performance analysis of unskewed asymmetrical rotor for lv...

Degree project in

Performance Analysis of UnskewedAsymmetrical Rotor for LV Induction

Motors

USMAN SHAUKAT

Stockholm, Sweden 2012

XR-EE-E2C 2012:018

Electrical EngineeringMaster of Science

Performance Analysis of Unskewed AsymmetricalRotor for LV Induction Motors

USMAN SHAUKAT

Master of Science Thesis in Electrical Energy Conversion (E2C)School of Electrical EngineeringRoyal Institute of Technology

Stockholm, Sweden, October 2012

Supervisor: Alexander Stening (KTH), Rathna Chitroju (ABB)Examiner: Chandur Sadarangani, Professor

XR-EE-E2C 2012:018

i

AbstractThis master thesis presents a comparative analysis of the starting per-formance and losses at rated operation for a 15 kW , 4-pole industrial in-duction motor, mounted with standard skewed, unskewed and unskewedasymmetrical die-cast aluminium rotors through measurements and sim-ulations.

It is a well-known fact that rotor skewing suppresses the synchronoustorques at low speeds and also reduces the audible noise of the machine.However, the casting process results in a low resistive path between therotor bars and the iron laminations, for skewed rotors, this promotesthe flow of inter-bar currents. These currents, flowing between the rotorbars, increase the harmonic torques during a start and create additionallosses at rated operation. For standard unskewed rotors, these losses areideally zero, but these rotors may produce high audible noise. Studieshave shown that rotors with asymmetrical rotor slot pitch can reducethe audible noise level in unskewed machines. By removing the skew, theinter-bar current losses are suppressed to a negligible level; ultimatelyincreased machine efficiency is obtained. In this work the electrical per-formance is verified through measurements on the built prototypes.

Direct-on-line starts and rated performance for motors with differentrotor slot arrangements is simulated using 2D FEM tool FCSmek. Thethree prototypes are tested in the laboratory according to IEC 60034-2-1 standard and the simulation results are in good agreement withthe measured results. An additional test for the measurement of highfrequency delta connected stator winding currents for each prototypemachine is also performed, in order to study the losses induced in thestator winding.

Results have shown that by introducing the proposed asymmetry inthe rotor slots, the synchronous torques at low speeds are suppressedeffectively, thus, improving the starting performance of the asymmet-rical rotor compared to the standard unskewed rotor. Additionally, ahigher pull-out torque is obtained for the unskewed rotor motor com-pared to the standard skewed rotor motor. However, the losses weremore or less re-distributed in the unskewed rotor motor, resulting insimilar efficiency as the standard skewed rotor motor. One importantobservation is that; to capture the inter-bar current losses which are es-timated to be 5.5% of the total losses, requires more accurate methodsof measurements than the existing. And sufficient repeatability must beachieved; alternatively one should rely on statistical data obtained frommeasurements on several number of motors.

Keywords: Induction motor, skewed rotor, inter-bar currents,starting performance, asymmetrical rotor slots, 2D FEM, syn-chronous torques.

ReferatUtvärdering av Asymmetrisk Rotor Utan

Spårsneddning för AsynkronmotorerDetta examensarbete presenterar en jämförande studie av startegenska-per och förluster vid märkdrift för en 15 kW , 4-polig asynkronmotor,utrustad med olika typer av högtrycksgjutna aluminiumrotorer. Mät-ningar och simuleringar utförs då motorn utrustas med en symmetriskrotor, med och utan spårsneddning, samt även för en asymmetrisk rotorutan spårsneddning.

Sneddning av rotorspåren i asynkronmotorer är vanligt förekommande,det används för att reducera asynkrona och synkrona moment vid lågavarvtal samt för att minska det elektromagnetiska ljudet från motorn.Men avsaknaden av spårisolation i högtrycksgjutna rotorer resulterar iatt strömmar kan ledas fritt mellan spåren, dessa tvärströmmar ökarkraftigt med spårsneddningen. Tvärströmmarna ökar de asynkrona mo-menten vid en start, och kan även öka tillsatsförlusterna vid märkdrift.Genom att använda rotorer utan spårsneddning kan dessa strömmar re-duceras till en försumbar nivå, och optimalt uppnås därmed en högrevekningsgrad.

Genom simuleringsprogrammet FCSmek, vilket använder Finita Ele-ment Metoden, beräknas startegenskaper och förluster vid märkdrift dåmotorn utrustas med de olika rotorerna. Motsvarande mätningar utförspå prototypmotorer enligt mätstandarden IEC 600-34-2-1, resultaten vi-sar god korrelation med simuleringarna. Ett extra prov utförs även föratt studera hur de högfrekventa strömmarna i statorlindningen påver-kas av rotordesignen. Resultaten visar att ett högre startmoment ochett högre kipmoment uppnås då rotorn ej har spårsneddning, men tillen kostnad av ökade synkrona moment vid låga varvtal. Detta problemreducerades då en spårasymmetri infördes. Förbättrade startegenskaperuppnås därför med den asymmetriska rotorn jämfört med den symmet-riska rotorn utan spårsneddning. Mätningarna visade att de totala för-lusterna var mer eller mindre oförändrade för de tre rotorerna, vilketäven gällde tillsatsförlusterna. Detta indikerar att förlusterna omförde-lades i motorn, snarare än att de reducerades, då spårsneddning ej an-vändes. Det konstaterades även att noggrannare mätmetoder med högrerepeterbarhet krävs för att med precision kunna jämföra förändringenav tvärströmsförlusterna, som för de studerade motorerna uppskattadestill 5.5% av de totala förlusterna.

Sökord: Asynkronmotor, induktionsmotor, tvärströmmar, starte-genskaper, asymmetrisk rotor, 2D FEM, synkrona moment.

Acknowledgment

This master degree project was conducted in cooperation between ABB LV Motors(Västerås) and KTH (Stockholm). First of all, I would like to thank Allah Almightyfor helping me all the way to reach this point. I would like to thank the ElectricalEnergy Conversion (E2C) department in KTH and ABB for providing me the op-portunity to work at one of the world’s leading engineering company.

I would like to thank Alexander Stening who was my supervisor in KTH. He hadgiven me patient explanation about the theory, guided me very well, taught me howto work in the right direction and encouraged me all the way.

Also, I would like to thank my supervisor Rathna Chitroju from ABB LV Mo-tors, who gave me great help during my thesis work and did the detailed proofreadof the report. I am especially thankful to him for introducing me to LaTeX. Hisprofound nature gave me great impression.

I also want to thank my examiner Professor Chandur Sadarangani for his kind-ness and valuable guidance for my studies at KTH.

Last but not least, I would like to thank my parents, my brother, sisters and all thefamily members for their moral and financial support throughout my study work;my dear friends Arslan Zafar, Amit Kumar Jha, Ibrahim Bilal and Ali Shaheen forbeing an inspirational source and a helping hand through this journey.

Usman ShaukatVästerås, October 2012.

iii

List of Symbols & Abbreviations

I Average line current [A]I0 No-load average line current [A]Is Stator current [A]Ir Rotor current [A]Iab Winding current in one phase [A]L Axial length of machine [mm]Nr Rated mechanical speed [rpm]PN Rated power [W ]PT , Ploss Total power losses [W ]Pin Input power [W ]P0 No-load input power [W ]Pout Output power [W ]PK Constant losses [W ]PL Load losses [W ]PLL Additional load losses or Stray load losses [W ]Pfe Iron losses [W ]Pfw Friction & windage losses [W ]Pcu1, Ps Stator copper losses [W ]Pcu2, Pr Rotor copper losses [W ]Qs Number of stator slots [...]Qr Number of rotor slots [...]R Stator line to line resistance [Ω]RN Stator winding resistance from rated load test [Ω]R||0 Line to line resistance from no-load test [Ω]Rs Stator winding resistance [Ω]Rr Rotor winding resistance [Ω]R′

r Rotor winding resistance refer to stator side [Ω]Rtot Total locked rotor resistance [Ω]T Output torque [Nm]TN Rated torque [Nm]UN Rated voltage [V ]U0 No-load voltage [V ]

v

vi

Xtot Total locked rotor reactance [Ω]Xsl Stator leakage reactance [Ω]X ′

rl Rotor leakage reactance referred to stator side [Ω]

f(1) Fundamental frequency [Hz]fN Rated frequency [Hz]fs Sampling frequency [Hz]fc Cut-off frequency [Hz]ω Synchronous speed [rad/sec]ωr Rotor mechanical speed [rad/sec]ρ Inter-bar resistivity [Ωm]ϕ Induced flux [Wb]θ0 Initial slot pitch angle [rad]b0 Initial slot width [mm]θc Inlet cooling temperature during tests [0C]θw Reference winding temperature [0C]s(1) Fundamental slip [...]Ki,δi ith modulation coefficients [...]γ Correlation coefficient [...]kθ Winding temperature correction factor [...]sθ Corrected slip value according to reference

coolant temperature [...]p Number of stator pole pairs [...]m Number of machine phases [...]n Harmonic order [...]η Efficiency [...]nr Slot number [...]

FFT Fast Fourier TransformFEM Finite Element MethodFEMM Finite Element Method Magnetics

List of Figures

1.1 Efficiency classes for electrical motors(50 Hz, 4-pole) [5] . . . . . . . . . 21.2 A simple sketch of unskewed and skewed rotors [3] . . . . . . . . . . . . 31.3 Illustration of inter-bar current flow between bars through lamination core 31.4 Starting torque curves with different inter-bar resistivity for a)Skewed

rotor motor b)Unskewed rotor motor[7] . . . . . . . . . . . . . . . . . . 41.5 a) Components of stray-load losses for 0.2-37 kW induction motors

[7],[9]. b) Total rotor losses as a function of rotor skew and inter-barresistivity [3] . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5

2.1 Semi-closed stator slot shape . . . . . . . . . . . . . . . . . . . . . . . . 82.2 Distribution of losses in induction machines . . . . . . . . . . . . . . . . 92.3 Different test methods for efficiency measurement according to IEC 60034-

2-1 standard . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12

3.1 One-pole geometry of different asymmetrical rotor slot arrangementsa)Dual 24-32 rotor slot design b)Progressive sinusoidal rotor slot designc)Dual+progressive sinusoidal rotor slot design . . . . . . . . . . . . . . 14

3.2 One-pole geometry of rotor slots for different rotor designs a) Standardunskewed b) Dual 24-32+progressive c) Progressive sinusoidal . . . . . 16

3.3 Smoothening of residual loss data [14] . . . . . . . . . . . . . . . . . . . 20

4.1 2D FEM simulation results of the starting performance for three rotordesigns . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

4.2 2D FEM simulated flux density in the airgap for three rotor designs . . 244.3 Harmonic spectrum of airgap flux density for three rotor designs . . . . 244.4 2D FEM simulation for the rated torque of different rotor designs . . . . 254.5 Torque harmonics at rated operation for three rotor designs . . . . . . . 264.6 Delta circulating stator currents at rated operation for three rotor designs 284.7 High frequency harmonic components in 2D FEM simulated winding

current Iab at rated operation for three rotor designs . . . . . . . . . . . 29

5.1 Validation of 2D FEM simulation results for the starting performance ofstandard unskewed rotor case . . . . . . . . . . . . . . . . . . . . . . . . 32

vii

viii List of Figures

5.2 2D FEM simulated and measured results for high frequency stator wind-ing currents for standard unskewed rotor case . . . . . . . . . . . . . . . 35

5.3 Measured torque-speed and current-speed curves for standard skewedand standard unskewed rotor motors . . . . . . . . . . . . . . . . . . . . 36

5.4 Measured variation of torque with time at high sampling frequency forskewed and unskewed rotor motors . . . . . . . . . . . . . . . . . . . . . 39

5.5 Measured results for high frequency stator winding current for standardskewed and standard unskewed rotor motors . . . . . . . . . . . . . . . 41

