set 8 ), whichresults in the highest permeability and lowest iron losses, shows larger deviation ofBr and Hc than the worst laser setting (Set 2 ) .

• Laser cutting increases khyst and kec. That can be used for modeling the effect oflaser cutting on the iron losses.


4.4 Comparison between Mechanical and Laser Cut-

ting

4.4.1 Introduction

In this investigation, M400-50A, M270-50A and N020 laminations are considered. Forthis comparative study, laser cutting is performed with Set 8, which is the setting withhighest-power/highest-speed of fiber laser. The reason for this is that this laser settinghas the minimum impact on the magnetic properties of the steel laminations, as shown inSection 4.3. The high-power/high-speed laser cutting is also used for NO20 laminations.Thus, this study concerns the comparison between mechanical and the best possible lasercutting (best among fourteen laser settings).

4.4.2 M400-50A

Figures 4.51 and 4.52 depict the permeability and the iron losses respectively of a 10 mmwide M400-50A strip.

The permeability of a mechanically cut strip is up to 40% higher compared to thecorresponding laser cut strip. For induction levels that exceed 1.3 T, the laser cut striphas higher permeability and at 1.5 T, this difference reaches 26 %. The iron losses of themechanically cut strips are lower than the respective losses of the laser cut laminations,as shown in Figure 4.52. The difference in the iron losses drops as the induction levelincreases. Table 4.23 summarizes the above results.

0.1 T 0.5 T 1 T 1.5 TDeviation of µr -29% -39% -26% +26%Deviation of pFe +46% +17% +8% 0%

Table 4.23: Deviation of the relative permeability and the iron losses between laser andguillotine cut M400-50A strips of 10 mm width at 50 Hz. Guillotine cut laminations aretaken as reference.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6500

1000

1500

2000

2500

3000

3500

4000

4500

Rel

ativ

e pe

rmea

bilit

y

B (T)

GuillotineLaser Set 8

Figure 4.51: Relative permeability of 10 mmwide M400-50A strip at 50 Hz.

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

3.5

4

4.5

Iron

loss

den

sity

(W

/kg)

B (T)


Figure 4.52: Iron losses of 10 mm wide M400-50A strip at 50 Hz.

4.4. COMPARISON BETWEEN MECHANICAL AND LASER CUTTING 53

4.4.3 M270-50A

Figures 4.53 and 4.54 illustrate the permeability and the iron loss curves of the materialfor the two different cutting techniques.

As Figure 4.53 illustrates, the permeability of the material is up to 60% higher in thecase of mechanically cut laminations. Moreover, the iron loss characteristics of mechani-cally cut laminations are also better. However this difference drops as the induction levelincreases. The deviation of the relative permeability and iron losses are summarized inTable 4.24.

0.1 T 0.5 T 1 T 1.5 TDeviation of µr -49% -60% -46% +21%Deviation of pFe +75% +51% +23% +8%

Table 4.24: Deviation of the relative permeability and the iron losses between guillotineand laser cut M270-50A strips of 10mm width at 50 Hz. Guillotine cut laminations aretaken as a reference.

4.4.4 NO20

Figures 4.55 and 4.56 illustrate the permeability and the iron loss curves of the materialfor the two different cutting techniques.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60

1000

2000

3000

4000

5000

6000

B (T)

Rel

ativ

e pe

rmea

bilit

y


Figure 4.53: Relative permeability of 10mmM270-50A strips at 50 Hz.

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

3.5

B (T)

Iron

loss

den

sity

(W

/kg)


Figure 4.54: Iron losses of 10mm M270-50Astrips at 50 Hz.


0 0.5 1 1.50

1000

2000

3000

4000

5000

6000

7000

8000

Rel

ativ

e pe

rmea

bilit

y

B (T)


Figure 4.55: Relative permeability of 10mmNO20 strips at 50 Hz.

0 0.5 1 1.50

0.5

1

1.5

2

2.5

3

Iron

loss

den

sity

(W

/kg)

B (T)


Figure 4.56: Iron losses of 10mm NO20 stripsat 50 Hz.

As Figure 4.55 illustrates, the permeability of the material is up to 60% higher in thecase of mechanically cut laminations. Moreover, the iron loss characteristics of mechani-cally cut laminations are also better. However this difference drops as the induction levelincreases. The deviation of the relative permeability and iron losses are summarized inTable 4.25.

0.1 T 0.5 T 1 T 1.5 TDeviation of µr -57% -62% -52% 0%Deviation of pFe +146% +71% +40% +15%

Table 4.25: Deviation of the relative permeability and the iron losses between guillotineand laser cut NO20 strips of 10mm width at 50 Hz. Guillotine cut laminations are takenas a reference.

4.4.5 Summary

The comparison between mechanical and laser cutting can be summarized as follows.

• Laser cutting leads to higher iron losses than mechanical cutting.

• Laser cutting shows larger degradation of the magnetizing properties of the material.This is seen as a reduced permeability.

• At high induction, the differences between the two cutting techniques are reduced.

• The degradation of the permeability in the case of mechanical cutting is most sig-nificant in the region which is close to the knee of saturation. In the case of lasercut laminations, the corresponding degradation expands in a larger region of fluxdensities and specifically between 0.5 and 1 T.

• The shape of the hysteresis loop is different in the case of laser cut laminations thanfor mechanically cut ones.

4.5. WELDING 55

4.5 Welding

The welding effect is investigated based on Epstein measurements. The test specimens arelaser cut (Set 8) standard Epstein strips (30 mm wide) of M400-50A laminations with 1,3, 5 and 10 welding seams as depicted in Figure 4.57. Welding is used to join two strips.As reference, a non welded strip is used. All the specimens have 10 welding sockets,where the welding can be applied. The Epstein measurements use the same number ofstrips (16), as suggested by the standard [26]. Welding has also been investigated inliterature [6, 51, 52]. However, in those references, measurements on ring core topologyare conducted. In this project, a new method using Epstein measurements is proposed.

