A Magnetic-Geared Outer-Rotor Permanent-Magnet Brushless Machine for Wind Power Generation

954 IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 45, NO. 3, MAY/JUNE 2009

A Magnetic-Geared Outer-Rotor Permanent-MagnetBrushless Machine for Wind Power Generation

Linni Jian, Student Member, IEEE, K. T. Chau, Senior Member, IEEE, and J. Z. Jiang

Abstract—This paper presents a new permanent-magnet (PM)brushless machine for wind power generation. This machineadopts an outer-rotor topology, aiming at capturing wind powerdirectly. In order to achieve high power density, a high-speed PMbrushless generator is artfully integrated with a coaxial magneticgear. The design details, with emphasis on the special constraintsof wind power generation, are elaborated. By using the time-stepping finite element method, the static characteristics as wellas no-load and on-load operations are simulated. A prototype isalso built for experimentation. Both simulation and experimentalresults are given to verify the validity of the proposed machine.Finally, a quantitative comparison is made to justify that theproposed machine is of smaller size, lighter weight, and lower costthan its counterparts.

Index Terms—Finite-element method (FEM), magnetic gear,outer rotor, permanent-magnet (PM) machine, wind powergeneration.

I. INTRODUCTION

W IND POWER as an abundant clean renewable energyresource has attracted increasing attention for solving

problems arising from energy crisis and environmental pollu-tion [1].1 Wind power generation is a major method of utilizingthis natural resource. Generally, it can be classified as constant-speed constant-frequency (CSCF) wind power generation andvariable-speed constant-frequency (VSCF) wind power genera-tion. In the CSCF system, the squirrel-cage induction generatoris usually adopted [2]. It offers several advantages, such asa simple structure and high robustness. Furthermore, it candirectly connect to the power grid without using any powerconverters. However, since the turbine speed is kept constantregardless of the variation of the wind speed, the CSCF systemsuffers from very low efficiency and high mechanical stress.

Paper IPCSD 08-087, presented at the 2007 Industry Applications SocietyAnnual Meeting, New Orleans, LA, September 23–27, and approved forpublication in the IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS bythe Electric Machines Committee of the IEEE Industry Applications Society.Manuscript submitted for review May 6, 2008 and released for publicationJanuary 28, 2009. Current version published May 20, 2009. This work wassupported by the Research Grants Council, Hong Kong Special AdministrativeRegion, China, under Grant HKU 7105/07E.

L. Jian is with the Department of Electrical and Electronic Engineering, TheUniversity of Hong Kong, Hong Kong (e-mail: [email protected]).

K. T. Chau is with the Department of Electrical and Electronic Engineering,The University of Hong Kong, Hong Kong (e-mail: [email protected]).

J. Z. Jiang is with the Department of Automation, Shanghai University,Shanghai 200030, China (e-mail: [email protected]).

Digital Object Identifier 10.1109/TIA.2009.2018974

1http://www.nrel.gov/docs/fy05osti/37602.pdf

With the development of power electronics, low-cost powerconverters make it possible to produce constant-frequency elec-tric power with a variable turbine speed. Since the turbinespeed changes with the wind speed to capture the maximumwind power, the efficiency of the VSCF system is muchhigher. For the VSCF system, several types of generators havebeen adopted or proposed, such as the doubly fed inductionmachine [3], switched reluctance machine [4], doubly salientpermanent-magnet (PM) machine [5], PM hybrid machine [6],and double-stator PM machine [7]. However, mechanical gearsare generally engaged to match the low-speed operation of thewind turbine and the relatively high-speed operation of thegenerator. This not only increases the cost of manufacture andmaintenance but also reduces the efficiency and robustness.

In order to get rid of the nuisances arising from mechanicalgears, the direct-drive PM machine (DDM) has been proposedfor wind power generation [8]. Since the machine has to operateat low speeds, it needs to have a bulky size with a verylarge number of poles. Therefore, it is a tradeoff between theelimination of mechanical gears and the increase of machinesize. Sometimes, the use of mechanical gears plus a high-speedmachine can have the reduction of overall size and weight ascompared with the low-speed direct-drive machine.

