Axial-Flux Permanent-Magnet Brushless DC Traction Motor for ...

International Review of Electrical Engineering (I.R.E.E.), Vol. 6, N. 2 March-April 2011

Axial-Flux Permanent-Magnet Brushless DC Traction Motor for Direct Drive of Electric Vehicle N. A. Rahim, W. P. Hew, A. Mahmoudi Abstract This paper presents the design of an inside-out axial-flux permanent-magnet

brushless dc motor for direct traction drive in an electric vehicle. The prototype motor is a double-sided axial-flux permanent-magnet motor with non-slotted stator. The preliminary design had 16 rotor poles, for high torque density and stable rotation at low speed. The design was simulated via Finite Element Method Magnetics (FEMM) Software, for obtainment of design parameters. The motor was fabricated and tested in an in-wheel test-bed. There exist close agreements between the simulated and experimental results. Copyright © 2011 Praise Worthy Prize S.r.l. - All rights reserved.

Keywords: Axial-Flux Permanent-Magnet Motor, Electric Vehicle, Finite Element Analysis

Ki Bg fe fm A

Nomenclature Frm fro fst fl fr M g Cd S v v0 min

r Pmin m

a Paccel Pout m e(t) i(t) T Kp fe(t) fi(t) Epk Ipk Irms P Kw Nph

Vehicle driving resistance Rolling-resistance force Climbing-resistance force Aerodynamic-resistance force Rolling-resistance coefficient Vehicle mass [kg] Gravity acceleration [m/s2] Vehicle movement angle Air density Air-resistance coefficient Frontal projected area Vehicle speed Headwind speed Minimum required torque [N m] Position vector Minimum required power [W] Rotor angular speed [rad/s] Vehicle acceleration [m/s2] Power required to accelerate [W] Rated power Motor efficiency Number of phases Phase-air-gap EMF [V] Phase current [A] Period of one EMF cycle [s] Electrical power waveform factor Normalized EMF waveforms Normalized current waveforms Peak value of phase-air-gap EMF Phase current peak value Phase current rms value [V] Number of motor pole pairs Winding distribution factor Number of winding turns per phase

Do Di K Ke m1 Ar As KL Dtot Ltot den

Wcu Dave Kcu Js Lss Le Ls Lcs p

Bcs Lr Lcr Lpm Bcr Bu g µr Br

Manuscript received and revised March 2011, accepted April 2011

Current waveform factor Air-gap flux density [Wb/m2] Electrical frequency [Hz] Mechanical frequency [Hz] Electrical loading total [A] Diameter ratio Machine stator outer diameter [m] Machine stator inner diameter [m] Electrical loading ratio EMF factor Number of phases of each stator Rotor electrical loading [A] Stator electrical loading [A] Aspect ratio coefficient Machine outer diameter total [m] Machine axial length total [m] Torque density [N m/cm3] End-winding protrusion from iron stack [m] Machine stator average diameter [m] Copper fill factor Current density [A/m2] Stator slot depth [m] Effective axial length of motor [m] Stator axial length [m] Stator-core axial length [m] Average air-gap flux density to its peak value ratio Stator-core flux density [T] Rotor axial length [m] Rotor-core axial length [m] Permanent-magnet length [m] Rotor-disc flux density [T] Flux density on permanent-magnet surface [T] Air-gap length [m] Recoil relative permeability of magnet Permanent-magnet residual-flux density [T] Copyright © 2011 Praise Worthy Prize S.r.l. - All rights reserved

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N. A. Rahim, W. P. Hew, A. Mahmoudi

Kd Kc Kf Bgpk Pe Pm

II.

Leakage-flux factor Carter factor Peak value corrected factor of air-gap flux density Peak value of air-gap flux density [Wb/m2] Electrical power [W] Mechanical power [W]

I.

Design Procedure

II.1.