4.5.1 Measurement Results

The normalized iron loss density and relative permeability at 50 Hz can be seen at Figures4.58 and 4.59.

Figure 4.57: Samples used for the investigation of the welding effect. Red spots indicatethe welding seams.


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.81

1.1

1.2

1.3

1.4

1.5

1.6

1.7

B (T)

Iron

loss

den

sity

(pu

)

1 welding point3 welding points5 welding points10 welding points

Figure 4.58: Iron loss density of weldedM400-50A laminations at 50 Hz. Referenceis the non welded strip.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8

0.4

0.5

0.6

0.7

0.8

0.9

1

1.1

B (T)

Rel

ativ

e pe

rmea

bilit

y (p

u)

1 welding point3 welding points5 welding points10 welding points

Figure 4.59: Relative permeability of weldedM400-50A laminations at 50 Hz. Reference isthe non welded strip.

Table 4.26 summarizes the increase of the iron losses and the reduction of the relativepermeability at 1 T and 50 Hz.

1 weld. seam 3 weld. seams 5 weld. seams 10 weld. seamsDeviation of pFe +2% +10% +19% +50%Deviation of µr -14% -34% -50% -62%

Table 4.26: Deviation of iron loss density and relative permeability due to welding at 1T and 50 Hz. Reference is the non-welded material.

In the case of ten welding seams, both sides of the strip are welded. This introducesextra eddy currents in addition to the effect of having more welds. Welding seams at oneside of the flux path can introduces extra losses due to stresses and local eddy currents.However, welding at both sides of the flux path will introduce some extra eddy currentlosses in the total volume of strips between the welds.

4.5.2 Summary

The main results of the conducted experiments on welding are summarized as follows.

• A new method for investigating the welding effect is presented. This method isbased on Epstein measurements and no ring core geometry is required.

• An iron loss increase up to 50% is measured. The respective decrease of permeabilityis 62%, corresponding to 10 welding seams.

Chapter 5

Simulations

Iron loss models can show significantly improved accuracy if they include correction coef-ficients, which result from many years of experience gathered after the construction andevaluation of prototype motors [34]. However, this method is difficult to apply for dif-ferent motor geometries, since correction factors do not take into consideration machinegeometry and elevated level of the flux density in certain parts of the motor [28].

Incorporating manufacturing effects into the FEM model of the machine leads to theconsideration of the additional magnetomotive force (MMF), which is required in orderto develop a certain flux, while the geometry is also taken into account [28].

In this section, the experimental results regarding the degradation of the electricalsteel due to mechanical and laser cutting are considered in the finite element model of astandard induction motor.

5.1 Separation in Yoke and Teeth Regions

The proposed model is based on the approach that is discussed in [34], where the stator isseparated in yoke and teeth region. Material properties of ring cores with correspondingwidth are applied. In the model that is suggested in this thesis, both the stator and rotorare separated in yoke and teeth regions for an induction motor, as illustrated in Figure5.1.

57

58 CHAPTER 5. SIMULATIONS

Figure 5.1: Illustration of the FEM model of the induction motor which includes thedegradation profile.

The suggested approach is implemented in a FEM model of the motor. The softwarethat is used for the FEM modeling and simulations is Adept, which includes a 2D-FEMpackage.

The magnetic material properties that are assigned in each region are based on Epsteinmeasurements of strips with different widths. However, the Epstein measurements areconducted on 30, 15, 10 and 7.5 mm wide strips. If a region has different width, thenlinear interpolation is applied between the BH values of the two adjacent widths. In thiscase, the iron loss coefficients are determined by an equation that gives khyst and kec forany width. This equation is determined by curve fitting the calculated iron loss coefficientsfor 30, 15, 10 and 7.5 mm. Figures 5.2 and 5.3 depict this curve fitting approach in thecase of mechanically cut M270-50A laminations.

5.2. MODEL FOR PERMEABILITY AT HIGH FLUX DENSITIES 59

5 10 15 20 25 300.016

0.017

0.018

0.019

0.02

0.021

0.022

0.023

Width (mm)

k hyst

(W

kg−

1 T−

2 Hz−

1 )

Figure 5.2: Curve fitting of the hysteresisloss coefficients, khyst, for different widths ofM270-50A.

5 10 15 20 25 301.15

1.2

1.25

1.3

1.35

1.4x 10

−4

Width (mm)

k ec (

W k

g−1 T

−2 H

z−2 )

Figure 5.3: Curve fitting of the eddy currentloss coefficients, kec, for different widths ofM270-50A.

The fitting equation used is given by Equation 5.1.

kloss = α · e−b·w + c (5.1)

Where kloss is the loss coefficient and w is the width of the region. The constants a, b, care determined by the curve fitting.

The iron loss coefficients, i.e the khyst and kec, for 30, 15, 10 and 7.5 mm wide stripsare determined by surface fitting of the iron losses in each measured value of B and f ,based on the separation model, as presented in Section 4.2.

5.2 Model for Permeability at High Flux Densities

The magnetic measurements were conducted for flux densities up to 1.5 T. However,certain regions of the core materials reach higher levels of induction during the operationof the motor. That means that the BH curves that are applied in the FEM softwareshould contain values of the magnetizing field for flux densities higher than 1.5 T. Anapproach to extrapolate the values of B and H based on the last two measured values isintroduced in [57]. This method is called Exponential Law Extrapolation (ELE) and it isexpressed by Equations 5.2-5.3.