Recently, the concept of magnetic gears has been proposed[9], [10]. Because of physical isolation between the inputand output shafts, it offers some distinct advantages: namely,minimum acoustic noise, freedom from maintenance, improvedreliability, and inherent overload protection. By adopting thecoaxial topology, the utilization of PMs can be greatly im-proved, leading to the offering of a torque density comparablewith that of the mechanical gear. Very recently, it has beenintegrated into a PM motor to offer high-torque low-speedoperation for electric vehicles [11].

The purpose of this paper is to propose and implement amagnetic-geared outer-rotor PM brushless machine for windpower generation, which incorporates the attractive features ofboth the outer-rotor PM generator and the magnetic gear. InSection II, the configuration of a stand-alone wind power gener-ation system and the machine design details will be introduced.Section III will be devoted to using finite element analysisto deduce the static characteristics of the proposed machine.In Section IV, the mathematical modeling of the proposedmachine will be presented. Then, simulation and experimentalresults will be given to verify the validity of the proposedmachine in Sections V and VI, respectively. Consequently, aquantitative comparison between the proposed machine and itscounterparts will be made in Section VII. Finally, conclusionswill be drawn in Section VIII.

0093-9994/$25.00 © 2009 IEEE

JIAN et al.: MAGNETIC-GEARED OUTER-ROTOR PM BRUSHLESS MACHINE FOR WIND POWER GENERATION 955

Fig. 1. Wind power generation system.

Fig. 2. Configuration of proposed machine.

II. SYSTEM DESIGN

A. System Configuration

Fig. 1 shows the configuration of a stand-alone wind powergeneration system in remote areas. It consists of the proposedPM brushless machine to directly capture wind power, a three-phase bridge rectifier to perform simple ac–dc conversion, adc–dc converter to regulate the rectified dc voltage, a batterypack for energy storage, and an inverter to perform simpledc–ac conversion.

The detailed configuration of the proposed machine for windpower generation is shown in Fig. 2. The coaxial magnetic gearis artfully integrated into the PM generator such that they canshare the same high-speed rotor (namely, the outer rotor ofthe generator and the inner rotor of the magnetic gear). Thiscommon rotor is designed like a cup with PMs mounted on itsinside and outside surfaces. The magnetic gear employs PMson both the inner and outer rotors and has a stationary ringbetween the two rotors. In order to provide magnetic pathswhile reducing iron losses, both the stationary ring and the ironyokes of the two rotors are built of laminated ferromagneticmaterials. These laminated materials are based on cold-rolledsilicon steel (type BW315) with a thickness of 0.35 mm.Moreover, epoxy is filled in the slots of the stationary ring toenforce the structural strength for high torque transmission. Forcapturing wind power directly, the wind blades are mounted onthe gear outer rotor. Fig. 3 shows the mechanical assembly ofthe proposed machine, in which three bearings are employed toguarantee free rotations of both rotors.

Fig. 3. Mechanical assembly of proposed machine.

B. Machine Design

The operation of the magnetic gear relies on the use of astationary ring to modulate the magnetic fields [9]. By definingp1 and p2 as the pole-pair numbers of the outer and inner rotors,respectively, and ns as the number of ferromagnetic pole pieceson the stationary ring, it yields

ns = p1 + p2. (1)

The corresponding speed relationship can be expressed as

ω2 = −Grω1 (2)

Gr =p1

p2(3)

where ω1 and ω2 are the rotational speeds of the outer andinner rotors, respectively, and Gr is the so-called gear ratio.The minus sign indicates that the two rotors rotate in oppositedirections.