Vehicle Dynamics

A simple vehicle dynamics model to evaluate vehicle performance is presented. A simplified vehicle driving resistance or road load (Frm) consists of rolling resistance force (fro), climbing resistance force (fst), and aerodynamic drag force (fl):

Introduction

Frm

Protection of natural environments sparked interest in electric vehicle (EV), which is non-polluting. EV was first introduced in 1870; it had light electric motor and very heavy storage batteries. Battery, electric motor, motor drive circuit, and transmission gears make up EV power system. Range of EV driving speeds was limited. Researchers and designers keep attempting more-efficient and morereliable EV power systems. Improvements to each subsystem have increased overall efficiency and driving range [1]-[6]. Attempts at finding the most suitable EV motor are keen pursuits of researchers and engineers throughout the world. Permanent-magnet motors already developed for electric vehicles fulfill requirements for, e.g., high power-density, high efficiency, high starting torque, and high cruising speed. Low cost, high speed, low torqueripple, high reliability, established manufacturing technology that includes converter, and absence of position sensors make induction motor the preferred drive system [7]. Compactness, low weight, and high efficiency of permanent-magnet brushless DC motors are suitable options for EV propulsion [8]-[11]. Motors designed for EV drive can be classified as direct drive [12] or indirect drive [13]. Direct-drive motor is wheelmounted. Mechanical deferential and transmission gears, including the associated energy losses, are thus eliminated. Not only is efficiency improved, but vehicle weight is reduced. Slotless AFPM motors have over conventional radial-flux motors advantages such as high torque-toweight ratio, high efficiency, adjustable air gap, balanced motor-stator attractive forces, and better heat-removal [14]-[17]. This paper presents the design of and experimental work on slotless AFPM motor for EV. The motor is designed for placement inside the wheel of a motorcycle. Its specifications are according to typical vehicle dynamics. Sizing equations of TORUS AFPM machines are derived via generalized sizing equation, to calculate motors power-production potential. The sizing equation is used for optimum machine design. Finite-element analysis is then performed in field analysis of the proposed motor topology. Finally, a prototype motor is fabricated, and experiments performed, for information on possible current driving patterns.

f ro

f st

fl

(1)

Rolling resistance (fro) is caused by on-road tire deformation: f ro f r M g (2) where fr, M, and g are rolling resistance coefficient, vehicle mass, and gravity acceleration, respectively. Climbing resistance (fst with positive operational sign) and downward force (fst with negative operational sign) are given by: f st

M g Sin

(3)

where, is angle of vehicle movement relative to horizon. Aerodynamic drag force (fl) is air viscous resistance on vehicle: 1 2

fl

Cd S v v0

2

(4)

where, is air density, Cd is air-resistance coefficient, S is frontal projected area, v is vehicle speed, and v0 is headwind speed. Acting as propulsion, driving force is applied to wheels to overcome driving resistance. Driving force lower than driving resistance does not make vehicle roll. In angular movement, minimum required torque for vehicle propulsion is: r Frm

min

(5)

where, r is position vector. Minimum power required is thus: Pmin

(6)

min

where, m is rotor angular speed. Acceleration is important to vehicle movement; energy losses caused by it (a) must factor in calculations. Power required to accelerate EV is thus: Paccel

M va

(7)

Power at wheels is:

Pout

Copyright © 2011 Praise Worthy Prize S.r.l. - All rights reserved

Paccel

Pmin

(8)

International Review of Electrical Engineering, Vol. 6, N. 2

761


e(t) is phase air-gap EMF, i(t) is phase current, is machine efficiency, m is number of machine phases, and T period of one EMF cycle. Epk and Ipk are peaks of phase air-gap EMF and of current, respectively. Kp is electrical power waveform factor, defined as: Kp Fig. 1. Proposed Driving Cycles for Electric-Vehicle Design

e t i t dt

mK p E pk I pk

1 T

T 0

fe t

fi t dt (10)

I pk

1 1 T

T 0

(11)

2

i t I pk

dt

K e N ph Bg

f 1 p

2

Do2

(12)

Ke is EMF factor incorporating winding distribution factor (Kw) and per-unit portion of air-gap area-total spanned by machines salient poles (if any); Nph is number of turns per phase; Bg is flux density in air gap; f is converter frequency; P is machine pole pairs; is AFPM diameter ratio Di /Do; Do is diameter of machine outer surface; Di is diameter of machine inner surface. Equation (9)s peak phase current is:

Sizing Equation

I pk

A Ki

1

Do 2 2m1 N ph

(13)

where, m1 is number of phases of each stator, and A is electrical loading. Other authors have provided a general-purpose sizing equation for AFPM machines; it takes the following form: Pout

Main dimensions of each electrical machine are determined via electrical-machine-output power equation. Assuming negligible leakage inductance and resistance, rated power is expressed as [18]:

0

dt

I rms

E pk

TABLE II DESIGN RESTRICTIONS AND REQUIREMENTS Dimensional Constraints Stator Outer Diameter 460 mm Total Axial-Length 80 mm Air-Gap 1 mm Limits on Power Systems Permanent Remanence 1.3 T Rated Line-to-Line Voltage (rms) 70V Input Phase Current (rms) 30A Requirements Maximum Torque 36.5 N.m Output Power 1.8 kw Motor Efficiency >90%

Pout

E pk I pk

where, Irms is phase-current rms value. Table III lists typical waveforms and their corresponding powerwaveform factor (Kp) and current-waveform factor (Ki) [14]. Peak value of phase-air-gap EMF for equation (8)s AFPM motor is:

An optimum design would be maximized torque density while desired efficiency is maintained within design restrictions and requirements (see Table II).

T

e t i t

Ki

TABLE I PARAMETERS USED IN THIS STUDY Vehicle Specification Weight of Vehicle 80 kg Weight of Passengers 70 kg Wheel Radius (Rd) 0.30 m Tire Set 3 units Drive System Front drive Frontal Area (S) 0.4 m2 Air Resistance Coefficient (Cd) 0.35 Tire Resistance Coefficient (fr ) 2.5×10-3 Air Density ( ) 1.22 kg/m3 Maximum Speed (vmax) 60 km/h

m T

T 0

where fe(t)=e(t)/Epk and fi(t)=i(t)/Ipk are expressions for normalized EMF and current waveforms. For effect of current waveform, current waveform factor (Ki) is defined and presented:

To design EV motor propulsion, vehicle dynamics should first be determined. Fig. 1 is EV cruising scenario, which includes an EVs typical-trip elements such as increasing speed, constant speed, and braking action. Power needed by the vehicle is calculated from the proposed driving cycle in Fig. 1, together with Equations (1) to (8). Table I lists the parameters used in the study.

II.2.

1 T

1 m K e Ki K p K L Bg A 1 K m1 2 f 1 P

2

1 2

Do2

(14)

Le

m1 is number of phases of each stator; Le is effective axial length of the motor; K is electrical loading ratio on rotor and stator; KL is aspect ratio coefficient pertinent to a specific machine structure, with considerations for effects of losses, temperature rise, and the designs

(9)



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efficiency requirements. Also, machine torque density for volume total is defined as: Pout

den m

4

Le

Ls

3

Pout m f K e K p Ki ABg 1 2 m1 p

2

Lcs

Model

Do

TABLE III TYPICAL PROTOTYPE WAVEFORMS e(t) i(t)

(17)

Ki 2

0.5Cos

Sinusoidal

2

0.5

Rectangular

1

1

Trapezoidal

1.134

0.777

Triangular

3

Wcu

Di2 2

p Do

1

Lcr

(21)

(22)

L pm

r Bg

Br

Kf Kd

Kc g

(24)

Bg

where µr is magnets recoil relative permeability, Br is permanent-magnet material residual-flux density, Kd is leakage flux factor, Kc is Carter factor, Kf =Bgpk/Bg is peak value corrected factor of air-gap flux density in radial direction of AFPM motor. These factors can be obtained from FEM analysis [19]. In AFPM motors, air-gap flux density and diameter ratio are the two important design parameters having significant effect on motor characteristics. To optimize motor performance, diameter ratio and air-gap flux density must be chosen carefully. The optimum design should maximize power density while maintaining desired efficiency within design restrictions (Table II). In design studies, diameter ratio and air-gap flux density are design parameters. Fig. 2 shows power density variation as a function of air-gap flux density, and diameter ratio of the AFPM motor.

0.333

2 ADave K cu J s

(20)

4 pBcs

Lcr

where, Wcu is protrusion of end winding from iron stack, in radial direction. For back-to-back wrapped winding, protrusions exist towards machine axis as well as towards the outsides, and can be calculated as: Di

2 Lss

where Bcr is flux density in rotor disc core, and Bu is attainable flux density on permanent-magnet surface. Permanent-magnet length Lpm can be calculated as:

Kp

Sinusoidal

Lcs

Lpm is permanent-magnet length; axial length of rotor core Lcr is: Bu Do 1 Lcr (23) 8 pBcr

2

2Wcu

Bg

Lr

Machine outer diameter total Dtot for the TORUS motor is given by: Dtot

(19)

where Bcs is flux density in stator core, and p is ratio of average air-gap flux density to peak air-gap flux density. Axial length of rotor Lr becomes:

(16)

1

2g

Axial length of stator core Lcs can be written as:

m is rotor angular speed, Dtot and Ltot respectively are machine outer diameter total and machine length total including stack outer diameter and end-winding protrusion from radial and axial iron stacks. The generalized sizing equation approach can easily be applied to double-sided axial-flux permanent-magnet TORUS type motor. The outer surface diameter (Do) can be written as:

Do

2 Lr

Lr is axial length of rotor, and g is air-gap length. Axial length of stator Ls can be written as:

(15)

2 Dtot Ltot

Ls

III. Simulation and Finite Element Analysis

(18)

The design was simulated via Finite Element Method Magnetics (FEMM) Software. The simulation model reached the output (2.7 kW) targeted for the electric motorcycle.

where, Dave is average diameter of the machine, Js is current density, and Kcu is copper fill factor. Note that for slotted machines, depth of stator slot is Lss=Wcu. Axial length Le of machine is given by: Copyright © 2011 Praise Worthy Prize S.r.l. - All rights reserved


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of magnet inner to outer diameter (Dave=(Di+Do)/2). Corresponding materials and circuit currents were assigned to each block of the model; see Fig. 3(a). The motors 2-D model is symmetric, so 16 magnetic poles were sliced to reduce simulation/calculation time, and the FEMM model became six magnetic pole pieces. Results from the model were calculated via LUA programming language, to obtain values for the entire motor. For simulation, input parameters needing consideration were permanent-magnet thickness, air-gap width, and magnetic properties of all active materials. Fig. 3(b) shows the magnetic flux density generated by the permanent magnets.

Fig. 2. Torque density vs. air-gap flux density and diameter ratio

The FEMM 4.0 software allows calculating in 2-D space, so the actual motor had to be modified to the flat model, in which all curvatures were developed relative to average diameter placed middle of stator core or average

(a) AFPM motor model in 2D with Fine Meshing

(b) Magnetic-flux density FEMM simulation

Figs. 3. AFPM motor simulation using FEMM

The relatively symmetrical distribution of the magnetic-flux density relative to radial symmetrical axis of the magnets indicates currents negligible influence on resultant magnetic field. Maximum flux density is higher in the stator core than in the rotor because the stator core is laminated steel that saturates at much higher values. Table IV shows the parameters and the optimized TORUS motor dimensions calculated via sizing equations. TABLE IV MOTOR DIMENSIONS Rotor Inner Diameter

130 mm

Rotor Outer Diameter

230 mm

Number of Windings

48

Number of Turns

12

Magnetic Pole

16

Magnet Thickness Magnet Arc Magnet Material Back-Iron Thickness Rated Voltage Line-to-Line (rms) Rated Phase Current (rms) Output Power

FEA was for overview of saturation levels in various parts of the machine, for comparison of flux densities obtained from FEM with sizing analysis. Table V tabulates results for the comparison, which was done at no load and for various parts of the machine. The no-load flux density plots show consistency with sizing analysis, maximum flux density of rotor and of stator almost equal. Maximum and average air-gap flux densities from FEM and from sizing analysis agree, too. Fig. 4 compares calculated back-EMF against electrical angle of the designed motor, from FEA and from the no-load experiment at 700 rpm. The experiments peak back-EMF was 90.4V, slightly less than the 95V computed value, agreeing closely with the computed waveform. Fig. 5 compares experiment torque against electrical angle variation, and predicted torque from FEA; both closely agree. At 30A rated current and 700rpm rated speed, the motor produces 37.4Nm maximum torque while the simulation showed 36.78Nm.

7 mm 18o Nd-Fe-B, N35 12 mm

TABLE V FLUX DENSITY COMPARISON OF THE DESIGNED MOTOR Rotor Air-gap stator Bcr Bmax Bave Bcs FEM 1.2 0.81 0.52 1.15 Sizing Eq. 1.1 0.8 0.5 1.1

64V 30A 1.8 kW



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Fig. 4. Back-EMF at 700 rpm

Fig. 6. Surface-mounted permanent-magnet arrangement on rotor back-iron

Fig. 5. Mechanical torque at 30A

The results show the motors ability to fulfill the electric motorcycles power requirement.

IV.