B = µ0(Ms[1− e−βH ] +H) (5.2)

dB

dH= µ0 + µ0Msβe

−βH (5.3)

Where µ0 is the vacuum permeability and equals to µ0 = 4π ·10−7(T· mA

). Ms is the satura-tion magnetization and β is the exponential coefficient. Solving the system of Equations5.2 and 5.3 for the two last measured values of B and H leads to the determination ofMs and β. Then, each value of B and H can be obtained by Equation 5.2.


5.3 Simulations of an Induction Motor

The proposed FEM model is implemented for a 7.5 kW induction motor. The character-istics of this motor are presented in Table 5.1.

Operating Voltage 400 VNominal Power 7.5 kW

Number of Poles 4Nominal Speed 1466 rpmOuter Diameter 220 mmInner Diameter 136 mmActive Length 155 mm

Stator Tooth Width 6.1 mmStator Yoke Width 21.2 mmRotor Tooth Width 7.6 mmRotor Yoke Width 19.4 mm

Number of Stator Slots 36Number of Rotor Slots 28

Lamination M270-50A

Table 5.1: Main characteristics of the reference induction motor.

The cross section of the reference motor is illustrated in Figure 5.1. The simulationsconcern nominal operation and the cases of a model with and without degradation ofmaterial. The data that are used for the case called ”no degradation” concern the mea-surement results of 30 mm wide mechanically cut M270-50A strips, as found in Chapter4. The results of the time stepping simulations for mechanically and laser cut core lam-inations are given in Table 5.2. The results for 75% and 50% of the nominal load arepresented in Tables 5.3, 5.4.

5.3. SIMULATIONS OF AN INDUCTION MOTOR 61

100% Load No Degradation Mech. Cut Laser Set 8 Laser Set 2Torque (Nm) 48.9 48.9 48.9 48.9

Slip (%) 2.271 2.287 2.289 2.305Nominal current (A) 14.52 14.77 14.72 15.01

cosφ 0.8161 0.8043 0.8078 0.7946Friction losses (W) 35.6 35.6 35.6 35.6

Iron losses (W) 77.2 100.3 102.9 116.0Stator copper losses (W) 333.1 344.4 339.4 353.1Rotor copper losses (W) 232.6 231.4 229.3 229.2

Total losses (W) 678.4 700.1 704.7 733.9Efficiency (%) 91.70 91.46 91.41 91.09

Slip deviation (%) 0 +0.7 +0.8 +1.5Nominal current deviation (%) 0 +1.7 +1.4 +3.4

cosφ deviation (%) 0 -1.4 -1.0 -2.6Iron loss deviation (%) 0 +30 +33 +50

Stator copper loss deviation (%) 0 +3.4 +1.9 +6.0Rotor copper loss deviation (%) 0 -0.5 -1.4 -1.5

Total loss deviation (%) 0 +3.2 +3.9 +8.2Efficiency deviation (%) 0 -0.26 -0.31 -0.67

Table 5.2: Comparison of performance and loss characteristics of the reference motorbefore and after the implementation of the degradation profile for mechanically and lasercut M270-50A laminations. Reference in the comparison is the model which does not havedegradation profile and the material properties correspond to mechanically cut standardEpstein strips (30 mm). Results concern simulations under 100% load.



Slip (%) 1.522 1.534 1.532 1.543Nominal current (A) 11.71 12 11.93 12.27


Iron losses (W) 72.5 83.6 94.5 109.4Stator copper losses (W) 207.2 217.8 212 224.4Rotor copper losses (W) 131.7 130.1 127.6 126.9

Total losses (W) 447 467.1 469.6 496.3Efficiency (%) 92.64 92.33 92.29 91.89


cosφ deviation (%) 0 -2.2 -1.5 -3.9Iron loss deviation (%) 0 +15.3 +30.3 +50.6




5.3. SIMULATIONS OF AN INDUCTION MOTOR 63


Slip (%) 0.947 0.954 0.952 0.957Nominal current (A) 9.38 9.73 9.59 9.96


Iron losses (W) 70.1 80.7 91.5 106Stator copper losses (W) 130.1 140.1 133.2 143.6Rotor copper losses (W) 67.4 65.6 63 62.1

Total losses (W) 303.2 321.9 323.3 347.4Efficiency (%) 92.52 92.09 92.06 91.52


cosφ deviation (%) 0 -3.2 -1.7 -4.9Iron loss deviation (%) 0 +15.1 +30.5 +51.2




Laser cutting with Set 2 gives the worst performance. The efficiency drops by 0.67%while the iron losses increase by 50%, compared to simulations without any degradation

50 60 70 80 90 10091

91.2

91.4

91.6

91.8

92

92.2

92.4

92.6

92.8

Load (%)

Effi

cien

cy (

%)

No degr.GuillotineLaser Set 8Laser Set 2

Figure 5.4: Efficiency for different loads.

50 60 70 80 90 1001.02

1.04

1.06

1.08

1.1

1.12

1.14

1.16

1.18

Load (%)

Tot

al lo

sses

(pu

)

GuillotineLaser Set 8Laser Set 2

Figure 5.5: Total losses for different loads,with the simulation without degradation pro-file as reference.


profile. Correspondingly, mechanically and laser cut laminations with Set 8 show iron lossincrease of 30% and 33% respectively. The degradation of permeability causes increasednominal current. This leads to lower power factor and higher stator copper losses. Sim-ulations for mechanically cut laminations indicate a 3.4% increase of the stator copperlosses and 1.4% decrease of the power factor, compared to simulations without degrada-tion profile. The respective increase for laser cut laminations with Set 8 is 1.9%, while thepower factor decreases by 1%. For 50% and 75% load, the incorporation of degradationprofile leads to larger deviations, as shown in Figures 5.4 and 5.5.