Equations (1)–(3) indicate that the gear ratio is determined bythe pole-pair numbers of both rotors. Thus, we should chooseproper pole-pair numbers according to the wind-turbine andgenerator speeds. Based on the Betz theory [12], the mechanicalpower extracted from wind power by the wind turbine can beexpressed as

Pmech =12Cpρv3

wA (4)

where ρ is the air density, vw is the wind speed, A is the sweptarea of the wind-turbine rotor, and Cp is the power conversionefficiency factor. This Cp is a function of the tip speed ratio β,which is defined as

β =ωR

vw(5)

where R is the radius of the blades and ω is the rotational speedof the wind-turbine shaft. For the proposed machine, it equalsthe gear outer-rotor speed ω1. The characteristic of Cp = f(β)is approximately parabolic. When β takes the specific valueβopt, the maximum power conversion efficiency factor Cpmax

occurs. In order to extract the maximum mechanical power, ω1

should vary with the wind speed. Fig. 4 shows the mechanicalpower characteristics of a wind turbine under different wind


Fig. 4. Mechanical power versus turbine speed at different wind speeds.

speeds. The maximum power curve indicated by the dotted linecan be expressed as

Pmechmax =12

(CpmaxπR5ρ

β3max

)ω3

1 . (6)

Thus, the rated outer-rotor speed can be determined by thewind speed according to the power curve. For an averagewind speed of 7 m/s in some areas, from Fig. 4, ω1 shouldbe rated at 136 r/min. Meanwhile, the inner-rotor speed ω2

is rated at 1000 r/min for producing 50 Hz of electricity. Bychoosing p1, p2, and ns as 3, 22, and 25, respectively, the gearratio of 7.33 results so as to match the rated speeds of thetwo rotors.

Considering the power generation under gentle breeze, thereduction of the starting torque is another design goal. In PMmachines, cogging torque is a major factor which results in ahigh starting torque and deteriorates the low-speed operation.To reduce this cogging torque, the skewed slot design is usuallyadopted. However, it degrades the generated output powerand complicates the manufacturing process [5]. Therefore, theproposed machine does not adopt the use of skewing.

Generally, the fundamental order of the cogging torque isthe smallest common multiple between the stator slot numberand the rotor pole number. The higher the fundamental order,the lower the resulting harmonic magnitude. Two machineswith exactly the same size but different numbers of statorslots equal to 9 and 27 are selected for analysis. Since theinner-rotor pole number is six, the fundamental order of thecogging torque existing in the former machine is 18, and thatin the latter machine is 54. Fig. 5 shows the cogging torquewaveforms calculated by the finite element method (FEM).The magnitude of the cogging torque in the machine with27 stator slots is about 0.7 N · m which is much less than2 N · m in the machine with nine stator slots. The use of moreslots is helpful to reduce cogging torque; however, it leads to theincrease of the manufacturing complexity. To make a compro-mise, the use of 27 stator slots is adopted to reduce the coggingtorque.

Fig. 6 shows the stator winding connection of the pro-posed machine. The three-phase symmetric windings consist of27 double-layer coils. Each coil span covers five slot pitches,while the pole pitch is 9/2 of the slot pitch.

Fig. 5. Cogging torque in inner rotor. (a) Stator with 27 slots. (b) Stator withnine slots.

Fig. 6. Stator winding connection.

TABLE ISPECIFICATIONS OF PROPOSED MACHINE

III. FINITE-ELEMENT ANALYSIS

A. Magnetic Field Distributions

The specifications of the proposed machine are listed inTable I. Fig. 7 shows the magnetic field distributions of the


Fig. 7. Magnetic field distributions. (a) No-load. (b) Full-load.

proposed machine under no-load and full-load. It can be ob-served that a large portion of flux lines can pass through allthree air gaps. These flux lines dictate the torque transmissionand power conversion. There are also some flux lines directlyturning around the boundaries of adjacent PMs. These flux linesare useless for transmitting torque or power and arouse powerloss. Fig. 8 shows the radial flux density waveforms in theinner, middle, and outer air gaps at no-load. The radial fluxdensity waveforms at full-load are almost the same with thatat no-load, and their little differences are shown in Fig. 9. Itdemonstrates that the armature reaction just has a slight effecton the magnetic gear. This is due to the airlike permeabilityof the surface-mounted PMs which significantly increase thereluctance of the armature flux path.