Fabrication and Experiment Works Fig. 7. Toroidal windings for a slot-less stator

The motor costs relatively low to manufacture, as there are no stator teeth. The stator lamination silicon steels are rolled; no need to wire-cut or laser-cut. Components such as rotor plate and shaft are also designed simply and also cost relatively low to manufacture. Absence of teeth makes windings difficult to assemble. The motor uses encapsulated thermal conductor epoxy; good for releasing heat, but stands up to only 80°C before its rigidity decreases.

The shaft was embedded with stator components; see Fig 8. It kept the stator from returning to a direction opposite to rotor (wheel), so the shaft had to be strong enough for the stator to hold up to the motors torque.

IV.1. Manufacturing Design challenge in manufacturing the AFPM motor is maintaining air gap between stator and rotor. Magnetic interaction between rotor magnet and stator back-iron is quite large (752 N simulated value for this motor). The air-gap needs to be the smallest possible; in the design, 1mm. Fig. 6 shows the active parts assembled and the rotors fabricated surface-mounted permanent-magnet mount. Windings were professionally hand-made; see Fig. 7. They were placed on flat-stator-core surface. To prevent the windings from missing its position and from vibration during motor operation, a type of epoxy resin was applied, giving the windings characteristics such as stiffness in working temperature, original dimensions, and good thermal conductivity for heat-release.

Fig. 8. Shaft with hole-through for phase-winding terminal outlet

IV.2. Driving System To rotate the motor, the stator windings should be energized in sequence. Knowledge of the rotors positioning is important, to understand which winding is energized following the energizing sequence. Rotor



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position was sensed by Hall Effect sensors embedded in stator. Whenever rotor magnetic poles passed near the Hall sensors, they gave high or low signals, indicating which pole (the N or the S) is passing. A combination of three of the Hall sensor signals enables determination of the exact sequence of commutation. Fig. 9 exemplifies Hall sensor signals related to back-EMF and phase current. Fig. 10 shows the Hall sensors position on a three-phase coreless stator.

(a) Switching diagram

Fig. 9. Single Hall-sensor position signal (green) on three-phase back-EMF

windings

Hall sensors

(b) Current-Flow to the Motor Winding at One Commutation-Step Figs. 11. Switching sequence

stator back iron

Fig. 10. Hall-Sensor Position on Stator

Fig. 11(a) shows the switching sequence to be followed relative to signals of the Hall sensors. Fig. 11(b) shows current to the motor winding at one commutation step. Each commutation sequence of a three-phase motor has one winding energized to positive power (current enters the winding), another to negative (current exits it), and another is not energized. Figs. 12 show the commutator circuit and the six-gate drive. One Hall sensor changes state for every 60 electrical degrees of rotation. One electrical cycle (360 electrical degrees) takes six steps to complete. Every 60 electrical degrees, phase-current switching is synchronously updated. One electrical cycle, however, may not correspond to one rotor revolution (mechanical).

(a) Commutator circuit

(b) Six gate-drive Figs. 12. Commutator circuit and six gate-drives



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Completion of a mechanical rotation is determined by rotor-pole pairs, via number of electrical cycles to be repeated. One rotor-pole pair completes one electrical cycle: fm

fe 2P

(25)

fe and fm are electrical and mechanical frequency respectively.

V.

Experiment Results

Fig. 13 shows the experiment test-bench set-up in University of Malayas Department of Electrical Engineering, for performance test of the in-wheel motor. A National Instrument Data Acquisition System with LabVIEW interface was used to obtain test data and plot performance curves. Motor torque and back-EMF were the main performance parameters obtained. During cruising-speed test, secondary measurements such as temperature rise in the motors critical parts were also recorded. Figs. 14 and 15 show graphs of back-EMF and torque. Back-EMF maximum output was 180V peak-to-peak, and torque output at rated current (30A) was about 37.4 N.m. Back EMF was acquired by mechanically turning the wheel at a particular speed, and then measuring terminal voltage. Motor speed was captured on a tachometer, which obtained the speed from the Hall-sensor pulse train. Under such conditions, the machine then acted as generator. At no-load condition, terminal voltage of the machine equaled generated back-EMF. Motor torque was measured on a load-cell force sensor, which was mounted on a free-rolling shaft. Constant, controlled current was injected into the motor from an inverter. The wheel was loaded with roller brake. Torque could be increased to maximum value quickly, and to twice the rated value. Input power (Pi) is the electric energy that runs the motor. Mechanical power delivered by the motor (Pm) is torque and speed, and overall motor efficiency is the ratio of input power to mechanical power: Pi

3V ph I ph

Pm

m

Pm Pi

Fig. 13. Motor experiment test-bench set-up with NI® data logger

Fig. 14. Comparison of Back-EMF results obtained from experiment and from FEM simulation

Fig. 15. Comparison of torque results obtained from experiment and from FEM simulation

(26) (27) (28)

Fig. 16 is a plot of measurement results, showing input power, mechanical output power, and overall efficiency. Results indicate the machines efficiency is more than 90%. Fig. 16. The motors performance-test results Copyright © 2011 Praise Worthy Prize S.r.l. - All rights reserved


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VI.