5.4 Summary

• The advantage of the presented model is the consideration of different degrada-tion characteristics according to the geometry of the lamination. Thus, it can beimplemented in any motor design.

• Permeabilities which correspond to higher flux densities than the measured ones aredetermined. Exponential Law Extrapolation is used for this purpose.

• The implemented BH curves and iron loss coefficients are based on measurementresults. For non measured lamination widths, values through curve fitting are ap-plied. The four measured lamination widths (7.5, 10, 15 and 30 mm) allow curvefitting with good accuracy. These widths are representative of the width of yokeand teeth studied motors.

• The suggested degradation profile is incorporated in the FEM model of an inductionmotor. The new model results in the estimation of lower efficiency and power factor.The iron losses are found higher.

• For laser cut laminations with Set 2, the iron loss increase is 50%. For mechanicallycut laminations this increase is 30%.

• Due to degraded permeability, there is a 1.7 to 3.4% increased nominal current.That corresponds to a 1.4 to 2.6% drop, respectively, in the power factor. Theincreased nominal current causes 1.9 to 6% additional stator copper losses.

• Operation under lower load than the nominal leads to larger deviations due to theimplementation of degradation profile.

Chapter 6

Conclusions and Future Work

This chapter summarizes the most important conclusions of this thesis. Future investiga-tions in this field are also proposed.

6.1 Conclusions

In this thesis, the degradation of the magnetic materials due to the manufacturing processis investigated experimentally. In Chapter 3, a review of the most influential productionsteps, regarding the deterioration of the magnetic properties, is presented. Based on thisstudy, cutting the core laminations is recognized as the process that degrades the mostthe magnetic materials. This is the reason why the cutting effect is further investigatedthough experiments by means of Epstein measurements.

In Section 4.2, the effect of mechanical cutting is investigated. It is found that as thewidth of the cut lamination decreases, the relative permeability of the magnetic materialdrops and the specific iron losses increase. An outcome of this investigation is also thatthe cutting effect is reduced for thinner laminations. The maximum permeability drop ofa 3.2% Si lamination of 0.5 mm thickness is found to be 61%. The maximum iron lossdensity increase of the same material is 41%. A similar lamination with 0.2 mm thicknesshas a corresponding 41% permeability drop and 20% increased iron loss density. Thecutting effect also depends on the Si-content and the induction. Laminations with highSi-content are found to have higher degradation at low induction.

A next step in the study of the cutting effect is the investigation of the magneticproperties of the steel laminations after laser cutting. In Section 4.3, it is found thatfiber laser has less impact on the magnetic properties than CO2 laser. Another result ofthis study is also that the parameters power, speed, frequency and gas pressure of thelaser play a major role in the magnetic properties of the steel laminations. Particularly,for M400-50A laminations that are cut with high-speed/high-power settings (Set 8 ), thepermeability is 99% higher and the specific iron losses 27% lower compared to identicallaminations that are cut with low-speed/high-power settings (Set 2 ). Increased cuttingspeed could lead to better productivity. The cutting effect for laser cut laminationsis investigated for different strip widths. It is found that laser cutting with the high-speed/high-power setting (Set 8 ) causes up to 65% higher iron loss density and 65%lower permeability, compared to the standard 30 mm Epstein strips (also laser cut withSet 8 ).

65

66 CHAPTER 6. CONCLUSIONS AND FUTURE WORK

The magnetic measurements of identical laminations that are cut by means of guillo-tine and laser has indicated that there is a difference in the properties of the magneticmaterial, according to the different cutting technique. This result is valuable for motordesigners, since the laminations of mass production motors are cut mechanically by meansof punching, while the prototype motors are cut with laser. In Section 4.4, it is foundthat at 0.5 T, the permeability of a mechanically cut M400-50A lamination is 39% higherand the specific iron losses are 17% lower than a respective high-power/high-speed lasercut lamination.

In Section 4.5, experimental results on the welding effect are presented. It is foundthat the magnetic material is deteriorated as the number of welding seams increases.Increased iron loss density up to 50% and a corresponding permeability drop of 62% aremeasured.

The findings regarding the cutting effect on iron losses are valuable for motor designersonly if they can be incorporated during the design process. This is the reason why a FEMmodel which includes a degradation profile is developed in Chapter 5. In this model, themagnetization and loss characteristics of the motor change according to the geometry, i.ethe width of the lamination. It is found that the degraded magnetic material causes notonly additional iron losses but also copper losses due to the deteriorated magnetizationproperties. The advantage of this approach is that it considers the motor geometry andcan be implemented in any design. FEM simulations of a reference induction motor haveindicated a decrease of 2.6% in the power factor and an increase of up to 50 % in the ironlosses.

6.2 Future Work

In this thesis, aspects of the cutting and welding effect are investigated. Interesting topicsthat could be further investigated are:

• Comparison among welding, cleating and gluing: The major methods oflamination stacking are welding, cleating and gluing. A future work could includethe comparison of these techniques in terms of additional iron losses and mechanicaldurability.

• Stress relief annealing: A suggested investigation is annealing the strips thatare cut by different techniques and have different widths. This study would high-light the cases, where annealing can lead to significant improvement of the magneticproperties of the steel laminations. Furthermore, annealing with different combi-nations of temperatures and time would be an investigation that could lead to theoptimization of the annealing process.

• Investigation on the wear of the mechanical cutting tool: One of the disad-vantages of the guillotine or the punching tool is that it requires regrinding after aperiod of time. An interesting study would be to investigate the degradation of themagnetic material due to mechanical cutting as a function of the time after the lastregrinding.

• Magnetic measurements on punched strips: In this study, the mechanicalcutting is done by guillotine. However, it would be interesting to make similar

6.2. FUTURE WORK 67

measurements on punched laminations. This way, the cutting effect due to punchingcan be determined and compared with the corresponding due to guillotine cutting.