Moreover, Fig. 10 shows the radial flux density waveformand its harmonic spectrum in the outer air gap produced by theinner rotor alone. It can be seen that the highest asynchronousharmonic component is that with 22 pole pairs. On the otherhand, Fig. 11 shows the radial flux density waveform and itsharmonic spectrum in the middle air gap produced by the outerrotor alone. It can be found that the highest asynchronousharmonic component is that with three pole pairs. Therefore, itconfirms that the stationary ring can modulate the flux densityfrom 3 to 22 pole pairs.

Fig. 8. Radial flux densities at no-load. (a) Inner air gap. (b) Middle air gap.(c) Outer air gap.

Fig. 9. Radial flux density differences between no-load and full-load.(a) Inner air gap. (b) Middle air gap. (c) Outer air gap.


Fig. 10. Radial flux density in outer air gap produced by inner rotor. (a) Fluxdensity waveform. (b) Harmonic spectrum.

Fig. 11. Radial flux density in middle air gap produced by outer rotor. (a) Fluxdensity waveform. (b) Harmonic spectrum.

B. Torque Transmission

By calculating the Maxwell’s stress tensors in the outerand middle air gap, the torque transmission capability of themagnetic gear can be determined. While keeping the outerrotor at a standstill, the inner rotor is rotated step by step.The corresponding torque-angle curves are calculated as shownin Fig. 12. It can be found that the torque-angle curves varysinusoidally, in which the maximum torque values denote thepull-out torques. On the inner and outer rotors, the pull-outtorques are 14.2 and 103.4 N · m, respectively. Their ratio is7.28, which has a good agreement with the ratio of the pole-pair numbers of two rotors equal to 7.33. According to themachine size listed in Table I, the overall torque density is about87 kN · m/m3, which is much higher than that of a standardelectric machine with only about 10 kN · m/m3. There is a πdifference between the phase angles of two torque waveforms.This implies that the two rotors rotate in opposite directions,which is in accordance with (2).

Fig. 12. Torque-angle curves.

Fig. 13. Equivalent circuit.

IV. MATHEMATICAL MODELING

The back electromotive force (EMF) induced in the statorwindings can be expressed as

E = −dΨdt

= Uo + RsIs (7)

where Ψ is the flux linkage in the inner air gap, Uo and Is arethe output voltage and current, respectively, and Rs is the statorresistance. The flux linkage Ψ is given by

Ψ = Ψpm + Ψs = Ψpm + LsIs (8)

where Ψpm is the PM excited flux linkage, Ψs is the fluxlinkage due to the armature reaction, and Ls is the statorreactance. Ignoring the nonlinear factors in the magnetic path,it yields

Ψpm = Ψpm1 + Ψpm2 (9)

where Ψpm1 and Ψpm2 are the flux linkages in the inner airgap produced by the outer- and inner-rotor PMs, respectively.Therefore, the voltage equation is given by

Uo = − dΨpm1

dt− dΨpm2

dt− Ls

dIs

dt− Is

dLs

dθω2 − RsIs

≈E1 + E2 − jIsp2ω2Ls − RsIs (10)

where the variation of Ls with respect to the rotor positionangle θ is negligible because of the nonsaliency due to thesurface-mounted PMs. The corresponding equivalent circuitand phasor diagram are shown in Figs. 13 and 14, respectively.


Fig. 14. Phasor diagram.

V. SIMULATION RESULTS

In order to assess the proposed machine, a computer simula-tion based on time-stepping FEM is conducted. The simulationmodel consists of two equations: the finite element equationof the electromagnetic field of the proposed machine and thecircuit equation of the armature windings [13]. The 2-D elec-tromagnetic equation is governed by

Ω :∂

∂x

(v∂A

∂x

)+

∂

∂y

(v∂A

∂y

)

= −J − v

(∂Bry

∂x− ∂Brx

∂y

)+ σ

∂A

∂t

s1 : A = 0 (11)

where Ω is the region of calculation, A is the magnetic vectorpotential component along the z-axis, J is the current density,v is the reluctivity, σ is the electrical conductivity, Brx andBry are the remnant flux density components of the PM alongthe x- and y-axes, and s1 is the boundary of the region ofcalculation.