5.6 Measured torque-speed and current-speed curves for three rotor designs 415.7 Torque variation with time for motor with different rotor prototypes . . 435.8 Measured winding current Iab at rated operation for three rotor designs 455.9 High frequency components in measured winding current Iab at rated

operation for three rotor designs . . . . . . . . . . . . . . . . . . . . . . 46

A.1 Rotor slot dimensions . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53

B.1 Single phase equivalent circuit for locked rotor condition . . . . . . . . . 55

List of Tables

3.1 General motor specifications . . . . . . . . . . . . . . . . . . . . . . . . . 15

4.1 Simulated starting properties for the three rotor designs . . . . . . . . . 234.2 Summary of losses at rated operation for three rotor designs . . . . . . . 27

5.1 Starting performance parameters obtained from 2D FEM simulation andmeasurements for standard unskewed rotor case . . . . . . . . . . . . . . 33

5.2 Loss comparison of analytical, simulated and measured results for stan-dard unskewed rotor case . . . . . . . . . . . . . . . . . . . . . . . . . . 34

5.3 Measurement results of starting performance for standard skewed andstandard unskewed rotor motors . . . . . . . . . . . . . . . . . . . . . . 36

5.4 Equivalent circuit parameters for standard skewed and standard unskewedrotor motors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 37

5.5 Measured losses for standard skewed and standard unskewed rotor motors 405.6 Measurement results of the starting performance for three rotor designs 425.7 Equivalent circuit parameters for standard skewed, standard unskewed

and asymmetrical rotor motors . . . . . . . . . . . . . . . . . . . . . . . 425.8 Summary of measured losses for motor with three rotor designs . . . . . 44

6.1 Stray losses for three rotor designs . . . . . . . . . . . . . . . . . . . . . 48

A.1 Slot dimensions and pitch angles for standard rotor . . . . . . . . . . . . 53

ix

Contents

List of Figures vii

List of Tables ix

Contents xi

1 Introduction 11.1 Thesis background . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21.2 Thesis objectives . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51.3 Thesis outline . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

2 Harmonics in Induction Machines, Distribution of Losses & Stan-dard Efficiency Measurement Methods 72.1 Harmonics in induction machines . . . . . . . . . . . . . . . . . . . . 7

2.1.1 Space harmonics . . . . . . . . . . . . . . . . . . . . . . . . . 72.1.2 Slot harmonics . . . . . . . . . . . . . . . . . . . . . . . . . . 82.1.3 Phase belt harmonics . . . . . . . . . . . . . . . . . . . . . . 9

2.2 Distribution of losses in induction machines . . . . . . . . . . . . . . 92.3 Methods for efficiency measurement . . . . . . . . . . . . . . . . . . 102.4 Standards for efficiency measurement . . . . . . . . . . . . . . . . . . 11

3 Asymmetrical Rotor Slot Arrangement & The Investigated Motors 133.1 Asymmetrical rotor slot arrangement . . . . . . . . . . . . . . . . . . 13

3.1.1 Dual rotor slot arrangement . . . . . . . . . . . . . . . . . . . 133.1.2 Progressive sinusoidal rotor slot arrangement . . . . . . . . . 133.1.3 Dual + Progressive rotor slot arrangement . . . . . . . . . . 14

3.2 The investigated motors . . . . . . . . . . . . . . . . . . . . . . . . . 143.2.1 Rotor prototypes . . . . . . . . . . . . . . . . . . . . . . . . . 143.2.2 Simulation tool- FCSmek . . . . . . . . . . . . . . . . . . . . 153.2.3 Method of analysis . . . . . . . . . . . . . . . . . . . . . . . . 16

3.2.3.1 Simulations . . . . . . . . . . . . . . . . . . . . . . . 163.2.3.2 Measurements . . . . . . . . . . . . . . . . . . . . . 17

4 2D FEM Simulation Results 21

xi

xii CONTENTS

4.1 Starting performance for three rotor designs . . . . . . . . . . . . . . 214.2 Performance of motors at rated operation . . . . . . . . . . . . . . . 23

4.2.1 Air-gap flux density . . . . . . . . . . . . . . . . . . . . . . . 234.2.2 Rated torque . . . . . . . . . . . . . . . . . . . . . . . . . . . 254.2.3 Losses at rated operation . . . . . . . . . . . . . . . . . . . . 274.2.4 High frequency stator(delta connected) winding currents . . . 28

5 Measurement Results 315.1 Validation of simulation results . . . . . . . . . . . . . . . . . . . . . 32

5.1.1 Starting performance . . . . . . . . . . . . . . . . . . . . . . . 325.1.2 Losses at rated operation . . . . . . . . . . . . . . . . . . . . 345.1.3 High frequency stator(delta connected) winding currents . . . 35

5.2 Standard skewed verses standard unskewed rotor motors . . . . . . . 365.2.1 Starting performance . . . . . . . . . . . . . . . . . . . . . . . 365.2.2 Losses at rated operation . . . . . . . . . . . . . . . . . . . . 385.2.3 High frequency stator(delta connected) winding currents . . . 40

5.3 Asymmetrical verses symmetrical rotor motors . . . . . . . . . . . . 415.3.1 Starting performance . . . . . . . . . . . . . . . . . . . . . . . 415.3.2 Losses at rated operation . . . . . . . . . . . . . . . . . . . . 435.3.3 High frequency stator(delta connected) winding currents . . . 44

6 Conclusions and Future Work 476.1 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 476.2 Future work . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48

Bibliography 51

Appendices 52

A Rotor slot dimensions for standard rotor prototype 53

B Calculation of equivalent circuit parameters from locked rotortest 55

Chapter 1

Introduction

Electrical machines provide a low cost and efficient solution for converting electricalenergy to mechanical energy and vice versa. From hydro and turbo generators ofseveral hundred mega watts to few watts micro-motors used for surgical applica-tions, electrical machines have a wide range of applications with different powerratings [1]. The huge advancement in power electronics have made it possible tohave highly efficient electric machines and drives with torque and speed control.

Induction motor, invented in 1886, is one of the most commonly used electricalmachine in industrial sector. Direct on-line start, low-cost, simplicity and robust-ness are key features associated with these motors.

Need for high efficiency motorsElectric motors used in industrial applications consume 30% to 40% of generatedelectrical energy of the world. In European countries, the electric motor systemshave 70% share in the total electricity consumption [2]. In Sweden, single and threephase induction motors, ranging from 0.75 to 375 kW , have 90% share in the totalelectricity consumption by all kind of electric motors [3]. These statistics suggestthat a small reduction in the losses of induction motor will have significant impacton the energy consumption, keeping in view the number of motors being manufac-tured and the life time of each motor.

Today, the focus is on producing high efficiency motors to minimize the energyconsumption. According to the new EU regulations, all the motors purchased mustmeet IE2 efficiency standards. It has also been decided that by 2015, the direct-on-line start motors, ranging between 7.5 kW to 375 kW must meet IE3 efficiencystandard or IE2 efficiency standard with variable speed drive attached.[5]. A newversion of efficiency standard IEC-60034-30 [4], including the efficiency classes ofthree-phase single-speed cage-induction motors, has been introduced. The energyefficiency class IE4 (Super-Premium Efficiency) is also under consideration. Thedifferent efficiency classes for synchronous and asynchronous motors (50 Hz 4-pole)

1

CHAPTER 1. INTRODUCTION

Figure 1.1: Efficiency classes for electrical motors(50 Hz, 4-pole) [5]

are shown in figure 1.1.

The emerging calculation methods, powerful computational capacity, the ad-vancement in materials and improved manufacturing techniques have made it pos-sible for the motor designers to work on possible solutions for improving the existingmotor designs in efficiency point of view.

1.1 Thesis background

Induction motors with skewed rotor

Small to medium size induction motors are generally equipped with die-cast alu-minium skewed rotors. Rotor skewing by one stator slot pitch is a common techniqueutilized by motor manufacturers to suppress the asynchronous torques [6] and re-duce the audible noise of the motor[7], [8]. A simple sketch of skewed and unskewedrotors is shown in figure 1.2.

It is known fact that casting process results in low contact resistive paths be-tween the rotor bars through the iron core laminations. Rotor skewing under theseconditions result in the flow of currents through these low contact resistive paths.These are called inter-bar currents or cross-currents which result in losses known asinter-bar current losses.

The concept of inter-bar current flow is illustrated in figure 1.3(a) and 1.3(b).Two small sections between the rotor bars are shown for both skewed and unskewedrotors. In case of an unskewed rotor, the flux(directed into the page) linked by these

2

1.1. THESIS BACKGROUND

s

(a) Unskewed rotor

s

(b) Rotor skewed by one stator slot pitch

Figure 1.2: A simple sketch of unskewed and skewed rotors [3]

sections is independent of the axial position of the rotor, As a result, the currentinduced will flow only between rotor bars and the short-circuit ring, completing thecurrent path, as shown in 1.3(a).

For a skewed rotor, the flux linked by the element and the resulting inducedcurrent both depend upon the axial position of rotor. Considering this effect alongthe total length of the rotor, it can be found that current will flow between the rotor

(a) Unskewed rotor (b) Skewed rotor

Figure 1.3: Illustration of inter-bar current flow between bars through laminationcore

3


(a) (b)

Figure 1.4: Starting torque curves with different inter-bar resistivity for a)Skewedrotor motor b)Unskewed rotor motor[7]

bars taking into consideration the low resistive path as shown in figure 1.3(b). Themagnitude of these currents will depend upon the time derivative of flux creatingthem. The harmonic inter-bar currents produce large asynchronous torques duringthe direct-on-line start of induction motors.

Figure 1.4(a) is an interesting result from [7]. It shows that the total torquefor skewed rotor motors largely depends upon the inter-bar resistivity value. Forlow resistivity values, even skewing is ineffective in suppressing the asynchronoustorques. The resultant asynchronous torques in skewed rotor motors are prolongedeven above their synchronous speed, resulting in a decrease of accelerating torquewith a reduction of pull-out torque. In some cases, the motor could face a startingproblem. For unskewed rotor motors, the starting torque and pull-out torque wasfound to be independent of the inter-bar resistivity as shown in figure 1.4(b). How-ever, large asynchronous torques are present at low speed for unskewed rotors.

Measurements in the past have shown that for small to medium sized inductionmotors, the stray losses can vary from 0.5% to 3% of the total input power [12].The different stray loss components are shown in figure 1.5(a). The surface losses1

contributes around 40% of the total stray losses. The inter-bar current losses con-tribute 30% in the total stray losses. For a 4-pole, 15 kW 36/28 slot inductionmotor the variation of total rotor losses, as a function of rotor skew and inter-barresistivity, is shown in figure 1.5(b) [7], [10]. It can be seen that the stray lossesincrease with increase in skewing angle, reducing the overall efficiency of inductionmotor. These above mentioned reasons make skewing an unattractive choice withdie-cast aluminium rotor bars.

1The surface losses are caused by high frequency flux that results in eddy currents on the rotorsurface. These losses can be suppressed using non-machined rotors [9].

4

1.2. THESIS OBJECTIVES

Surface

losses1

40%

Inter-bar

current losses

30%

Pulsation

losses

17%

High

frequency

losses

10%

Leakage flux

losses

3%

(a) (b)

Figure 1.5: a) Components of stray-load losses for 0.2-37 kW induction motors[7],[9]. b) Total rotor losses as a function of rotor skew and inter-bar resistivity [3]

Unskewed rotors on the other hand help in eliminating inter-bar currents, sim-plify and improve the casting process. The total torque for unskewed rotor motorsis independent of the inter-bar resistivity. However, the drawbacks associated withsymmetrical unskewed rotors are high audible noise and large synchronous and asyn-chronous torques at low speed [6] which can prolong the starting time and limit themotor from reaching its rated speed [3], [7].