• Validation of FEM results with measurements on real motors: A com-parison of the FEM with measurement results would validate the proposed FEMmodel.

68 CHAPTER 6. CONCLUSIONS AND FUTURE WORK

Appendix A

Guillotine Cutting

A.1 M400-50A

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60

5

10

15

20

25

B (T)

PF

e (W

/kg)

30 mm/ 50 Hz15 mm/ 50 Hz10 mm/ 50 Hz7.5 mm/ 50 Hz30 mm/ 100 Hz15 mm/ 100 Hz10 mm/ 100 Hz7.5 mm/ 100 Hz30 mm/ 200 Hz15 mm/ 200 Hz10 mm/ 200 Hz7.5 mm/ 200 Hz

Figure A.1: Iron loss curves of M400-50A for different strip widths and frequencies.

69

70 APPENDIX A. GUILLOTINE CUTTING

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6−70

−60

−50

−40

−30

−20

−10

0Deviation from 30 mm

B (T)

Dev

iatio

n of

µr (

%)

15 mm/ 50 Hz10 mm/ 50 Hz7.5 mm/ 50 Hz15 mm/ 100 Hz10 mm/ 100 Hz7.5 mm/ 100 Hz15 mm/ 200 Hz10 mm/ 200 Hz7.5 mm/ 200 Hz

Figure A.2: Deviation of the relative permeability of 15 , 10 and 7.5 mm wide M400-50AEpstein strips from a corresponding 30 mm wide strip.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.65

10

15

20

25

30

35


B (T)

Dev

iatio

n of

PF

e (%

)


Figure A.3: Deviation of the iron losses of 15 , 10 and 7.5 mm wide M400-50A Epsteinstrips. Reference is the standard 30 mm wide strip.

A.1. M400-50A 71

0.20.4

0.60.8

11.2

50

100

150

2000

5

10

15

20

25

B (T)f (Hz)

p Fe (

W/k

g)

Figure A.4: Surface fitting to separation model with two terms of the iron loss density ofM400-50A for 30 mm strips.

The error of the fitting model is expressed by the residuals, which are illustrated in FigureA.5.

Figure A.5: Residuals of surface fitting to separation model of the iron loss density ofM400-50A for 30 mm strips.

Based on Figure A.5, the maximum values of the relative error are given in Table A.1.


Max. relative error B of max. relative error50 Hz 18% 0.5 T100 Hz 14% 0.5 T200 Hz 2% 1.4 T

Table A.1: Relative error of the surface fitting in the case of mechanically cut 30 mmwide M400-50A.

A.2 M270-50A

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60

2

4

6

8

10

12

14

16

18

20

B (T)

PF

e (W

/kg)


Figure A.6: Iron loss curves of M270-50A for different strip widths and frequencies.

A.2. M270-50A 73

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6−70

−60

−50

−40

−30

−20

−10

0


B (T)

Dev

iatio

n of

µr (

%)


Figure A.7: Deviation of the relative permeability of 15 , 10 and 7.5 mm wide M270-50AEpstein strips from a corresponding 30 mm wide strip.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60

10

20

30

40

50

60


B (T)

Dev

iatio

n of

PF

e (%

)

15 mm10 mm7.5 mm15 mm10 mm7.5 mm15 mm10 mm7.5 mm

Figure A.8: Deviation of the iron losses of 15 , 10 and 7.5 mm wide M270-50A Epsteinstrips. Reference is the standard 30 mm wide strip.


0.20.4

0.60.8

11.2

50

100

150

2000

2

4

6

8

10

12

14

16

B (T)f (Hz)

p Fe (

W/k

g)

Figure A.9: Surface fitting to separation model with two terms of the iron loss density ofM270-50A for 30 mm strips.

The residuals of the fitting are illustrated in Figure A.10.

Figure A.10: Residuals of surface fitting to separation model of the iron loss density ofM270-50A for 30 mm strips.

Based on Figure A.10, the maximum values of the relative error are given in TableA.2.

A.3. NO20 75

Max. relative error B of max. relative error50 Hz 14% 0.6 T

100 Hz 12% 0.6 T200 Hz 1% 0.5 T

Table A.2: Relative error of the surface fitting in the case of mechanically cut 30 mmwide M270-50A.

A.3 NO20

0 1000 2000 3000 4000 5000 60000

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

H (A/m)

B (

T)


Figure A.11: BH curve of NO20 for different strip widths and frequencies.


0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.60

2

4

6

8

10

12

14

B (T)

PF

e (W

/kg)


Figure A.12: Iron loss curves of NO20 for different strip widths and frequencies.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6−5

0

5

10

15

20

25

30


B (T)

Dev

iatio

n of

PF

e (%

)


Figure A.13: Deviation of the iron losses of 15 , 10 and 7.5 mm wide NO20 Epstein strips.Reference is the standard 30 mm wide strip.

A.3. NO20 77

0.20.4

0.60.8

11.2

50

100

150

2000

1

2

3

4

5

6

7

8

9

10

B (T)f (Hz)

p Fe (

W/k

g)

Figure A.14: Surface fitting to separation model with two terms of the iron loss densityof NO20 for 30 mm strips.

The residuals of the surface fitting are illustrated in Figure A.15.

Figure A.15: Residuals of surface fitting to separation model of the iron loss density ofNO20 for 30 mm strips.

Based on Figure A.15, the maximum values of the relative error are given in TableA.3 .


Max. relative error B of max. relative error50 Hz 6% 1.1 T

100 Hz 5% 1.1 T200 Hz 4% 1.4 T

Table A.3: Relative error of the surface fitting in the case of mechanically cut 30 mmwide NO20.