According to the equivalent circuit shown in Fig. 13, thearmature circuit equation is given by

(Rs + RL)is + (Ls + LL)disdt

− l

S

∫∫Ω

∂A

∂tdΩ = 0 (12)

where RL is the load resistance, LL is the load inductance, l isthe axial length of the iron core, and S is the conductor area ofeach turn of phase winding.

Fig. 15 shows the output phase voltages at no-load and with aresistive load of 300 Ω, when the outer rotor rotates at the ratedspeed of 136 r/min.

VI. EXPERIMENTAL RESULTS

As shown in Fig. 16, the proposed machine is prototyped forexperimentation. First, by driving the outer rotor at the ratedspeed of 136 r/min, the measured speed at the inner rotor is993 r/min, verifying that the speed reduction ratio is 7.3 whichagrees with the theoretical Gr = 7.33. Then, the correspondingno-load EMF waveform and the output-voltage waveform with

Fig. 15. Output voltages at rated speed. (a) No-load. (b) On-load.

Fig. 16. Proposed machine experimentation. (a) Gear outer rotor. (b) Station-ary ring. (c) Gear inner rotor. (d) Assembled machine. (e) Test bed.

a resistive load of 300 Ω are measured, as shown in Fig. 17. Itcan be seen that there is a good agreement with the simulationwaveforms, as shown in Fig. 15. Furthermore, Fig. 18 showsa comparison of the measured and simulated amplitudes of


Fig. 17. Measured output voltages at rated speed (150 V/div, 4 ms/div).(a) No-load. (b) On-load.

Fig. 18. No-load EMF amplitudes at different outer-rotor speeds.

Fig. 19. Measured efficiency versus load current at different speeds.

the no-load EMF per phase at different outer-rotor speeds. Asexpected, they are in good agreement, and the voltage is linearlyproportional to the rotational speed.

Moreover, the efficiency of the proposed machine is mea-sured at different load currents and outer-rotor speeds. Asshown in Fig. 19, the efficiency at the rated speed of 136 r/minand the rated power of 500 W is about 72%, whereas themaximum efficiency is about 80% which occurs at 150 r/min.

TABLE IIMACHINE RATINGS FOR COMPARISON

In order to figure out the reasons for a relatively low effi-ciency of the prototype machine, a breakdown of all losses isestimated. First, the no-load loss, which combines the iron andmechanical losses, is directly measured at the rated speed. It isfound to be 174 W at 136 r/min. Second, the corresponding ironloss is calculated by using the FEM. It is estimated to be 36.7 W.Hence, the corresponding mechanical loss can be deduced as137.3 W. Such high mechanical loss should be due to frictionresulting from manufacturing imperfections. Third, the copperloss at the rated power of 500 W can be easily calculated,namely, 18.5 W. Thus, the breakdown of all losses of themachine operating at the rated power and rated speed is 9.6% incopper loss, 19.1% in iron loss, and 71.3% in mechanical loss.Based on this loss analysis, it can be realized that the major lossis due to the mechanical loss which will be less significant whenthe power level of the machine is elevated.

It should be noted that the prototype generator is mainly usedfor experimental verification. Its power level should be elevatedto about 50 kW so as to justify the use of magnetic gearing withthe rated rotational speed of 136 r/min.

VII. COMPARISONS

To further demonstrate the merits of the proposed magnetic-geared PM machine (MGM), a quantitative comparison with itstraditional counterparts, namely, the DDM and the planetary-geared PM machine (PGM), is conducted. For a fair com-parison, the machine ratings are unified as listed in Table II.Furthermore, all these machines adopt the outer-rotor topologyand three-phase star-connected stator windings.