Conclusion

[12] F. Caricchi, F. Crescimbini, F. Mezzetti, E. Santini, Multistage Axial-Flux PM Machine for Wheel Direct Drive, IEEE Transactions on Industry Applications, Vol. 32 n. 4, July-August 1996, pp. 882-888. [13] C. C. Chan, K. T. Chau, J. Z. Jiang, W. Xia, M. Zhu, R. Zhang, Novel Permanent Magnet Motor Drives for Electric Vehicles. IEEE Transactions on Industrial Electronics, Vol. 43 n.2, April 1996, pp. 331-339. [14] A. Mahmoudi, N. A. Rahim, W. P. Hew, Analytical Method for Determining Axial-Flux Permanent-Magnet Machine Sensitivity to Design Variables, International Review of Electrical Engineering (IREE), vol. 5, no. 5, September-October 2010, pp. 2039-2048. [15] S. Asghar Gholamian, M. Ardebili. K. Abbaszadeh, Selecting and Construction of High Power Density Double-Sided Axial Flux Slotted Permanent Magnet Motors for Electric Vehicles, International Review of Electrical Engineering (IREE), vol. 4. n. 3, June 2009, pp. 477-484. [16] D. C. Hanselman, Brushless Permanent Magnet Motor Design (McGraw-Hill New York, 1994). [17] K. Sitapati and R. Krishnan, Performance Comparison of Radial and Axial Field Permanent Magnet Brushless Machines, IEEE Transactions on Industry Applications, vol.37, n. 5, SeptemberOctober 2001, pp. 1219-1226. [18] S. Huang, J. Luo, F. Leonardi and T. A. Lipo, A Comparison of Power Density for Axial Flux Machines Based on the General Purpose Sizing Equation, IEEE Transaction on Energy Conversion, Vol.14 n.2, June 1999, pp. 185-192. [19] J. F. Gieras, R. J. Wang, M. J. Kamper, Axial Flux Permanent Magnet Brushless Machines (Kluwer Academic Publisher, 2008).

The design, simulation, and testing of an AFPM wheel motor have been presented. Its high torque-density was the parameter of concern. The aim was for maximumtorque-density double-sided AFPM motor. Flux-densities of various parts of the motor were compared via sizing analysis, FEM, and experiment, each at no-load, all agreed in their results. Results of experiment and simulation show the motors actual back-EMF value during testing to be 90.4 Vmax at 700rpm. The test result was 4.8% less than that of the simulated result (95 Vmax). The difference could be due to the winding arrangement, which was slightly different during fabrication. Torque produced during experiment was 37.4Nm, with 30A input current, whereas torque produced in simulation was about 36.78Nm for the same condition. The motors design achieved the required motor specification. Its efficiency was 90%, and its design suits EV application.

References [1]