Appendix B

Laser Cutting

Name of Set-ting

Power (pu) Speed(pu)

Freq.(pu)

Gas pres-sure (pu)

FL-lp/lsNO20

0.06 0.1 0.4 0.5

FL-hp/hsNO20

1 1 0.4 0.5

CL NO20 0.16 0.14 1 0.65

Table B.1: Laser settings of NO20

In Table B.1, FL-lp/ls stands for low-power/low-speed settings of fiber laser, while FL-hp/hs stands for high-power/high-speed settings of the same laser. In this section, themagnetic characteristics of NO20, after laser cutting with different settings and laser typeare investigated. Figures B.1-B.2 depict the relative permeability and the iron loss densityrespectively of NO20 strips.

0 0.5 1 1.50

500

1000

1500

2000

2500

3000

Rel

ativ

e pe

rmea

bilit

y

B (T)

FL high power/high speedFL low power/low speedCO2 laser

Figure B.1: Relative permeability of 10mmNO20 at 50 Hz.

0 0.5 1 1.50

0.5

1

1.5

2

2.5

3

3.5

Iron

Los

ses

(W/k

g)

B (T)

FL high power/high speedFL low power/low speedCO2

Figure B.2: Iron losses of 10mm NO20 at 50Hz.

It can be seen from Figure B.1 that the relative permeability of FL-hp/hs is up to

79

80 APPENDIX B. LASER CUTTING

23% higher than the respective one of the FL-lp/ls technique. As far as the iron lossdensity is concerned, cutting with FL-hp/hs settings leads to lower iron losses than therespective losses of FL-lp/ls.

It is also worth mentioning that CO2 laser deteriorates the most the magnetic material.Indicatively, the permeability is up to 74% higher, while the iron loss density is 30% lowercompared the FL-hp/hs cutting.

Tables B.2 summarizes the variation of the magnetic characteristics of NO20, whenfiber laser is concerned.

0.1 T 0.5 T 1 T 1.5 TDeviation of µr 7% 23% 20% 1%Deviation of pFe -8% -9% -6% -3%

Table B.2: Deviation of relative permeability and iron loss density between FL-lp/ls andFL-hp/hs laser cut NO20 at 50 Hz. Results from FL-lp/ls settings are used as reference.

Table B.3 summarizes the difference of the magnetic characteristics of NO20 aftercutting with FL-hp/hs and CO2 laser is performed.

0.1 T 0.5 T 1 T 1.5 TDeviation of µr 41% 74% 71% 9%Deviation of pFe -30% -29% -22% -16%

Table B.3: Deviation of relative permeability and iron loss density between CO2 andFL-hp/hs laser cut NO20 at 50 Hz. Results from CO2 laser are used as reference.

Bibliography

[1] E. Commission, EU Action on Climate,http://ec.europa.eu/clima/policies/. 2015.

[2] F. Fleiter, W. Eichhammer, and J. Schleich, Energy efficiency in electric motor sys-tems: Technical potentials and policy approaches for developing countries. UnitedNations- Industrial Development Organization, 2011.

[3] ABB, Technical Note: IEC 60034-30-1 standard on efficiency classes for low voltageAC motors. 2014.

[4] G. Bertotti, “Chapter 7 - magnetic domains and domain walls,” in Hysteresis inMagnetism (G. Bertotti, ed.), Electromagnetism, pp. 189 – 221, San Diego: AcademicPress, 1998.

[5] P. Beckley, Electrical Steels for Rotating Machines. IET Power and Energy Series 37,2002.

[6] A. Schoppa, J. Schneider, and C.-D. Wuppermann, “Influence of the manufactur-ing process on the magnetic properties of non-oriented electrical steels,” Journal ofMagnetism and Magnetic Materials, vol. 215–216, no. 0, pp. 74 – 78, 2000.

[7] N. A. Spaldin, Magnetic Materials: Fundamentals and Applications. CambridgeUniversity Press, 2011.

[8] A. Broddefalk and M. Lindenmo, “Dependence of the power losses of a non-oriented3and Magnetic Materials, vol. 304, no. 2, pp. e586 – e588, 2006. Proceedings of the17th International Symposium on Soft Magnetic Materials Proceedings of the 17thInternational Symposium on Soft Magnetic Materials.

[9] A. Krings, Iron Losses in Electrical Machines- Influence of Material Properties, Man-ufacturing Processes and Inverter Operation. PhD thesis, KTH Royal Institute ofTechnology, 2014.

[10] A. Goldman, Modern Ferrite Technology. Springer, 2nd ed., 2010.

[11] S. Chikazumi, Physics of Ferromagnetism (The International Series of Monographson Physics). The International Series of Monographs on Physics, Oxford UniversityPress, USA, 2 ed., 2009.

[12] G. Bertotti, “Physical interpretation of eddy current losses in ferromagnetic materi-als. ii. analysis of experimental results,” Journal of Applied Physics, vol. 57, no. 6,1985.

81

82 BIBLIOGRAPHY

[13] C. Serpico, C. Visone, I. D. Mayergoyz, V. Basso, and G. Miano, “Eddy currentlosses in ferromagnetic laminations,” Journal of Applied Physics, vol. 87, no. 9, 2000.

[14] C. Steinmetz, “On the law of hysteresis,” Proceedings of the IEEE, vol. 72, pp. 197–221, Feb 1984.

[15] M. Reinlein, T. Hubert, A. Hoffmann, and A. Kremser, “Optimization of analyticaliron loss approaches for electrical machines,” in Electric Drives Production Confer-ence (EDPC), 2013 3rd International, pp. 1–7, Oct 2013.