A. Overall Size

It is formidable to make a precise assessment on the sizes ofdifferent types of machines. Nevertheless, a rough comparisoncan be achieved based on the following sizing equation [14]:

D2i l =

4Pr

kekikpBδ maxAωr(13)

A =2mNIrms

πDi(14)

where Di is the diameter of the air gap (inner air gap for theMGM), l is the effective axial length, Pr is the rated power, ke

is the EMF waveform factor, ki is the current waveform factor,kp is the power waveform factor, ωr is the rated rotor speed(inner-rotor speed for the MGM), Bδ max is the maximum fluxdensity in the air gap (inner air gap for the MGM), A is theelectric loading, m is the number of phases, N is the number


of turns of the phase winding, and Irms is the rms value of thephase current.

Since the rectangular waveforms can produce a higher out-put power than the sinusoidal ones, the waveform factors areselected as ke = π, ki = 1, and kp = 1 for all three typesof machines. Based on engineering practice, A is selected as15 000 A/m for all of them. Then, Bδ max is selected as 1.1 Tfor the DDM and PGM, whereas Bδ max is set as 0.85 T forthe MGM since it involves a relatively long magnetic path. Asdesired, ωr is 1000 r/min for the MGM and PGM, whereas ωr

is 136 r/min for the DDM. Since the ratio λ = l/Di is governedby ωr, λ is selected as 0.33 for the MGM and PGM, whereasλ is set as 0.15 for the DDM. By substituting all these valuesinto (13), it can be deduced that the Di’s of the MGM, DDM,and PGM are 30.7, 71.2, and 28.2 cm, respectively, and thecorresponding l’s are 10.1, 10.7, and 9.3 cm, respectively.

Based on the previous design of the proposed MGM, the sizeof the magnetic gear outer part of the MGM can be estimatedas 20 101 cm3. Therefore, the overall size of the MGM is27 577 cm3. In the outer rotor of the DDM and PGM, the PMthickness is chosen as 0.8 cm which is sufficient to excite thedesired air-gap flux density. To avoid unnecessary saturation,the thickness of the iron yoke is chosen as 2.0 cm for theDDM and 1.0 cm for the PGM. Thus, the size of the DDMcan be deduced as 49 567 cm3, and that of the PGM (withoutthe planetary gear) is 7386 cm3. Taking into account the size ofthe planetary gear, which is 24 389 cm3 for the rating of 10 kWand the ratio of 1000 to 136 r/min, the overall size of the PGMis 31 775 cm3.

B. Overall Weight

Based on the sizes of the different materials used in themachine, namely, the copper windings, iron cores, and PMs,the weights of these materials and, hence, the overall weightof the machine can be calculated. For the MGM, the requiredcopper-winding material is 292 cm3, the iron-core material is15 521 cm3, and the PM material is 1715 cm3. Since thedensities of the copper-winding, iron-core, and PM materialsare 9.0, 7.8, and 7.6 g/cm3, respectively, the overall weight canbe calculated as 137 kg. Similarly, for the DDM, the requiredcopper-winding material is 718 cm3, the iron-core material is42 179 cm3, and the PM material is 1936 cm3. Therefore,based on their densities, the overall weight can be calculated as350 kg.

For the PGM (without the planetary gear), the requiredcopper-winding material is 247 cm3, the iron-core material is5816 cm3, and the PM material is 678 cm3. Therefore, based onthe corresponding densities, the weight of the PGM (without theplanetary gear) can be calculated as 53 kg. Taking into accountthe weight of the planetary gear, which is made of steel withthe density of 7.9 g/cm3, the overall weight of the PGM can beobtained as 138 kg.

C. Material Cost

It is hardly possible to accurately compare the overall costof the three machines, unless they are under mass production,since the assembly cost heavily depends on the maturity and

TABLE IIICOMPARISON WITH COUNTERPARTS

complexity of those machines. Therefore, in this stage, the costcomparison of these three machines is simply focused on thematerial cost only.