Asymmetrical slot rotors have been the focus of research for many years in or-der to gain the advantage of having unskewed rotor bars in terms of increasing theefficiency, suppressing the braking torques and reducing the audible noise in the mo-tor. Studies carried out in [3], [7], [10] and [11] have shown promising results in thisregard. It has been found that by proper asymmetrical modulation of rotor slots,the radial airgap forces associated with high noise level can be suppressed, whilethe joule losses are maintained at a reasonable level with reduction in synchronoustorque at low speed. Hence, induction motor with asymmetrical unskewed rotor isexpected to achieve higher efficiency with acceptable noise level compared to thestandard skewed rotor motor.

1.2 Thesis objectivesThe objective of this thesis is to analyse the performance of a 15 kW , 4-pole,36/28 slot industrial induction motor mounted with standard skewed, unskewed andunskewed asymmetrical rotors. The performance analysis is based on measuring thelosses in the motor at rated operation according to IEC standard test procedures,and torque measurements during a direct-on-line start. Accurate determination of

5


high frequency stator winding currents to determine the high frequency copper lossesare also of interest. Based on the measurement results, a comparative analysis ismade between different rotor types and the differences are highlighted and explained.

1.3 Thesis outlineThe outline of chapters in this thesis report are described as follows:

• Chapter 2 includes details about harmonics in induction machines, the dis-tribution of losses in induction machine and a review on standard efficiencymeasurement methods.

• In Chapter 3, theory about the asymmetrical rotor slots is explained. It alsoincludes details about the investigated motor with different rotor prototypes.The method of analysis used in carrying out the 2D FEM simulations andmeasurements is presented in this chapter.

• Chapter 4 includes the 2D FEM simulation results for the starting perfor-mance and losses at rated operation for the different rotor prototypes. Thesimulation results for the high frequency stator(delta connected) winding cur-rents for the three rotor designs is presented in this chapter.

• Chapter 5 includes the measurement results and the comparisons for theinvestigated motors mentioned in chapter 3. Firstly, the simulation resultspresented in previous chapter are validated against measurements for the stan-dard unskewed rotor case. Secondly, a comparison of measurement results forstandard skewed and standard unskewed motors is presented. Finally, themeasurement results of standard skewed, standard unskewed and asymmetri-cal unskewed rotor motors are compared and explained.

• Chapter 6 includes conclusions and future work.

6

Chapter 2

Harmonics in Induction Machines,Distribution of Losses & StandardEfficiency Measurement Methods

2.1 Harmonics in induction machines

2.1.1 Space harmonics

The harmonics in the airgap magnetic field are known as space harmonics. Theseharmonics cause additional noise, vibration and adversely affect the starting perfor-mance of induction motors [17]. Space harmonics are by-products of the rotatingmagnetic field production mechanism inside electrical machines, hence they cannotbe completely eliminated. However, they can be suppressed to a greater level byproper design optimization. The space harmonics in an electrical machine can pro-duce forward or backward rotating waves, depending upon the harmonic order. Forsymmetrical three-phase machines, the general expression for space harmonics oforder n, created by the stator winding is;

Space harmonics, n = 1 ± 6k (2.1)

where, k = 0, 1, 2, 3, ....

Harmonics such as 7th and 13th result in forward rotating mmf waves with speedequals to ω/7 and ω/13, respectively. On the other hand, 5th and 11th space har-monics result in backward rotating mmf waves with speed equals to ω/5 and ω/11,respectively. The nth space harmonic in general has a speed equal to 1/n times thespeed of the fundamental synchronous speed of the motor, but contains n timesmore peaks than the fundamental. Therefore, the space harmonics and the fun-damental component created by the stator winding have same frequency in statorreference frame [6].

7

CHAPTER 2. HARMONICS IN INDUCTION MACHINES, DISTRIBUTION OFLOSSES & STANDARD EFFICIENCY MEASUREMENT METHODS

The space harmonic of order n induces a current in rotor of frequency given by;

f2(n) = [n(1 − s(1)) ± 1]·f(1) (2.2)

In equation 2.2, the positive(+) or negative (-) sign depends on the value of n givingbackward or forward rotating waves, respectively [6].

Studies have shown that the for 36/28 slot induction motors, the torque dips dueto 7th(space harmonic) and 19th(stator slot harmonic) forward rotating waves canaffect the starting characteristics [7]. If the resultant torque dip becomes lower thanthe required load torque, it can result in an unsuccessful start of the motor.

2.1.2 Slot harmonics

The stator of the induction machine consists of semi-closed slots as shown in figure2.1.

Figure 2.1: Semi-closed stator slot shape

The stator slot openings result in the variation of air gap permeance, distortingthe main flux and create space harmonics. These harmonics are known as thestator slot harmonics. In contrast, with closed rotor slot design, the high frequencyrotor slot harmonics are expected to have low magnitude compared to stator slotharmonics [18]. The general expression for the stator and rotor slot harmonics isgiven by;

Stator slot harmonics = 1 ± n·Qs

p(2.3)

Rotor slot harmonics = 1 ± n·Qr

p(2.4)

where n is an integer representing the order of slot harmonic, p is the number ofpole pairs. Qs and Qr are the total number of stator and rotor slots, respectively.

The stator slot harmonics induces currents in rotor of frequency given by;

[n·Qs

p(1 − s(1)) ± s(1)]·f(1) (2.5)

8

2.2. DISTRIBUTION OF LOSSES IN INDUCTION MACHINES

The rotor slot harmonics induces currents in stator of frequency given by;

[n·Qr

p(1 − s(1)) ± 1]·f(1) (2.6)

where, f(1) is the fundamental line frequency and s(1) is the fundamental slip. In astandard unskewed rotor the slot harmonics result in;

• High frequency currents which can cause an increase in stray losses of themotor [7].

• Production of torque harmonics, noise and vibrations in induction machinedue to interaction of stator fields and rotor harmonic currents [3].

2.1.3 Phase belt harmonicsThey are created due to concentration of MMF in slots [7]. The general expressionfor phase belt harmonics for three phase machine is given by;

n = 1 ± 6k (2.7)

2.2 Distribution of losses in induction machinesThe classification of the different losses occurring in induction machine is well de-scribed in [13] and [14]. The total losses in the induction machine are usually definedas the difference between the input and the output power.

PT = Pin − Pout (2.8)

where PT , Pin and Pout are the total power losses, the input power and the outputpower of the machine in watts, respectively.

According to the IEC 60034-2-1 [14], the total power losses in an induction ma-chine can be mainly categorized as shown in figure 2.2.

Figure 2.2: Distribution of losses in induction machines

9


Constant losses, Pk

The constant losses, as seen from figure 2.2, consist of iron losses, friction andwindage losses. The constant losses always occur in the machine, independent ofload. The iron losses are further categorized as hysteresis losses and eddy currentlosses. These losses are dependent upon the magnetic properties of the material,lamination thickness and frequency. Some harmonic losses occurring on the surfaceof the iron and in the teeth are also included in iron loss category [14].

Load losses, PL

The load losses consist of the ohmic losses occurring in the stator winding and inthe rotor bars. These are given as;

Stator copper losses, Pcu1 = 3·I2s ·Rs (2.9)

Rotor copper losses, Pcu2 = 3·I2r ·Rr (2.10)

Additional load losses or stray load-losses, PLL

The losses that occur at rated load condition in the active metal parts other thanconductor and in active iron are considered as additional load losses or stray-loadlosses. The load current dependent flux pulsations that cause eddy currents are alsoconsidered in this category of losses [14].

2.3 Methods for efficiency measurementTwo common methods for determining the efficiency of electrical machines are pre-sented in [16]. These are;

• Direct method

• Indirect method

In case of the direct method, the efficiency of the motor is determined simply byfinding the ratio of input active electric power and the output mechanical power.

η = Pout

Pin(2.11)

where, Pout=T ·wr is output mechanical power in watts. Pin is the input electricalpower in watts.

The indirect method of efficiency measurement is based on the input active electricpower and the total losses in the machine calculated from tests.

η = Pin − Ploss

Pin(2.12)

10

2.4. STANDARDS FOR EFFICIENCY MEASUREMENT

where, the total losses consist of the constant losses, the load losses and the addi-tional stray load-losses.

Ploss = PT = PK + PL + PLL (2.13)

Indirect method is considered more accurate for determining the efficiency ofmachines; provided that the stray-load losses are determined accurately [16].

2.4 Standards for efficiency measurementFour major standards for measuring the efficiency of electrical machines are pre-sented in [16]. These are;

• IEEE 112

• IEC 60034-2-1

• CSA C390

• JEC37

IEC 600-34-2-1 and IEEE 112B standards include both direct1 and indirectefficiency measurement methods. IEC 600-34-2-1 was introduced in September 2007and it includes the method to determine the additional stray losses. The olderversion of IEC standard, IEC 600-34-2, defined stray losses as 0.5% of the totallosses occurring in the machine.

Comparison of IEEE and IEC standards

A comparison between the loss determination methods, utilized in IEC and IEEEstandards [15] [16], are presented as follows.

• In order to determine the stator copper losses, the IEEE standard requiresthe measurement of reference stator winding in cold condition and at differentoperating conditions along with temperature measurements. This requires thetemperature sensor to be used during the test hence, this procedure cannotbe performed on motors that are in service. On the other hand, according toIEC standard test method, the resistance is directly measured for the highestand the lowest operating points, without the need of temperature sensors.

• In IEC standard, the method to find the iron losses includes the factor ofvoltage drop across the stator resistance; as the load on the motor changes. Onthe other hand, the IEEE standard considers the iron losses to be independentof load, which could lead to inaccurate results.

1The direct method is known as "Method A" in the IEEE 112 standard.

11


• In case of stray losses, the two standards deploy similar techniques. Firstthe residual stray losses in the machine are determined. A linear regressionanalysis technique is then used to smoothen the curve in order to determinethe actual stray losses in the machine [16]. A small difference between the twostandards is the correlation coefficient value that is used in the curve fittingprocess. IEEE 112 specifies a value of 0.9 while IEC uses 0.95 [15].

• Regarding the methods to find the rotor copper and friction losses, both stan-dards describes similar techniques.

In this thesis work, the test methods for determining losses and efficiency ofdifferent machines are followed according to IEC 600-34-2-1 standard [14]. A blockdiagram showing different test methods for efficiency measurement according to IECstandard are shown in figure 2.3.

Efficiency measurement

methods [IEC 60034-2-1]

Direct

Method

Torque Meter Test

Dynamometer Test

Dual Supply Back to

Back Test

Indirect

Method

Calorimetric Method

Single Supply Back to

Back Test

No-Load Test

Heat-run test (Load

test

Equivalent Circuit

Method

Reversed Rotation and

Removed Rotor Test

EH-Star Test

Load test with torque measurment

(residual loss method)

Gives total losses

Gives constant losses

Gives stray losses

Gives load losses

Figure 2.3: Different test methods for efficiency measurement according to IEC60034-2-1 standard

12

Chapter 3

Asymmetrical Rotor Slot Arrangement& The Investigated Motors

3.1 Asymmetrical rotor slot arrangementThe noise and vibration can be reduced to a greater extent by introducing asym-metry in rotating structure [3]. Asymmetry can be introduced in both rotor andstator, later is more complex from the manufacturing point of view. A detail methodof introducing asymmetry in rotor slots has been presented in [3]. Based on thisstudy an improved modulation function for designing asymmetrical rotors has beendeveloped in [10]. The different asymmetrical rotor slot arrangements are discussedas follows.

3.1.1 Dual rotor slot arrangement

The main theme of this type of asymmetrical rotor slot design is to combine twodifferent rotor slot combinations, with their respective slot pitch angles, into a singlerotor. In this way, the combined advantages of both rotor slot arrangements canbe achieved. An example of a dual rotor slot geometry, combining 24 and 32 rotorslots in a single rotor design, is shown in figure 3.1(a).

3.1.2 Progressive sinusoidal rotor slot arrangement

Asymmetrical rotor slot arrangement can also be achieved by varying the slot pitchangles and corresponding slot width, as a function of sine-wave as given in equation3.1 and equation 3.2, respectively [10].