[16] J. Reinert, A. Brockmeyer, and R. De Doncker, “Calculation of losses in ferro- andferrimagnetic materials based on the modified steinmetz equation,” in Industry Ap-plications Conference, 1999. Thirty-Fourth IAS Annual Meeting. Conference Recordof the 1999 IEEE, vol. 3, pp. 2087–2092 vol.3, 1999.

[17] “Overview and comparison of iron loss models for electrical machines,” Journal ofElectrical Engineering, vol. 10(3), pp. 162 – 169, 2010.

[18] H. Jordan, “Die ferromagnetischen konstanten fur schwache wechselfelder.,” Elektr.Nach. Techn., vol. 1, p. 8, 1924.

[19] R. H. Pry and C. P. Bean, “Calculation of the energy loss in magnetic sheet materialsusing a domain model,” Journal of Applied Physics, vol. 29, no. 3, 1958.

[20] S. Jacobs, F. Henrotte, M. Herranz Gracia, K. Hameyer, P. Goes, and D. Hectors,“Magnetic material optimization for hybrid vehicle PMSM drives,” World ElectricVehicle Journal (www.evs24.org/wevajournal), vol. 3, 2009.

[21] I. Mayergoyz, “CHAPTER 1 - the classical preisach model of hysteresis,” in Math-ematical Models of Hysteresis and Their Applications (I. Mayergoyz, ed.), Electro-magnetism, pp. 1 – 63, New York: Elsevier Science, 2003.

[22] IEC, 60404-6: Magnetic Materials- Part 6: Methods of measurement of magneticallysoft metallic and powder materials at frequencies in the range 20 Hz to 200 kHz bythe use of ring specimens. 2003.

[23] A. Krings, M. Cossale, J. Soulard, A. Boglietti, and A. Cavagnino, “Manufacturinginfluence on the magnetic properties and iron losses in cobalt-iron stator cores forelectrical machines,” in Energy Conversion Congress and Exposition (ECCE), 2014IEEE, pp. 5595–5601, Sept 2014.

[24] A. Clerc and A. Muetze, “Measurement of stator core magnetic degradation duringthe manufacturing process,” Industry Applications, IEEE Transactions on, vol. 48,pp. 1344–1352, July 2012.

[25] F. Fiorillo, Characterization and Measurement of Magnetic Materials (Electromag-netism). Academic Press, 2005.

[26] IEC, 404-2: Magnetic materias-Part 2: Methods of measurement of the magneticproperties of electrical steel sheet and strip by means of Epstein frame. 1996.

BIBLIOGRAPHY 83

[27] A. Schoppa, J. Schneider, and J.-O. Roth, “Influence of the cutting process on themagnetic properties of non-oriented electrical steels,” Journal of Magnetism and Mag-netic Materials, vol. 215–216, no. 0, pp. 100 – 102, 2000.

[28] M. Bali, H. De Gersem, and A. Muetze, “Finite-element modeling of magnetic ma-terial degradation due to punching,” Magnetics, IEEE Transactions on, vol. 50,pp. 745–748, Feb 2014.

[29] A. Clerc and A. Muetze, “Measurement of stator core magnetic characteristics,” inElectric Machines Drives Conference (IEMDC), 2011 IEEE International, pp. 1433–1438, May 2011.

[30] A. Smith and K. Edey, “Influence of manufacturing processes on iron losses,” inElectrical Machines and Drives, 1995. Seventh International Conference on (Conf.Publ. No. 412), pp. 77–81, Sep 1995.

[31] M. Emura, F. Landgraf, W. Ross, and J. Barreta, “The influence of cutting tech-nique on the magnetic properties of electrical steels,” Journal of Magnetism andMagnetic Materials, vol. 254–255, no. 0, pp. 358 – 360, 2003. Proceedings of the 15thInternational Conference on Soft Magnetic Materials (SMM15).

[32] A. Moses, N. Derebasi, G. Loisos, and A. Schoppa, “Aspects of the cut-edge ef-fect stress on the power loss and flux density distribution in electrical steel sheets,”Journal of Magnetism and Magnetic Materials, vol. 215–216, no. 0, pp. 690 – 692,2000.

[33] R. Rygal, A. Moses, N. Derebasi, J. Schneider, and A. Schoppa, “Influence of cuttingstress on magnetic field and flux density distribution in non-oriented electrical steels,”Journal of Magnetism and Magnetic Materials, vol. 215–216, no. 0, pp. 687 – 689,2000.

[34] A. Boglietti, “A first approach for the iron losses building factor determination,” inIndustry Applications Conference, 1999. Thirty-Fourth IAS Annual Meeting. Con-ference Record of the 1999 IEEE, vol. 1, pp. 489–493 vol.1, 1999.

[35] A. Boglietti, A. Cavagnino, L. Ferraris, and M. Lazzari, “The annealing influenceonto the magnetic and energetic properties in soft magnetic material after punch-ing process,” in Electric Machines and Drives Conference, 2003. IEMDC’03. IEEEInternational, vol. 1, pp. 503–508 vol.1, June 2003.

[36] A. Moses, N. Derebasi, G. Loisos, and A. Schoppa, “Aspects of the cut-edge ef-fect stress on the power loss and flux density distribution in electrical steel sheets,”Journal of Magnetism and Magnetic Materials, vol. 215–216, no. 0, pp. 690 – 692,2000.

[37] K. H. Schmidt, “Der einfluss des stanzens auf die magnetischen eigenschaften vonelektroblech mit 1Materials, vol. 2, no. 1–3, pp. 136 – 150, 1976.

[38] L. Vandenbossche, S. Jacobs, F. Henrotte, and K. Hameyer, “Impact of cut edgesin magnetization curves and iron losses in e-machines for automotive traction,” in

84 BIBLIOGRAPHY

Proceedings of 25th World Battery, Hybrid and Fuel Cell Electric Vehicle Symposium& Exhibition, EVS, (Schenzhen, China), November 2010.