F. J. Perez-Pinal, C. Nunez, R. Alvarez, M. Gallegos, Step by Step Design of the Power Stage of a Light Electric Vehicle, International Review of Electrical Engineering (IREE), vol. 3. n. 1, January-February 2008, pp. 100-109. [2] P. Naderi, M. Mirsalim, S. M. T. Bathaee, Driving/Regeneration and Stability Enhancement for a Two-Wheel-Drive Electric Vehicle, International Review of Electrical Engineering (IREE), vol. 4 n. 1, January-February 2009, pp. 57-65. [3] S. Meo, F. Esposito, The EVALUATOR Suite for the Computer-aided Analysis of Advanced Automotive Electrical Power System, International Review of Electrical Engineering (IREE), vol. 2 n. 6, December 200, pp. 751-762. [4] F. Esposito, V. Isastia, S. Meo: Overview on Automotive Energy Storage Systems, International Review of Electrical Engineering (IREE), vol. 4 n. 6, November-December 2009, pp. 1122-1144. [5] F. Esposito, G. Gentile, V. Isastia, S. Meo, A New Bidirectional Soft-Switching Multi-Input DC-DC Converter for Automotive Applications, International Review of Electrical Engineering (IREE), vol. 5 n. 4, July-August 2010, pp. 1336-1346. [6] F. Esposito, V. Isastia, S. Meo, PSO Based Energy Management Strategy for Pure Electric Vehicles with Dual Energy Storage Systems, International Review of Electrical Engineering (IREE), vol. 5 n. 5, September-October 2010, pp. 1862-1871. [7] L. Chang, Comparison of AC Drives for Electric Vehicles-a Report on Experts' Opinion Survey, IEEE Aerospace and Electronic Systems Magazine, Vol. 9 n. 8, August 1994, pp. 7-11. [8] Y. P. Yang, Y. P. Luh, C. H. Cheun, Design and Control of AxialFlux Brushless DC Wheel Motors for Electric Vehicles-Part I: Multiobjective Optimal Design and Analysis. IEEE Transactions on Magnetics, Vol. 40 no. 4, pp. July 2004, 1873-1882. [9] B. B. Salah, A. Moalla, S. Tounsi, R. Neji, F. Sellami, Analytic Design Of A Permanent Magnet Synchronous Motor Dedicated To EV Traction With A Wide Range Of Speed Operation, International Review of Electrical Engineering (IREE), Vol. 3. n. 1, February 2008, pp. 110-122. [10] S. Kreuawan, F. Gillon, P. Brochet, Comparative Study of Design Approach for Electric Machine in Traction Application, International Review of Electrical Engineering (IREE), Vol. 3. n. 3, June 2008, pp. 455-465. [11] F. Dubas, C. Espanet, Exact Analytical Model of the No-Load Flux Density in the Air-gap, the Permanent Magnets and the Rotor Yoke for the Surface Mounted Permanent Magnet Motors, International Review of Electrical Engineering (IREE), Vol. 2. n. 1, June 2007, pp. 425-437.

Authors’ information Corresponding Author: Tel: +60136778050 Fax: 03-7967 5317 E-mail: [email protected] Nasrudin Abd. Rahim was born in Johor, Malaysia, in 1960. He received his B.Sc. (Hons.) degree in 1985, and his M.Sc. degree in 1988, both from the University of Strathclyde, Glasgow, UK. His Ph.D. degree was awarded in 1995 by Heriot-Watt University, Edinburgh, U.K. He is a Professor at the Department of Electrical Engineering, University of Malaya, Malaysia, Director of the University of Malaya Power Electronics, Drives, Automation and Control (UMPEDAC) Research Centre, and Chairman of University of Malaya Advanced Engineering & Technology Research Cluster. Dr. Rahim is a Fellow of the Institution of Engineering and Technology, UK, and a Chartered Engineer. He had been Chairman of IEEEs Power Engineering Society/Electric Machinery Committee Motor Subcommittee Working Group 8 (WG-8) covering reluctance motors. His research interests include power electronics, real-time control systems, electrical drives, and renewable energy (solar and wind). Hew Wooi Ping was born in Kuala Lumpur, Malaysia, in 1957. He obtained his Bachelor of Engineering (Electrical) degree in 1981, and his Master of Electrical Engineering degree from University of Technology, Malaysia. His Ph.D. degree was awarded in 2000 by University of Malaya, Kuala Lumpur, Malaysia. He is an Associate Professor at the Department of Electrical Engineering, University of Malaya. Dr. Hew is a Member of IET and a Chartered Engineer. His research interests include electrical drives, electrical machine design, application of fuzzy logic/neural network to electrical-machine-related applications, and renewable energy (solar and wind).



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Amin Mahmoudi was born in Bandar Abbas, Iran, in 1983. He received the B.S. degree in electrical engineering from Shiraz University, Shiraz, Iran in 2005 and the M.S. degree in electrical power engineering was awarded from Amirkabir University of Technology, Tehran, Iran, in 2008. He is currently a lecture at the Department of Engineering, HELP College of Arts and Technology, Kuala Lumpur, Malaysia. Mr. Mahmoudi is working toward the PhD degree in the Department of Electrical Engineering at University of Malaya, Kuala Lumpur, Malaysia. His research interests are numerical methods in electrical engineering, modeling and design of electrical machinery.



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