θ(nr) = θ0(nr) +n∑

i=1Ki sin( i.θ0(nr) + δi) (3.1)

bnr = b0(nr) θ(nr + 1) − θ(nr)θ0(nr + 1) − θ0(nr)

(3.2)

13

CHAPTER 3. ASYMMETRICAL ROTOR SLOT ARRANGEMENT & THEINVESTIGATED MOTORS

where nr is the slot number, θ0 is the initial slot pitch angle and b0 is the initialslot width. Ki and δi are the modulation coefficients.

An example of progressive rotor slot arrangement is shown in figure 3.1(b).

3.1.3 Dual + Progressive rotor slot arrangement

Another method to introduce asymmetry in rotor slots is by combining the prop-erties of both dual and progressive asymmetrical rotor slot arrangements. Thisresults in design named as dual+progressive rotor slot arrangement. This is shownin 3.1(c).

(a) (b) (c)

Figure 3.1: One-pole geometry of different asymmetrical rotor slot arrange-ments a)Dual 24-32 rotor slot design b)Progressive sinusoidal rotor slot designc)Dual+progressive sinusoidal rotor slot design

3.2 The investigated motorsThe general specifications of the four motors investigated in this work are shownin table 3.1. The stator was kept the same in all the tests with different rotorprototypes.

3.2.1 Rotor prototypes

The four rotor prototypes namely; a)Standard unskewed b)Standard skewed c) Dual24-32+progressive d)Progressive sinusoidal were investigated. Some details aboutthe rotor slot arrangement, for each prototype, is given as follows:

• The two prototypes: standard skewed and standard unskewed rotors havesame rotor slot design.

14

3.2. THE INVESTIGATED MOTORS

Table 3.1: General motor specificationsRated power, PN 15 kWRated voltage, UN 400 VRated frequency, fN 50 HzNo. of stator poles 4Stator winding Delta connected

3-cross concentricNo. of stator slots 36No. of rotor slots 28Core material M600-50AMachine length 205 mmShaft height 160 mmAirgap length 0.55 mmInsulation class B

• The dual 24-32+progressive rotor prototype have asymmetrical rotor slots.The rotor slot dimensions and slot pitch angles were obtained by setting mod-ulation coefficients i = 24 and K24 = 1.7 in equation 3.2.

• The rotor slot dimensions and slot pitch angles for progressive sinusoidal rotorprototype were obtained by setting modulation coefficients i = 4, 8 ; K4 = 2.65and K8 = 2.2 in equation 3.2.

The slot widths and slot pitch angles for each prototype are given in appendixA. The idea behind choosing particular modulation coefficients for asymmetricalrotor slot design is to suppress the synchronous torques at zero speed as presentedin [10].

3.2.2 Simulation tool- FCSmekA 2D FEM simulation tool FCSmek1 was used to simulate the different rotor designsand the performance of motors were analyzed. Some features of FCSmek and thegeometry creation process in FCSmek are illustrated as follows:

• The losses simulated in FCSmek includes the losses due to high frequencycomponents of current.

• At present, it is not possible to simulate skewed rotor cases in FCSmek.

• The simulation time largely depends upon the time step chosen for each sim-ulation and the mesh settings including size, shape & order of element.

1FCSmek is a base program in ADEPT. ADEPT is an in-house ABB calculation tool.

15


• In FCSmek, the stator and rotor geometries are created using FEMM2. De-pending upon the machine input parameters in ADEPT, FCSmek generatesthe FEMM files for the corresponding stator and rotor geometries along witha LUA script [20] file readable in FEMM.

• The LUA script file for the standard rotor geometry was modified in order toobtain the desired asymmetrical rotor geometries.

The one-pole geometry of standard stator and different rotor designs, created inFCSmek are shown in figure 3.2.

(a) (b) (c)

Figure 3.2: One-pole geometry of rotor slots for different rotor designs a) Standardunskewed b) Dual 24-32+progressive c) Progressive sinusoidal

3.2.3 Method of analysis

The method of analysis used during simulation and measurement work is presentedin this section.

3.2.3.1 Simulations

The starting performance of the machine with different rotor designs were simu-lated using a time step of 250 µ sec for a total of 80 fundamental periods. Thiscorresponds to a sampling frequency of 4 kHz. The simulation was initialized witha rotor speed of -750 rpm. The purpose of doing this is to avoid the switchingtransients in the motoring region during the direct on-line start condition and alsoto capture the asynchronous torques, due to 5th space harmonics, occurring in neg-ative speed range. A rotor inertia value of 2.25 Kgm2 was set in order to reduce theacceleration time of the rotor and to capture the different phenomenon of interest in

2FEMM is a free-ware tool known as Finite Element Method Magnetics [19].

16


the torque curve. In general, a higher value of rotor inertia is avoided as it prolongsthe simulation time.

The starting performance curves obtained from 2D FEM simulations contain highfrequency components in it. A 3rd order low-pass butter worth filter was used inorder to filter out these curves. The sampling frequency, fs for the low pass filterwas set to 50 Hz with cut-off frequency, fc to 35 Hz.

For rated operation, the simulations were carried out with time stepping of 100µ sec for a total of 200 periods. The simulation results were used in order to de-termine the high frequency stator winding currents, air-gap flux density, torqueharmonics and losses in the machines at rated operation.

3.2.3.2 Measurements

Starting performance

The torque-speed curves for the different cases were recorded by operating the motorbetween rated and zero speed range. However, due to the low sampling frequencyof recording equipment in the laboratory, it was difficult to determine the impactof synchronous and asynchronous torques at low speed. To capture these low speedparasitic effects, an oscilloscope with a 5 kHz sampling frequency was used tocapture the torque variation with time for different cases.

High frequency stator winding currents

An oscilloscope of 50 kHz sampling frequency was used to capture the high fre-quency currents, flowing in stator winding, for different cases.

Efficiency

To determine efficiency, indirect method was used in this study. The standardmethods used to determine individual loss components given in equation 2.13, arediscussed as follows.

1. Constant losses, PK using no-load test: The constant losses, PK consistsof frictional and windage losses, Pfw and the iron losses, Pfe. These losses aredetermined by performing a no-load test [14] on the machine, recording thevariation of input power at no-load with terminal voltage, Uo. The voltagewas varied between 20% to 125% of the rated voltage. During no-load test,the total losses in the machine equal the no-load input power, Po, and is thesum of stator copper losses, Ps and the constant losses, PK . The constantlosses are given as shown in equation 3.3.

PK = Po − Ps = Pfe + Pfw (3.3)

17


Ps = 1.5·Io2·R||o (3.4)

where Io, is the no-load average line current in amperes. R||o, the line to lineresistance in ohm, is determined after the lowest voltage reading during theno-load test.

Using the no-load loss points that show no significant saturation effect, acurve is developed showing relationship between the constant losses and thesquare of no-load terminal voltage. This curve is extrapolated to zero-voltageaxis and the intercept to that gives the friction and windage losses in themachine. These losses are independent of the load.

After finding the friction and windage losses, another curve can be plottedfor different load points giving relationship between Pfe = PK-Pfw and theno-load terminal voltage, Uo. At any desired load point, the iron losses can befound corresponding to voltage, Ur taking resistive voltage drop in the statorwinding into account as given in equation 3.6.

Ur =

√(U −

√3.I.R. cos φ

2)2 + (

√3.I.R. sin φ

2)2 (3.5)

where,cos(φ) = P1√

3.U.I, sin(φ) =

√1 − cos(φ)

U is the average terminal voltage, P1 is the input power in watts, R is thewinding resistance in ohms and I is the average line currents in amperes.

2. Load losses, PL using heat-run test: The load losses include the stator cop-per losses, Ps and rotor copper losses, Pr, at rated load operation. During theload test or heat-run test, the machine is loaded at rated supply power andis operated until thermal equilibrium is achieved (gradient of 2 K per hour).At the end of test the power, current, voltage, resistance, slip, frequency andwinding temperature at rated conditions are noted.

The incorrect stator copper losses are determined at any load point usingequation 3.6.

Ps = 1.5·I2·R (3.6)

The stator winding resistance is corrected to reference coolant temperature of25oC to obtain the correct value of stator copper losses;

Ps,θ = Ps·kθ (3.7)

kθ = 225 + θw + 25 − θc

225 + θw(3.8)

18


where, kθ is the temperature correction factor for winding, θc is the inletcoolant temperature during test and θw is the reference winding temperaturedepending upon the insulation class.

Similarly, the rotor copper losses for a given load point is determined usingequation 3.9.

Pr = (P1 + Ps + Pfe)·s (3.9)

The corrected rotor copper losses are given as;

Pr,θ = (P1 − Ps,θ − Pfe)·sθ (3.10)

sθ = s·kθ (3.11)

3. Stray losses, PLL using residual loss method: The additional load lossesare determined using partial load test with torque measurement. Firstly,the residual losses, PLr for each load point are determined by subtracting theoutput power, load losses and constant losses from the input power. This isshown in equation (3.12).

PLr = P1 − P2 − PL − PK (3.12)

The residual losses can be expressed as a function of square of torque givenas;

PLr = A·T 2 + B (3.13)

The losses given by equation 3.13 are smoothed using linear regression analysisas shown in figure 3.3. In figure 3.3, A is the slope, B is the intercept, i is thenumber of load points and γ is the correlation coefficient. Once the value of Ais stable, the value of additional load losses for every load point is determinedby the following expression.

PLL = A·T 2 (3.14)

19


Figure 3.3: Smoothening of residual loss data [14]

20

Chapter 4

2D FEM Simulation Results

4.1 Starting performance for three rotor designsIn this section the simulated starting performance and key performance character-istics of the investigated motors having different rotor designs is presented. As thesimulation of skewed rotor case needs 3D FEM tool, whereas FCSmek is a 2D FEMtool, only the unskewed cases (symmetrical and asymmetrical) were possible to sim-ulate using FCSmek. The 2D FEM simulated starting performance curves for thethree different rotor designs are shown in figure 4.1. The rotor speed was varied from-750 rpm to 1500 rpm as shown in figure 4.1(a) and the corresponding torque varia-tion for three cases, at sampling frequency of 4 kHz, is shown in figure 4.1(b), 4.1(c)and 4.1(d). For comparative analysis, the higher order frequency components werefiltered out using a low-pass filter. The 2D FEM simulated current-speed curves forthree rotor designs are also shown in figure 4.1(f).

Figure 4.1(e) shows the filtered starting torque-speed curves for the three rotordesigns. It can be observed that:

• The impact of 5th, 7th space harmonic and 17th, 19th stator slot harmonicscan be seen in the resulting torque speed curves. The 7th and 19th harmonicshaving forward rotation create torque dips in the motoring region.

• At zero speed, the synchronous torques for standard unskewed rotor are negli-gible. For asymmetrical rotor designs, small synchronous torques are presentat zero speed. However, the synchronous torques at zero speed are well sup-pressed with chosen modulation coefficients, compared to results presented in[11]. Hence, the motors with asymmetrical rotor prototypes were expected toundergo successful start during prototype testing.