[39] P. Baudouin, A. Belhadj, F. Breaban, A. Deffontaine, and Y. Houbaert, “Effects oflaser and mechanical cutting modes on the magnetic properties of low and mediumsi content nonoriented electrical steels,” Magnetics, IEEE Transactions on, vol. 38,pp. 3213–3215, Sep 2002.

[40] R. Siebert, J. Schneider, and E. Beyer, “Laser cutting and mechanical cutting of elec-trical steels and its effect on the magnetic properties,” Magnetics, IEEE Transactionson, vol. 50, pp. 1–4, April 2014.

[41] W. Arshad, T. Ryckebusch, F. Magnussen, H. Lendenmann, B. Eriksson, J. Soulard,and B. Malmros, “Incorporating lamination processing and component manufactur-ing in electrical machine design tools,” in Industry Applications Conference, 2007.42nd IAS Annual Meeting. Conference Record of the 2007 IEEE, pp. 94–102, Sept2007.

[42] T. Ryckebusch, “Quantifying spread of iron losses in virgin and processed lamina-tions for electrical motors and machines,” Master’s thesis, KTH, Royal Institute ofTechnology, 2005.

[43] W. M. Arshad, T. Ryckebush, A. Broddefalk, F. Magnussen, H. Lendenmann, andM. Lindenmo, “Characterization of electrical steel grades for direct application toelectrical machine design tools,” Journal of Magnetism and Magnetic Materials,vol. 320, no. 20, pp. 2538 – 2541, 2008.

[44] T. Nakata, M. Nakano, and K. Kawahara, “Effects of stress due to cutting on mag-netic characteristics of silicon steel,” Magnetics in Japan, IEEE Translation Journalon, vol. 7, pp. 453–457, June 1992.

[45] O. Hubert and E. Hug, “Influence of plastic strain on magnetic behaviour of non-oriented fe–3si and application to manufacturing test by punching,” Materials Scienceand Technology, vol. 11, no. 5, pp. 482–487, 1995.

[46] L. Vandenbossche, S. Jacobs, X. Jannot, M. McClelland, J. Saint-Michel, and E. At-trazic, “Iron loss modelling which includes the impact of punching, applied to high-efficiency induction machines,” in Electric Drives Production Conference (EDPC),2013 3rd International, pp. 1–10, Oct 2013.

[47] A. Kedous-Lebouc, B. Cornut, J. Perrier, P. Manfe, and T. Chevalier, “Punchinginfluence on magnetic properties of the stator teeth of an induction motor,” Journalof Magnetism and Magnetic Materials, vol. 254–255, no. 0, pp. 124 – 126, 2003.

[48] J. Powell and A. F. H. Kaplan, “Laser cutting technology — a commercial perspec-tive,” Laser Technik Journal, vol. 9, no. 2, pp. 39–41, 2012.

[49] A. Belhadj, P. Baudouin, F. Breaban, A. Deffontaine, M. Dewulf, and Y. Houbaert,“Effect of laser cutting on microstructure and on magnetic properties of grain non-oriented electrical steels,” Journal of Magnetism and Magnetic Materials, vol. 256,no. 1–3, pp. 20 – 31, 2003.

BIBLIOGRAPHY 85

[50] A. Schoppa, H. Louis, F. Pude, and C. von Rad, “Influence of abrasive waterjetcutting on the magnetic properties of non-oriented electrical steels,” Journal of Mag-netism and Magnetic Materials, vol. 254–255, no. 0, pp. 370 – 372, 2003.

[51] A. Schoppa, J. Schneider, C.-D. Wuppermann, and T. Bakon, “Influence of weldingand sticking of laminations on the magnetic properties of non-oriented electricalsteels,” Journal of Magnetism and Magnetic Materials, vol. 254–255, no. 0, pp. 367– 369, 2003.

[52] A. Krings, S. Nategh, O. Wallmark, and J. Soulard, “Influence of the welding processon the magnetic properties of a slot-less permanent magnet synchronous machinestator core,” in Electrical Machines (ICEM), 2012 XXth International Conferenceon, pp. 1333–1338, Sept 2012.

[53] G. B. K. Hamid A. Toliyat, Handbook of Electric Motors. CRC Press, 2 ed., 2004.

[54] H. Domeki, Y. Ishihara, C. Kaido, Y. Kawase, S. Kitamura, T. Shimomura, N. Taka-hashi, T. Yamada, and K. Yamazaki, “Investigation of benchmark model for esti-mating iron loss in rotating machine,” Magnetics, IEEE Transactions on, vol. 40,pp. 794–797, March 2004.

[55] M. Oka, T. Ogasawara, N. Kawano, and M. Enokizono, “Estimation of suppressediron loss by stress-relief annealing in an actual induction motor stator core by usingthe excitation inner core method,” Magnetics, IEEE Transactions on, vol. 50, pp. 1–4, Nov 2014.

[56] R. Siebert, A. Lutke, M.and Wetzig, H. Thonig, and E. Beyer, “Experimental find-ings on high power laser cutting of non-oriented electrical steel,” in InternationalConference Magnetism and Metallurgy at Ghent, Belgium, 2012.

[57] A. Umenei, Y. Melikhov, and D. Jiles, “Models for extrapolation of magnetizationdata on magnetic cores to high fields,” Magnetics, IEEE Transactions on, vol. 47,pp. 4707–4711, Dec 2011.

Documents

Manufacturing Effects on Iron Losses in Electrical Machines847378/... · 2015-08-20 · Manufacturing E ects on Iron Losses in Electrical Machines Konstantinos Bourchas Master of