According to the indicative prices of those materials usedin the three machines, namely, $11.2/kg for copper windings,$1.22/kg for iron cores, $45.7/kg for PMs, and $2.43/kg forthe steel planetary gear, the material costs can readily bededuced from their weights. Therefore, the overall materialcosts of the MGM, DDM, and PGM are $769, $1141, and $396,respectively.

D. Discussion

The estimated data of the three types of machines, namely,the overall size, overall weight, and material cost, are sum-marized in Table III for direct comparison. It can be observedthat the proposed MGM takes the definite advantage of thesmallest size and lightest weight as compared with the DDMand PGM. In terms of material cost, the MGM is morecostly than the PGM, but still much cheaper than the DDM.The additional cost of the MGM over that of the PGM is welljustified and outweighed by the significant reduction in size andthe aforementioned advantages of the magnetic gear. Therefore,the proposed MGM is highly competitive for wind powergeneration.

It should be noted that the proposed generator and its proto-type are mainly used to illustrate the concept of magnetic gear-ing for wind power generation. The corresponding equationsgiven by (4)–(6), the comparisons listed in Tables II and III,as well as the illustrations shown in Figs. 1 and 4 are used todescribe general wind-turbine applications.

VIII. CONCLUSION

In this paper, a new magnetic-geared PM brushless machine,which is particularly attractive for wind power generation, hasbeen designed, analyzed, and implemented. Compared with thetraditional wind power generators, the proposed machine offerssome distinct advantages. First, the low-speed outer-rotor topol-ogy can enable direct coupling with the wind blades to capturewind power with high efficiency. Second, the integration of acoaxial magnetic gear can enable the PM brushless generatorto be designed for high-speed operation, hence achieving highpower density. Third, the use of the magnetic gear can providephysical isolation between the inner rotor and the outer rotor,thus minimizing the maintenance cost and the acoustic noise.Finally, a quantitative comparison has illustrated that the pro-posed machine is of smaller size and lighter weight than boththe direct-drive PM brushless machine and the planetary-geared


PM brushless machine, with also lower material cost than thedirect-drive one.

REFERENCES

[1] N. B. Negra, O. Holmstrom, B. J. Birgitte, and S. Poul, “Windfarm gen-eration assessment for reliability analysis of power systems,” Wind Eng.,vol. 31, no. 6, pp. 383–400, 2007.

[2] A. Grauers, “Efficiency of three wind energy generator systems,” IEEETrans. Energy Convers., vol. 11, no. 3, pp. 650–657, Sep. 1996.

[3] S. Muller, “Doubly fed induction generator system for wind turbines,”IEEE Ind. Appl. Mag., vol. 8, no. 3, pp. 26–33, May/Jun. 2002.

[4] D. Torrey, “Switched reluctance generators and their control,” IEEETrans. Ind. Electron., vol. 49, no. 1, pp. 3–14, Feb. 2002.

[5] Y. Fan, K. T. Chau, and M. Cheng, “A new three-phase doubly salientpermanent magnet machine for wind power generation,” IEEE Trans. Ind.Appl., vol. 42, no. 1, pp. 53–60, Jan./Feb. 2006.

[6] K. T. Chau, Y. B. Li, J. Z. Jiang, and S. Niu, “Design and control of aPM brushless hybrid generator for wind power application,” IEEE Trans.Magn., vol. 42, no. 10, pp. 3497–3499, Oct. 2006.

[7] S. Niu, K. T. Chau, J. Z. Jiang, and C. Liu, “Design and controlof a new double-stator cup-rotor permanent-magnet machine for windpower generation,” IEEE Trans. Magn., vol. 43, no. 6, pp. 2501–2503,Jun. 2007.

[8] J. Chen, C. V. Nayar, and L. Xu, “Design and finite-element analy-sis of an outer-rotor permanent magnet generator for directly coupledwind turbines,” IEEE Trans. Magn., vol. 36, no. 5, pp. 3802–3809,Sep. 2000.