• At 214 rpm, (17 of the synchronous speed of motor), there are big fluctuations

in the torque-speed curves for all the cases. Synchronous torques are experi-enced by all the rotor designs at this speed due to 28 rotor slots. Synchronoustorque is also present around 150 rpm for all rotor designs due to 36 stator

21

CHAPTER 4. 2D FEM SIMULATION RESULTS

0 0.5 1 1.5 2 2.5−750

−500

−250

0

250

500

750

1000

1250

1500

Time [s]

Spee

d[r

pm

]

(a) Speed variation with time

0 0.5 1 1.5 2 2.5−400

−200

0

200

400

600

800

1000

Time [s]

Tor

que

[Nm

]

Original

Filtered

(b) Standard unskewed

0 0.5 1 1.5 2 2.5−400

−200

0

200

400

600

800

1000

Time [s]

Tor

que

[Nm

]

Original

Filtered

(c) Dual 24-32+progressive

0 0.5 1 1.5 2 2.5−400

−200

0

200

400

600

800

1000

Time [s]

Tor

que

[Nm

]

Original

Filtered

(d) Progressive sinusoidal

−500 −250 0 250 500 750 1000 1250 1500−100

0

100

200

300

400

500

600

700

Speed [rpm]

Tor

que

[Nm

]

Standard unskewedDual 24−32+progressiveProgressive sinusoidal

Pull−out torque

Starting torque

(e) Torque-speed curves comparison

0 500 1000 15000

50

100

150

200

250

Speed [rpm]

Curr

ent

[A]

Standard unskewedDual 24−32+ProgressiveProgressive sinuosidal

(f) Current-speed curves comparison

Figure 4.1: 2D FEM simulation results of the starting performance for three rotordesigns

22

4.2. PERFORMANCE OF MOTORS AT RATED OPERATION

slots. It can also be seen that the synchronous torques are suppressed for theprogressive sinusoidal rotor design compared to the other two rotor designs.This shows that the chosen modulation coefficients, for introducing asymme-try in rotor designs, not only helped in suppressing the synchronous torquesat zero speed but also suppressed the synchronous torques due to 7th spaceharmonic.

• It is difficult to analyse the asynchronous torque created by 7th space harmonicfor the three cases, because of the synchronous torque occurrence at the samespeed.

The values of starting torque, pull-out torque and starting current(RMS), sim-ulated for three rotor designs, are summarized in table 4.1. These results are con-sistent with the findings in [10] and show that the studied asymmetrical rotors arecapable of starting during direct-on-line.

Table 4.1: Simulated starting properties for the three rotor designsStarting Maximum Starting

Rotor design torque torque current

[Nm] [Nm] [A]

Standard unskewed 258 303 209.7

Dual 24-32+progressive 237 303 209.2

Progressive sinusoidal 253 306 208.5

4.2 Performance of motors at rated operationThis section contains the performance of the investigated motors, having differentrotor designs, simulated at rated operation.

4.2.1 Air-gap flux densityFigure 4.2 shows the simulation results of the airgap flux density of motor withthree rotor designs for one electrical cycle. The fundamental component of airgapflux density can also be seen for each case. In order to analyse the harmonic contentwith their possible sources in the airgap flux density, FFT spectral analysis of fluxdensity for the three rotor designs is shown in figure 4.3. It can be observed that:

• The value of the fundamental component of flux density was approximately0.75 T for all the three cases. In figure 4.3, the magnitude of higher orderflux density components is presented in terms of percentage of fundamentalcomponent value (taken as 100%).

23


0 0.5 1 1.5 2 2.5 3

−1

−0.5

0

0.5

1

Space co-ordinate in the air gap [radians]

Flu

xden

sity

[T]

Standard unskewed

0 0.5 1 1.5 2 2.5 3

−1

−0.5

0

0.5

1


Flu

xden

sity

[T]

Dual 24−32+progressive

0 0.5 1 1.5 2 2.5 3

−1

−0.5

0

0.5

1


Flu

xden

sity

[T]

Progressive sinusoidal

Figure 4.2: 2D FEM simulated flux density in the airgap for three rotor designs

1 3 5 7 9 11 13 15 17 19 21 23 25 27 29 31 33 35 37 390

5

10

15

20

25

30

35

40

45

50

Harmonic number

%am

plitu

de

offu

ndam

enta

lflux

den

sity

Standard


Progessive Sinuosidal

1 ± 2m

1 ±Qs

p

1 ± 2Qr

p

1 ± 2Qs

p

1 ±Qr

p1 ± 4m

Figure 4.3: Harmonic spectrum of airgap flux density for three rotor designs

24


• The 5th and 7th harmonics are due to the 1st order phase belt harmonics. Thegeneral expression for these type of harmonics is 2mn±1, where m representstotal number of phases of machine and n represents the order of harmonic.Similarly, the 11th and 13th harmonics are due to second order phase beltharmonic, with n equal to 2.

• The 17th, 19th, 35th and 37th harmonics are the stator slot harmonics, givenby equation 2.3. These are due to 36 stator slots.

• The 13th, 15th, 27th and 29th harmonics are the rotor slot harmonics, givenby equation 2.4. These are due to 28 rotor slots.

• For progressive sinusoidal rotor, the 17th harmonic has the highest magnitude,around 33% of fundamental, compared to other rotor designs.

4.2.2 Rated torqueThe rated torque can be calculated analytically using equation 4.1.

TN = PN

2·π·Nr·60 = 97.44 Nm (4.1)

where,PN = 15kW (rated power)Nr = 1470 rpm (rated speed)TN is rated torque in Nm.

The rated torque computed from equation 4.1 is 97.44 Nm. The 2D FEM sim-ulated torque variation at the rated operation, for the three rotor designs, is shownin figure 4.4. It can be seen that the rated torque obtained from the 2D FEMsimulation is in good agreement with the calculated value from equation 4.1.

0.99 0.992 0.994 0.996 0.998 160

70

80

90

100

110

120

130

140

150

Time [s]

Tor

que

[Nm

]

Standard unskewedDual 28−32+progressiveProgressive sinusoidal

Figure 4.4: 2D FEM simulation for the rated torque of different rotor designs

25


Figure 4.5 shows the harmonic content in the output torque for the three rotordesigns. The torque harmonics can be explained with respect to the air gap fluxdensity harmonics shown in figure 4.3.

0 10 20 30 400

5

10

15

20

Harmonic number

%Tor

que

amplitu

de

Standard unskewed

(a) Standard unskewed

0 10 20 30 400

5

10

15

20

Harmonic number%

Tor

que

amplitu

de


(b) Dual 24-32+progressive

0 10 20 30 400

5

10

15

20

Harmonic number

%Tor

que

amplitu

de

Progessive Sinusoidal

(c) Progressive sinusoidal

Figure 4.5: Torque harmonics at rated operation for three rotor designs

From figure 4.5, it can be seen that;

• The 6th torque harmonic, resulting from 5th and 7th air gap flux harmonic, isnegligible for all the three cases.

• For standard unskewed rotor, 12th harmonic, due to 11th and 13th airgap fluxharmonics, is the only significant component with 12% amplitude comparedto fundamental. For progressive sinusoidal rotor design, the 12th harmonic isaround 7% of the fundamental harmonic and is lower compared to standardunskewed and dual rotor designs.

• For dual 24-32+progressive rotor design, the 12th and 18th harmonic are dom-inant with amplitude around 11% of the fundamental.

26


• For progressive sinusoidal rotor design, the 18th harmonic has the highestmagnitude, with amplitude of 16% of the fundamental. The main reason forthis torque harmonic is the higher 17th order air gap flux harmonic as seen infigure 4.3.

• The 28th torque harmonic has same amplitude for both asymmetrical rotordesigns.

• The higher magnitude of torque harmonics for the asymmetrical rotor designsmay lead to an increase in noise and vibrations in motor.

4.2.3 Losses at rated operationThe different losses occurring in the stator and rotor along with other performancecharacteristics are summarized in table 4.2.

Table 4.2: Summary of losses at rated operation for three rotor designsParameter Standard Dual 24-32 Progressive

unskewed +progressive sinusoidal

Total stator I2R losses [W ] 515 522 529

Fundamental rotor I2R losses [W ] 309 310 310

Total rotor I2R losses [W ] 415 445 468

Total iron losses [W ] 382 377 370

Friction losses [W ] 69 69 69

Stray losses [W ] — — —

Total losses [W ] 1381 1413 1436

Output power [kW ] 15 15 15

Efficiency [%] 91.6 91.4 91.3

Cosϕ 0.84 0.83 0.83

Winding temperature [0C] 74 74 74

Fundamental stator current [A] 28.5 28.6 28.8

Torque [Nm] 97.5 97.4 97.5

From table 4.2, it can be seen that:

• For the dual 24-32+progressive and progressive sinusoidal rotor designs, thetotal rotor copper losses increased by 7.2% and 12.7%, respectively comparedto standard unskewed rotor. The major reason of this increase in losses is thehigh frequency currents that are induced in the rotor with the introduction of

27


asymmetry in slot designs. These high frequency currents are due to the har-monics with higher magnitude in the air gap flux density for the asymmetricalrotor designs compared to standard rotor, as shown in figure 4.3.

• The iron losses are slightly lower for asymmetrical rotor designs compared tostandard unskewed rotor; this difference is negligible.

• The total losses for dual 24-32+progressive and progressive sinusoidal rotordesigns are 2.3% and 4% higher, respectively compared to standard unskewedrotor.

However, it was expected that during the prototype testing the above mentionedincrease in rotor copper losses for asymmetrical rotor designs will be much lowercompared to the increase in stray losses for standard skewed rotor. Thus an overallreduction of losses for asymmetrical rotor designs compared to standard skewedrotor was expected.

4.2.4 High frequency stator(delta connected) winding currents

The delta connected stator winding arrangement for induction machine is shown infigure 4.6(a). For comparison purpose, the simulated winding current Iab in one ofthe three phases for all cases is shown in figure 4.6(b).

Ia

Ib

Ic

Iab

Ibc

Ica

(a) Delta-connected stator winding

0.05 0.06 0.07−40

−20

0

20

40

Time[s]

win

din

gcu

rren

t,Iab

[A]

Standard unskewedDual 24−32+progessiveProgressive sinusoidal

(b) 2D FEM Simulated winding current Iab

Figure 4.6: Delta circulating stator currents at rated operation for three rotor de-signs

Additional high frequency harmonic components are expected to be present inthe winding currents of all rotor designs. FFT was performed on the winding currentfor all the three cases and the high frequency components having magnitude above5% of the fundamental are shown in figure 4.7. It can be seen that:

28


144 146 148 150 152 1540

2.5

5

7.5

10

12.5

15

Frequency[Hz]

%offu

ndam

enta

lw

indin

gcu

rren

t,Iab

Progressive sinusoidalDual 24−32+progessiveStandard unskewed

(a)

630 632 634 636 638 6400

2.5

5

7.5

10

12.5

15

Frequency[Hz]

%of

fundam

enta

lw

indin

gcu

rren

t,Iab


(b)

730 732 734 736 738 7400

2.5

5

7.5

10

12.5

15

Frequency[Hz]

%of

fundam

enta

lw

indin

gcu

rren

t,Iab


(c)

830 832 834 836 838 8400

2.5

5

7.5

10

12.5

15

Frequency[Hz]

%of

fundam

enta

lw

indin

gcu

rren

t,Iab


(d)

Figure 4.7: High frequency harmonic components in 2D FEM simulated windingcurrent Iab at rated operation for three rotor designs

• The magnitude of frequency component at 150 Hz, due to 3rd order harmonic,is almost the same for the three cases.

• The frequency components at 636 Hz and 736 Hz have the highest magnitudefor standard unskewed rotor and are slightly lower for progressive sinusoidalrotor. These high frequency components are due to 1st order rotor slot har-monics, as given by equation 2.6.

• For asymmetrical rotor designs, additional high frequency components ap-peared at 734 Hz and 834 Hz. The magnitude of these harmonics is higherfor progressive sinusoidal design compared to dual 24-32+progressive design.According to equation 2.6, these high frequency harmonics should appear fora 32 slot rotor. This shows that for asymmetrical rotors, the dominant rotorslot harmonic is not always due to the apparent slot numbers, but it is possi-

29


ble to have an inherent slot number with a significantly high slot permeancevalue [10].

• For asymmetrical rotors, these additional high frequency components can bea source of further increase in stator copper losses.

30

Chapter 5

Measurement Results

As mentioned in the previous chapters, the major objective of this project is to testthe performance of standard industrial induction motor mounted with four differentrotor prototypes. The measurements were performed according to IEC 60034-2-1standard. The stator was kept the same in all the measured cases. The standardskewed and unskewed rotor motors were tested to study the influence of rotor skewon the motor performance. The two asymmetrical slot rotors were tested to studythe influence of different slot modulation concepts on the motor performance com-pared to the standard rotors.