[9] K. Atallah and D. Howe, “A novel high performance magnetic gear,”IEEE Trans. Magn., vol. 37, no. 4, pp. 2844–2846, Jul. 2001.

[10] P. O. Rasmussen, T. O. Andersen, F. T. Jorgensen, and O. Nielsen, “De-velopment of a high-performance magnetic gear,” IEEE Trans. Ind. Appl.,vol. 41, no. 3, pp. 764–770, May/Jun. 2005.

[11] K. T. Chau, D. Zhang, J. Z. Jiang, C. Liu, and Y. Zhang, “Design ofa magnetic-geared outer-rotor permanent-magnetic brushless motor forelectric vehicles,” IEEE Trans. Magn., vol. 43, no. 6, pp. 2504–2506,Jun. 2007.

[12] R. Datta and V. T. Ranganathan, “A method of tracking the peak powerpoints for a variable speed wind energy conversion system,” IEEE Trans.Energy Convers., vol. 18, no. 1, pp. 163–168, Mar. 2003.

[13] Y. Wang, K. T. Chau, C. C. Chan, and J. Z. Jiang, “Transient analysisof a new outer-rotor permanent-magnet brushless dc drive using circuit-field-torque coupled time-stepping finite-element method,” IEEE Trans.Magn., vol. 38, no. 2, pp. 1297–1300, Mar. 2002.

[14] S. Huang, J. Luo, F. Leonardi, and T. A. Lipo, “A general ap-proach to sizing and power density equations for comparison of elec-trical machines,” IEEE Trans. Ind. Appl., vol. 34, no. 1, pp. 92–97,Jan./Feb. 1998.

Linni Jian (S’07) received the B.Eng. degree fromHuazhong University of Science and Technology,Wuhan, China, in 2003, and the M.Eng. degreefrom the Institute of Electrical Engineering, ChineseAcademy of Science, Beijing, China, in 2006. Cur-rently, he is working toward the Ph.D. degree in elec-trical and electronic engineering at The University ofHong Kong, Hong Kong.

His research interests are in the areas of electricdrives, electric vehicles, and power electronics. Hecurrently focuses on the design of magnetic gears

and integrated permanent-magnet machines, as well as high-efficiency machinecontrol strategies.

K. T. Chau (M’89–SM’04) received the first-classhonors B.Sc.(Eng.), M.Phil., and Ph.D. degrees inelectrical and electronic engineering from The Uni-versity of Hong Kong, Hong Kong, in 1988, 1991,and 1993, respectively.

He is currently a Professor with the Departmentof Electrical and Electronic Engineering and theDirector of the International Research Center forElectric Vehicles, The University of Hong Kong. Histeaching and research interests focus on three mainareas: electric vehicles, electric drives, and power

electronics. In these areas, he has published over 200 refereed technical papers.He is also the coauthor of a monograph, Modern Electric Vehicle Technology(Oxford University Press, 2001).

Prof. Chau is a Fellow of the Institution of Engineering and Technology.He was the recipient of the Outstanding Young Researcher Award from TheUniversity of Hong Kong in 2003, the University Teaching Fellowship Awardfrom The University of Hong Kong in 2004, and the Award for Innovative Ex-cellence in Teaching, Learning and Technology at the International Conferenceon College Teaching and Learning in 2005.

J. Z. Jiang received the B.E.E. and M.E.E. de-grees from Shanghai Jiao Tong University, Shanghai,China, in 1962 and 1965, respectively, and theDr. Ing. degree from the Technical University ofBraunschweig, Braunschweig, Germany, in 1988.

He is currently a Professor with the Depart-ment of Automation, Shanghai University, Shanghai,China. His research interests are in high-performancevariable-speed drives, electric machine design, elec-tric vehicles, and wind power generation.

A Magnetic-Geared Outer-Rotor Permanent-Magnet Brushless Machine for Wind Power Generation

Documents

Transcript of A Magnetic-Geared Outer-Rotor Permanent-Magnet Brushless Machine for Wind Power Generation