The slot shapes on the lamination sheets for asymmetrical slot rotors were man-ufactured using laser cut technique, unlike the standard rotor slots which werepunched. The laser cut manufactured rotor laminations for asymmetrical rotorsleft some sharp edges on one of the lamination sides, which led to ineffective castingprocess. Aluminium had leaked between the laminations and at the core edges.As a result, one of the prototype rotor, the dual(24-32)+ progressive prototypewas not successfully manufactured and the other asymmetrical prototype had poorcasting issues. To summarize, the three prototypes; a)Standard skewed b)Standardunskewed c)Progressive sinusoidal were tested.

The measurement results include the torque-speed and current-speed measurementsfor determining the starting performance of the different rotor motors. Locked ro-tor test was performed on each rotor to determine the locked rotor torque and toextract equivalent circuit parameters. The no-load, heat run and partial load testswere performed for determining different losses in the motor. An additional test formeasuring the high-frequency currents, flowing in the delta-connected stator for allthe motors, was also performed. This test gives a good idea about the impact ofrotor slot harmonics on the stator winding currents and the effect of rotor skewingon these harmonics.

31

CHAPTER 5. MEASUREMENT RESULTS

5.1 Validation of simulation results

5.1.1 Starting performance2D FEM simulation and measurement results for torque-speed and current-speedcurves for standard unskewed rotor are shown in figure 5.1(a) and 5.1(b), respec-tively.

0 250 500 750 1000 1250 15000

100

200

300

400

500

Speed [rpm]

Tor

que

[Nm

]

Simulated (2D FEM)Measured (IEC 60034−2−1)

Starting torque

Pull−out Torque

(a) Torque-speed curve

0 250 500 750 1000 1250 15000

50

100

150

200

250

Speed [rpm]

Curr

ent

[A]

Simulated (2D FEM)Measured (IEC 60034−2−1)

(b) Current-speed curve

1 1.5 2 2.5−100

0

100

200

300

400

500

Time [s]

Tor

que

[Nm

]

2D FEM simulated−original2D FEM simulated−filtered

(c) 2D FEM simulation

0 5 10 15 20 25−100

0

100

200

300

400

500

Time [s]

Tor

que

[Nm

]

Measured−originalMeasured−filtered

(d) Measurement results

Figure 5.1: Validation of 2D FEM simulation results for the starting performanceof standard unskewed rotor case

From figure 5.1(a), the simulation results include the impact of synchronoustorques at 150 rpm and 214 rpm (due to 36/28 slot combination) [11], and alsothe impact of asynchronous torques due to 7th and 19th harmonics (with forwardrotating wave) in the motoring region [7]. However, it was not possible to analyzethe impact of synchronous and asynchronous torques at low speed in the measuredtorque-speed curve due to low sampling frequency of the recording equipment. An

32

5.1. VALIDATION OF SIMULATION RESULTS

oscilloscope with a high sampling frequency, was used to record the measurementresults of the torque variation with respect to time. The motor was simulated fromzero to rated speed, whereas, from rated speed to zero speed during measurements.The simulation results are shown in figure 5.1(c) and the corresponding measure-ment results are shown in figure 5.1(d).

Table 5.1: Starting performance parameters obtained from 2D FEM simulation andmeasurements for standard unskewed rotor case

Analytical Simulated Measured

Quantity (OSKAR) 2D FEM (IEC

(time step= 60034-2-1)

0.25msec)

Starting torque [Nm] 292 258 286

Maximum torque [Nm] 311 303 320

Starting current(RMS) [A] 216 210 216

The analytical, simulated and measured results for starting performance aresummarized in table 5.1. The reasons for difference in measurement and simulationresults for starting performance, seen from figure 5.1 and listed in table 5.1, areexplained as follows:

• The simulation results for the starting torque and maximum torque are 11%and 8% lower, respectively, compared to the measurement results as seen fromfigure 5.1(c) and 5.1(d). One reason for this high starting torque obtained inmeasurement result could be a higher starting current value measured com-pared to simulation as shown in figure 5.1(b). Additionally, the simulationwas carried out with no-load settings for the motor. This is evident from thesimulation results shown in figure 5.1(c) where the torque value is zero atsteady state. However, there is a DC-offset of 20 Nm still present in the mea-sured results which shows the motor was not ideally under no-load condition.The other possibility is that the torque meter was not calibrated exactly tozero before starting the measurements.

• The time for the motors to reach from zero speed to the rated speed is differentin simulation and measurement results. This is due to different rotor inertiasettings in both cases. As a result, the rate of rotor speed during simulationand measurement was different.

• The properties of low pass filter (mentioned in chapter 4), used to filter out thehigh frequency component from the torque curve, can also marginally effectthe absolute values in the filtered curve.

33


Table 5.2: Loss comparison of analytical, simulated and measured results for stan-dard unskewed rotor case

Analytical Simulated MeasuredQuantity (OSKAR) 2D FEM (IEC

(time step= 60034-2-1)0.1msec)

Total I2R stator losses [W ] — 515 —Fundamental I2R stator losses [W ] 528 — 538Fundamental I2R rotor losses [W ] 313 309 277Total I2R rotor losses [W ] — 415 —Total iron losses [W ] 311 382 325Friction losses [W ] 69 69 69Stray losses [W ] 316 — 309Total losses [W ] 1537 1381 1518Output power [kW ] 15 15 15Efficiency [%] 90.7 91.6 90.9Cosϕ 0.84 0.84 0.83Winding temperature [0C] 74 74 74Current [A] 28.4 28.5 28.7Torque [Nm] 97.4 97.5 97.2

5.1.2 Losses at rated operation2D FEM simulated and measured losses at rated operation, for standard unskewedrotor motor, are presented in table 5.2. Results from analytical calculations1 arealso listed in this table. From table 5.2, it can be observed that;

• The total I2R losses in stator and rotor determined by 2D FEM simulationincludes the fundamental and high frequency current losses. But, losses onlydue to fundamental component are included in the measurement results.

• For comparisons, the fundamental I2R losses for rotor was obtained as afunction of fundamental slip and airgap power from total simulated I2R losses.The simulated fundamental I2R rotor losses are 11% higher compared tomeasured value. The reason for this could be the difference in rotor bar andend-ring resistance for measured prototype compared to the simulation.

1ABB in-house analytical tool OSKAR was used in this study. The calculated quantities inOSKAR are adjusted with correction factors provided in OSKAR. Hence the analytical resultsserves as good starting reference.

34

5.1. VALIDATION OF SIMULATION RESULTS

• The total iron losses determined by FEM are 17% higher compared to mea-sured results. The measured iron losses can be different from simulated, de-pending upon the magnetic shaft properties which changes with rotor tem-perature. The magnetic properties of lamination sheets can also vary duringpunching and casting process, effecting the iron losses in motor [1].

• The friction losses are given as input value in analytical and in 2D FEM. Itwas set according to measured value.

• The stray losses were determined separately in measurement and analyticalmethods, which are in good agreement to each other. It was not possible todetermine the stray losses separately in 2D FEM simulation. Some compo-nents of stray load losses like the additional high frequency losses are includedin the total stator copper, rotor copper and iron losses in 2D FEM simulations.


The 2D FEM simulated and measured results in one of the stator windings areshown in figure 5.2(a). The corresponding harmonic analysis of the simulated andmeasured results is shown in figure 5.2(b). For simplicity, current in only one of thethree stator windings is considered in this study.

0 0.005 0.01 0.015 0.02 0.025 0.03 0.035 0.04−40

−30

−20

−10

0

10

20

30

40

Time[s]

win

din

gcu

rren

t,Iab

[A]

SimulatedMeasured

(a)

0 100 200 300 400 500 600 700 800 900 10000

2.5

5

7.5

10

12.5

15

Frequency[Hz]

%am

plitu

de

ofw

indin

gcu

rren

t,Iab

SimulatedMeasured

636 Hz

736 Hz

(b)

Figure 5.2: 2D FEM simulated and measured results for high frequency statorwinding currents for standard unskewed rotor case

From figure 5.2(a) and 5.2(b), it can be seen that;

• The percentage fundamental peak value of measured and simulated currentsdiffer only by 3%.

35


• Due to first order rotor slot harmonics, the high frequency currents inducedin stator winding at 636 Hz and 736 Hz, given by equation 2.6, are also ingood agreement in both cases.

From the above discussion, the simulation results are validated against the mea-surement results and are in acceptable range. Moreover, further analysis will be ofrelative comparisons between different rotor prototypes, hence absolute values areof relatively minor interest.

5.2 Standard skewed verses standard unskewed rotormotors

5.2.1 Starting performanceThe measured torque-speed and current-speed curves for the skewed and unskewedrotor motors are shown in figure 5.3(a) and 5.3(b), respectively. The starting per-formance parameters are summarized in table 5.3.

0 250 500 750 1000 1250 15000

50

100

150

200

250

300

350

400

Speed [rpm]

Tor

que

[Nm

]

Standard unskewedStandard skewed


0 250 500 750 1000 1250 15000

50

100

150

200

250

Speed [rpm]

Curr

ent

[A]

Standard skewedStandard unskewed


Figure 5.3: Measured torque-speed and current-speed curves for standard skewedand standard unskewed rotor motors

Table 5.3: Measurement results of starting performance for standard skewed andstandard unskewed rotor motors

Quantity Standard skewed Standard unskewedStarting torque [Nm] 273 286Maximum torque [Nm] 283 320Starting current(RMS) [A] 208 216

36

5.2. STANDARD SKEWED VERSES STANDARD UNSKEWED ROTOR MOTORS

It can be seen that;

• The starting torque for unskewed rotor motor is 5% higher compared to skewedrotor motor.

• The pull-out torque for unskewed rotor motor is 13% higher compared toskewed rotor motor.

• The starting current for unskewed rotor motor is 4% higher compared toskewed rotor motor.

The equivalent circuit parameters obtained from the stator resistance measure-ment and locked rotor test for the two motors is shown in table 5.4. These are usefulin understanding the difference in the starting performance characteristics for thetwo motors.

Table 5.4: Equivalent circuit parameters for standard skewed and standardunskewed rotor motors

Quantity Standard Standardskewed unskewed

Rtot(Y-connected) [ohm] 0.614 0.542Xtot(Y-connected) [ohm] 1.602 1.511Rs(Y-connected) [ohm] 0.217 0.217R′

r(Y-connected, with skin effect) [ohm] 0.397 0.325Xsl(Y-connected) [ohm] 0.534 0.534X ′

rl(Y-connected) [ohm] 1.068 0.967

One reason for large reduction of the pull-out torque for skewed rotor motor isthe 22% higher locked rotor resistance, compared to unskewed rotor motor. Duringthe locked rotor test, the inter-bar currents create additional rotor losses in skewedrotor motor, seen as an increased rotor resistance. The harmonic inter-bar currentsproduce asynchronous torques with large magnitudes well above their synchronousspeeds. These braking torques reduces the rotor accelerating torque and as a result,the pull-out torque of the motor reduces [7].

The total leakage reactance for unskewed rotor motor is 6% lower compared toskewed rotor motor. The total leakage reactance generally comprises of primary,secondary, differential and skew leakage reactance [1]. The absence of skew leakagereactance for unskewed rotor results in the reduction in total leakage reactance ofmotor compared to skewed rotor motor and the starting current and resulting start-ing torque get increased for unskewed rotor motor. An illustration of the equivalentcircuit parameter calculation is attached in Appendix B.

37


Although the starting torque and pull-out torque can be determined from figure5.3(a), the impact of synchronous and asynchronous torques cannot be observedbecause of low sampling frequency of the measuring instrument. Using differentinstrument (capable of recording torque curves at high sampling frequency), thevariation of torque with respect to time was recorded for the two motors (by vary-ing the speed between zero and rated speed range) and the results are shown infigure 5.4(a) and 5.4(b).

As seen from the filtered curves for two cases in figure 5.4(c), the starting torqueand pull-out torque are in agreement with the values obtained from initial measuredcurves at low sampling frequency. However, the impact of braking torques is clearlyseen in this figure. In case of skewed rotor motor, the synchronous torques at 150rpm, 214 rpm and the asynchronous torques due to 5th and 19th harmonic areeffectively suppressed. As a result, the motor should have much lower noise levelespecially at low speed range. For unskewed rotor motor, the asynchronous torquesare very large due to presence of first order stator slot harmonics, this should resultin a higher noise level at low speed range. For this particular case, the impact ofsynchronous torques for two motors is hard to observe because of the presence ofasynchronous torque due to 7th harmonic at the same speed.

5.2.2 Losses at rated operation

The loss components for skewed and unskewed rotor motors are shown in table 5.5.It can be seen that;

• The fundamental stator I2R losses for unskewed rotor motor are 2.6% lowercompared to skewed rotor motor.

• The fundamental rotor I2R losses for unskewed rotor motor are 7.2% lowercompared to skewed rotor motor. The high rotor I2R losses for skewed rotormotor could be due to presence of inter-bar currents, creating additional lossesin motor(the affect of which is seen as increased rotor resistance as shown intable 5.4).

• The iron losses for unskewed rotor motor are 5.2% lower than skewed rotormotor. One reason for this could be the reduction of leakage reactance forunskewed rotor motor compared to skewed rotor case. Another reason for thisdifference could be that the iron losses can vary depending upon the propertiesof shaft that depends on rotor frequency, temperature and load. The magneticproperty of lamination sheets around the punched slots can also effect the ironlosses[1].

• The stray losses have increased by 8.8% for unskewed rotor motor. Althougha reduction in overall stray losses for unskewed rotor motor was expectedaccording to [11].

38

5.2. STANDARD SKEWED VERSES STANDARD UNSKEWED ROTOR MOTORS

0 5 10 15 20 25−100

0

100

200

300

400

500

Time [s]

Tor

que

[Nm

]

Measured(Standard Skewed)Filtered(fc = 35 Hz, fs = 5 kHz)

(a) Standard skewed rotor

0 5 10 15 20 25−100

0

100

200

300

400

500

Time [s]

Tor

que

[Nm

]

Measured(Standard Unskewed)Filtered(fc = 35 Hz, fs = 5 kHz)

(b) Standard unskewed rotor

0 5 10 15 20 25−100

0

100

200

300

400

500

Time [s]

Tor

que

[Nm

]

Standard unskewed

Standard skewed

− Synchronous torques at 150rpm and 214rpm due tostator & rotor slot number, respectively.− Asynchronous torques due to 7th space harmonic and19th stator slot harmonic.

(c) Filtered curve comparision

Figure 5.4: Measured variation of torque with time at high sampling frequency forskewed and unskewed rotor motors

39


Table 5.5: Measured losses for standard skewed and standard unskewed rotor motorsQuantity Standard skewed Standard unskewed

Fundamental I2R stator losses [W ] 552 538Fundamental I2R rotor losses [W ] 297 277Total iron losses [W ] 343 325Friction losses [W ] 76 69Stray losses [W ] 284 309Total losses [W ] 1552 1518Output power [kW ] 15 15Efficiency [%] 90.7 90.9Cosϕ 0.82 0.83Winding temperature [0C] 74.5 73.7Current [A] 29.0 28.7Torque [Nm] 97.3 97.2

• The overall efficiency measured for both motors is the same with negligibledifference between the values.


The delta circulating current in stator windings was measured for the skewed andunskewed rotor motors and is shown in figure 5.5(a). The FFT was performed inorder to investigate the high frequency currents in stator windings and results areshown in figure 5.5(b). For comparison purpose, the results of only one phase wind-ings are presented here. Due to non-identical stator windings in reality, the currentsmeasured in each phase winding varied slightly but the difference was negligible.

From figure 5.5(a), it can be seen that the winding current waveform for skewedrotor motor is more close to sinusoidal compared to winding current waveform forunskewed rotor motor. The FFT analysis shown in figure 5.5(b) confirms the pres-ence of high frequency current harmonics with 7% of the fundamental current at636 Hz and 13% of the fundamental current at 736 Hz. These are due to first orderrotor slot harmonics for unskewed rotor motor.

For skewed rotor motor, apart from high frequency current which are multipleof three, there are negligible high frequency components. The rotor skewing hadeffectively suppressed the first order rotor slot harmonics. Since the motor withunskewed rotor was found to be lot noisier than skewed motor, so it would beinteresting to investigate whether these harmonics are the source for noise in motor.

40

5.3. ASYMMETRICAL VERSES SYMMETRICAL ROTOR MOTORS

0 0.005 0.01 0.015 0.02 0.025 0.03 0.035 0.04−40

−30

−20

−10

0

10

20

30

40

Time[s]

Win

din

gcu

rren

t,Iab[A

]

Standarad skewedStandard unskewed

(a)

0 100 200 300 400 500 600 700 800 900 1 0000

2.5

5

7.5

10

12.5

15

Frequency [Hz]

%am

plitu

de

ofw

indin

gcu

rren

t,Iab

Standard unkewedStandard skewed

150 Hz636 Hz

736 Hz

(b)

Figure 5.5: Measured results for high frequency stator winding current for standardskewed and standard unskewed rotor motors

5.3 Asymmetrical verses symmetrical rotor motors

5.3.1 Starting performanceThe torque-speed curves and the current-speed curves for the three rotor designsare shown in figure 5.6.

0 250 500 750 1000 1250 15000

50

100

150

200

250

300

350

400

Speed [rpm]

Tor

que

[Nm

]

Standard unskewedProgressive sinusoidalStandard skewed


0 250 500 750 1000 1250 15000

25

50

75

100

125

150

175

200

225

250

Speed [rpm]

Curr

ent

[A]

Standard skewedStandard unskewedProgressive sinusoidal


Figure 5.6: Measured torque-speed and current-speed curves for three rotor designs

The starting performance characteristics are summarized in table 5.6. It can beobserved that;

• The starting torque for asymmetrical rotor motor is 7.7% higher compared tostandard skewed motor.

41


Table 5.6: Measurement results of the starting performance for three rotor designsSymmetrical Asymmetrical

Quantity Standard Standard Progressiveskewed unskewed sinusoidal

Starting Torque [Nm] 273 286 294Pull-out Torque [Nm] 283 320 322Starting Current [A] 208 216 219

Table 5.7: Equivalent circuit parameters for standard skewed, standard unskewedand asymmetrical rotor motors

Symmetrical AsymmetricalQuantity Standard Standard Progressive

skewed unskewed sinusoidalRtot(Y-connected) [ohm] 0.614 0.542 0.555Xtot(Y-connected) [ohm] 1.602 1.511 1.494Rs(Y-connected) [ohm] 0.217 0.217 0.217R′

r(Y-connected, with skin effect) [ohm] 0.397 0.325 0.338Xsl(Y-connected) [ohm] 0.534 0.534 0.534X ′

rl(Y-connected) [ohm] 1.068 0.977 0.960

• The pull-out torque for asymmetrical rotor motor is 13.8% higher comparedto standard skewed motor.

• The starting current for asymmetrical rotor motor is 5% higher compared tostandard skewed motor.

• The starting performance characteristics for asymmetrical rotor motor arevery similar to the standard unskewed motor.

The equivalent circuit parameters for the three motors are tabulated in table5.7. From equivalent circuit parameters, it can be seen that the total leakage reac-tance and rotor resistance is similar for standard unskewed and asymmetrical rotormotors and it is 7% lower compared to standard skewed motor. As explained ear-lier, for both unskewed motors, the absence of rotor skew leakage and the harmonicinter-bar currents results in a higher pull-out torque compared to skewed rotor mo-tor. A higher pull-out torque can be advantageous specially if the motor is to beused in traction applications.

42


0 5 10 15 20 25−100

0

100

200

300

400

500

Time [s]

Tor

que

[Nm

]

Standard unskewedProgressive sinusoidalStandard skewed

− Synchronous torques at 150rpm and 214rpm due tostator & rotor slot number, respectively.− Asynchronous torques due to 7th space harmonic and19th stator slot harmonic.

Figure 5.7: Torque variation with time for motor with different rotor prototypes

To investigate the impact of synchronous and asynchronous torques on the start-ing performance, the torque variation from rated to zero speed range was recordedat a high sampling frequency for the three motors. The resulting curves are shownin figure 5.7.

From figure 5.7, it can be seen that the braking torques at low speed have smallerimpact for asymmetrical rotor motor compared to standard unskewed motor. It isalso evident that with proper modulation of rotor slots (by choosing the right mod-ulation coefficients) results in the suppression of synchronous torques at low speed.The measurement results also validates the 2D FEM simulated starting performanceshown in figure 4.1(e). The motor starting characteristics with asymmetrical rotorare much improved compared to standard unskewed rotor motor.

5.3.2 Losses at rated operation

The losses at rated operation for the three motors are listed in table 5.8.From table 5.8, it can be observed that;

• The fundamental I2R stator losses for asymmetrical rotor motor are 2.6%higher compared to standard unskewed motor whereas, they are same com-pared to standard skewed motor.

• The fundamental I2R rotor losses for asymmetrical rotor motor are 9% highercompared to standard unskewed motor whereas, the difference compared tostandard skewed motor is negligible.

43


Table 5.8: Summary of measured losses for motor with three rotor designsSymmetrical Asymmetrical

Quantity Standard Standard Progressiveskewed unskewed sinusoidal

Fundamental I2R stator losses [W ] 552 538 552Fundamental I2R rotor losses [W ] 297 277 303Total iron losses [W ] 343 325 308Stray losses [W ] 284 309 293Friction losses [W ] 76 69 88Total losses [W ] 1552 1518 1544Output power [kW ] 15 15 15Efficiency [%] 90.7 90.9 90.7Cosϕ 0.82 0.83 0.83Winding temperature [0C] 74.5 73.7 77.7Current [A] 29.0 28.7 28.9Torque [Nm] 97.3 97.2 97.4

• The iron losses for asymmetrical rotor motor are 5.5% and 11.3% lower com-pared to standard unskewed and standard skewed motors, respectively.

• The friction losses differ slightly for three motors. This could be due to vari-ation in mounting process of rotors in the same stator.

• The stray losses for asymmetrical rotor motor are 5% lower compared tostandard skewed motor. However, the increasing trend of stray losses forunskewed rotors compared to skewed rotor motors is very surprising and needsfurther investigation. Since an increase in efficiency was expected with theremoval of rotor skew that results in elimination of inter-bar current lossesfrom stray losses category [3], [7].

• The total losses for asymmetrical rotor motor are almost similar to standardskewed motor and are 1.7% higher compared to standard unskewed motor.It must be noted here that the asymmetrical rotor had casting issues duringmanufacturing. So a deteriorated performance during the tests was expectedfor asymmetrical rotor prototype.

5.3.3 High frequency stator(delta connected) winding currentsThe stator winding currents for all rotor cases were measured. For comparisonpurpose, winding current in one phase for each rotor designs is shown in figure 5.8.

44


0 0.005 0.01 0.015 0.02 0.025 0.03 0.035 0.04−40

−20

0

20

40

Time[s]

Win

din

gcu

rren

t,Iab[A

]

Standarad skewedStandard unskewedProgressive sinusoidal

Figure 5.8: Measured winding current Iab at rated operation for three rotor designs

As seen from figure 5.8, the winding current waveform for standard skewed motoris very close to sinusoidal, whereas, the winding current waveforms for standardunskewed and asymmetrical rotor motor are distorted. In order to investigate thehigh frequency stator winding currents, the FFT analysis was carried out on thewinding current Iab for each motor and the results are shown in figure 5.9. It canbe seen that;

• The magnitude of frequency component at 150 Hz due to 3rd order harmonic,is very similar for all the three motors.

• Apart from the 3rd order harmonic, the standard skewed motor has negli-gible high frequency components in stator winding. Skewing has effectivelysuppressed the 1st order rotor slot harmonics.

• The frequency components at 636 Hz and 736 Hz have highest magnitudefor standard unskewed motor. These high frequency components are due to1st order rotor slot harmonics, as given by equation 2.6.

• The high frequency components at 636 Hz and 736 Hz for asymmetricalrotor motor have lower magnitude compared to standard unskewed rotor mo-tor. However, additional high frequency components at 734 Hz and 834 Hzappeared for asymmetrical rotor motor which further distorts the resultantwinding current waveform. According to equation 2.6, these high frequencyharmonics should appear for a 32 slot rotor. This shows that for asymmetricalrotors, the dominant rotor slot harmonic is not always due to the apparent slotnumbers, but it is possible to have an inherent slot number with a significantlyhigh slot permeance value [10].

• The additional high frequency components in winding currents for asymmet-rical rotor designs may result in further increase in the stator winding lossesin motor.

45


146 148 150 152 1540

2.5

5

7.5

10

12.5

15

Frequency [Hz]

%of

fundam

enta

lw

indin

gcu

rren

t,Iab

Progressive SinusoidalStandard UnkewedStandard Skewed

(a)

632 634 636 638 6400

2.5

5

7.5

10

12.5

15

Frequency [Hz]

%of

fundam

enta

lw

indin

gcu

rren

t,Iab


(b)

732 734 736 738 7400

2.5

5

7.5

10

12.5

15

Frequency [Hz]

%of

fundam

enta

lw

indin

gcu

rren

t,Iab


(c)

832 834 836 838 8400

2.5

5

7.5

10

12.5

15

Frequency [Hz]

%of

fundam

enta

lw

indin

gcu

rren

t,Iab


(d)

Figure 5.9: High frequency components in measured winding current Iab at ratedoperation for three rotor designs

46

Chapter 6

Conclusions and Future Work

6.1 Conclusions

Some important conclusions from this work are listed as the following:

• The 2D FEM simulation results from FCSmek showed good agreement withthe measured results for the three prototypes.

• For asymmetrical rotor machines, the synchronous torques at low speed wereeffectively suppressed with new modulation function and coefficients.

• Due to the absence of skew leakage and negligible inter-bar currents, a 13% higher pull-out torque for asymmetrical rotor machine was achieved withminimal increase in starting current compared to standard skewed machine.

• The high frequency components in stator winding currents are negligible forskewed rotors compared to unskewed rotors. Skewing effectively suppressesthese high frequency components due to the rotor slot harmonics.

• For unskewed asymmetrical rotors machines, apart from the high frequencycomponents due to apparent rotor slot number, additional high frequencycomponents, due to inherent slot number, are also introduced in stator windingcurrents. This results in a small increase in the total stator winding losses.

• A summary of measured stray losses for three rotor prototypes are presentedin table 6.1. The total losses more or less re-distributed in the unskewed rotormachines compared to skewed rotor machine. The measured stray losses forunskewed rotor machines have increased contradicting the hypothesis. Theasymmetrical rotor machine achieved the same efficiency compared to stan-dard skewed rotor machine.

47

CHAPTER 6. CONCLUSIONS AND FUTURE WORK

Table 6.1: Stray losses for three rotor designsOutput Total Stray Stray Estimated+ Inter-barpower losses losses losses inter-bar current Efficiency

Cases (% of current losses(% oftotal losses total

[W] [W] [W] losses) [W] losses) [%]Standard 15000 1552 284 18.3% 85.2 5.5% 90.7%skewed

Standard 15000 1518 309∗ 20.4% — — 90.9%unskewed

Progressive 15000 1544 293∗ 19.0% — — 90.7%sinusoidal

+The inter-bar currents are considered 30% of stray loss, as shown in figure 1.5(a).∗A reduction of 30% in stray losses and 5.5% in total losses was expected for unskewedcases, compared to skewed case [7].

6.2 Future workThe limitations and future work are suggested as follows:

• This study was limited to 36/28 slot combination motor. The 2D FEM toolFCSmek can be used to investigate other slot combinations as well, as it servesas a economical, time saving and good calculation tool to study these concepts.

• Also only one prototype of each kind was available for testing which makes itdifficult to draw any major conclusions out of the results. However, the 2DFEM tool FCSmek can be used to further improve the studied asymmetricaldesigns. One possible improvement work could be to suppress the synchronoustorques at 214 rpm for the progressive sinusoidal rotor design.

• The aluminium leakage problem for asymmetrical rotor prototypes need to beaddressed properly. The laser cut method needs to be improved in order toproduce lamination sheets with smooth surface and edges. This could help inproducing better asymmetrical rotor prototypes.

• It will be interesting to measure the audible noise of the unskewed machines,since these machines sounded noisier during tests compared to standard skewedrotor machine.

• Apart from direct on-line start, the asymmetrical rotor could be test in inverterfed drives to see the performance and limitations for these rotor designs.

48

6.2. FUTURE WORK

• One important observation is that to capture inter-bar current losses whichare estimated to be 5.5% of the total losses requires more accurate methods ofmeasurements than the existing where sufficient repeatability can be achieved;alternatively one can rely on statistical data obtained from measurements onseveral number of motors.

49

Bibliography

[1] C. Sadarangani, Electrical Machines-Design and Analysis of Induction and Per-manent Magnet Motors, KTH, Book Chapter 4, 1st Edition, 2006.

[2] Anibal T. de Almeida, Fernando J.T.E. Ferreira and Joao A. C. Fong, Standardsfor Super-Premium Efficiency Class for Electric Motors, IEEE/IAS Industrialand Commercial Power System Technical Conference, p 8 pp, 2009.

[3] R. Chitroju, Improved Performance Characteristics of Induction Machines withNon-skewed Asymmetrical Rotor Slots, Licentiate Thesis, Royal Institute ofTechnology, Stockholm, Sweden, 2009.

[4] IEC-60034-30, Efficiency classes of single-speed, three-phase, cage-induction mo-tors, Edition 1.0, 2008-10.

[5] ABB LV Motors, Efficiency Regulations for Low Voltage Motors - ABB,http://http://inside.abb.com/search.aspx?q=ie4$\%$20efficiency$\%$20standard, last accessed 2012-04-12.

[6] Electric Machines and Drives, Book, Chapter 2, KTH 2010.

[7] A. Stening, On Inter-bar Currents in Induction Motors with Cast Aluminiumand Cast Copper Rotors, Licentiate thesis, Royal Institute of Technology, Stock-holm, Sweden, 2010.

[8] S.L.Nau, The influence of the skewed rotor slots on the magnetic noise of three-phase induction motors, In Proc. of Int. Conf. on Electrical Machines and Drives,pp. 396-399, Cambridge, 1997.

[9] H. Nishizawa, K. Itomi, S. Hibino, and F. Ishibashi, Study on reliable reductionof stray load losses in three-phase induction motor for mass production, IEEETransactions on Energy Conversion, vol. EC-2, pp. 489-495, 1987.

[10] A. Stening, C. Sadrangani, Reduction of Synchronous Torques in Induction Ma-chines using Asymmetrical Rotor Slots, Accepted for publication at the ICEMS,Sapporo, Oct.2012.

51

BIBLIOGRAPHY

[11] A. Stening, Performance Analysis of Asymmetrical Rotor InductionMotors https://eeweb01.ee.kth.se/upload/publications/reports/2011/IR-EE-EME_2011_012.pdf, last accessed 2012-05-12.

[12] A. Hagen, A. Binder, M. Aoulkadi, T. Knopik, and K. Bradley, Comparison ofmeasured and analytically calculated stray load losses in standard cage inductionmachines, 18th International Conference on Electrical Machines, pp. 1-6, 2008.

[13] Energimyndigheten, Improving the Efficiency of squirrel cage induction mo-tors: technical and economical consideration, http://energimyndigheten.se/PageFiles/18106/Elmotorer20Beaktande20803970.pdf, last accessed 2012-04-10.

[14] IEC-60034-2-1, Standard methods for determining losses and efficiency fromtests, Edition 1.0, 2007-09.

[15] Wenping Cao, Comparison of IEEE 112 and new IEC Standard 60034-2-1,IEEE Transactions on Energy Conversion, v 24, n 3, p 802-8, Sept. 2009.

[16] Agamloh, Emmanuel B., A comparison of direct and indirect measurement ofinduction motor efficiency, IEEE International Electric Machines and DrivesConference, IEMDC ’09, p 36-42, 2009.

[17] Xiaodong Liang, Alta, Yilmaz Luy, Harmonic Analysis for Induction Motors,Electrical and Computer Engineering, 2006. CCECE ’06. Canadian Conferenceon Digital Object Identifier: 10.1109/CCECE.2006.277368, Publication Year:2006 , Page(s): 172-177.

[18] McClay, C.I.; van der Toorn, G.T., A comparison of time-stepped finite-elementtechniques for the calculation of losses in cage induction motors, Electrical Ma-chines and Drives, 1999. Ninth International Conference on (Conf. Publ. No.468) , pp. 35-39, 1999.

[19] Meeker, D., Finite Element Method Magnetics, Version 4.2, User Manual, 2007https://www.femm.info, last accessed 2012-04-25.

[20] LUA scripts, Programming in LUA https://http://www.lua.org/pil/, lastaccessed 2012-05-04.

52

Appendix A

Rotor slot dimensions for standard rotorprototype

Typical double cage rotor slot with its dimensions shown in figure A.1.

Figure A.1: Rotor slot dimensions

Table A.1: Slot dimensions and pitch angles for standard rotorSlot BSO2 BSYD2 BSOD2 BSID2 Slot angle# [mm] [mm] [mm] [mm] [degrees]1 6.5 2.0 6.3 2.3 6.4282 6.5 2.0 6.3 2.3 19.2783 6.5 2.0 6.3 2.3 32.1284 6.5 2.0 6.3 2.3 44.9785 6.5 2.0 6.3 2.3 57.8286 6.5 2.0 6.3 2.3 70.6787 6.5 2.0 6.3 2.3 83.528

53

Appendix B

Calculation of equivalent circuitparameters from locked rotor test

Standard unskewed rotor case:

A single phase circuit during locked rotor test is shown in figure B.1. The ironlosses and magnetizing reactance are negligible and can be ignored for simplicity.

Xsl Xrl

Rr´

Rs

Is Ir'=Is

Us

+

-

Figure B.1: Single phase equivalent circuit for locked rotor condition

The initial values from locked rotor test are as follows:

• UL−L = 79.43 V

• Is(L) = 28.73 A

• cosϕ = 0.34

• Stator resistance, Rs = 0.217 ohm (Y-connected, obtained from resistancemeasurement test @740C)

55

APPENDIX B. CALCULATION OF EQUIVALENT CIRCUIT PARAMETERS FROMLOCKED ROTOR TEST

• Stator leakage reactance, Xs = 0.534 ohm (Y-connected, obtained from FEMcalculation tool)

Now, the apparent, active and reactive powers are given as;

S =√

3.UL−L.Is(L) = 3948 V A (B.1)

P = Scosϕ = 1342 Watts (B.2)

Q = Ssinϕ = 3712 V AR (B.3)

Neglecting the magnetizing reactance and iron losses, the total leakage reactance isgiven in this case as;

Xtot = Q

3.I2s

= 1.499 ohm (B.4)

The rotor leakage seen from stator side is thus given as;

X ′r = Xtot − Xs = 0.967 ohm (B.5)

The total resistance is found as;

Rtot = P

3.I2s

= 0.542 ohm (B.6)

Thus the rotor resistance, including skin effect as seen from stator side is;

R′r = Rtot − Rs = 0.325 ohm (B.7)

56

performance analysis of unskewed asymmetrical rotor for lv...

Documents