Electric Drive System of Dual-Winding Fault-Tolerant Permanent-Magnet Motor for Aerospace Applications

Transcription

1 73 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 6, NO., DECEMBER 05 Electric Drive System of Dual-Winding Fault-Tolerant Permanent-Magnet Motor for Aerospace Applications Xuefeng Jiang, Student Member, IEEE, Wenxin Huang, Member, IEEE, Ruiwu Cao, Member, IEEE, Zhenyang Hao, and Wen Jiang Abstract In this paper, a new electric drive system based on a six-phase ten-pole dual-winding fault-tolerant permanent-magnet (DFPM) motor for aerospace applications is proposed and investigated. The proposed DFPM motor consists of optimal surface-mounted permanentmagnet (PM) rotor and -slot stator with two sets of independent three-phase concentrated armature windings on alternate teeth, which incorporates the merits of high power density and high efficiency of the PM motor and high fault tolerance of the dual-winding motor. To achieve the fault-tolerant performance of the electric drive system, the redundant fault-tolerant control strategy based on the failure diagnosis and redundancy communication is proposed and applied to the electric drive system. Then, the operational performance of the drive system under open-circuit and short-circuit fault has been verified by simulation and experiments. The research results show that the opencircuit and short-circuit fault can be detected immediately and the output performance of the DFPM motor under fault conditions is almost the same as the normal condition, which verified the proposed electric drive system, DFPM motor, and fault-tolerant control strategy. Index Terms Dual-winding motor, fault-tolerant drive, open-circuit or short-circuit fault, permanent-magnet machine, redundancy. I. INTRODUCTION WITH the increasing development of aircraft, new moreelectric and all-electric aircraft in the aviation sector have attracted increasingly more attention [] [3]. The typical characteristic of the more-electric or all-electric aircraft is that part or all pneumatic and hydraulic systems are replaced by Manuscript received October 7, 04; revised February 3, 05 and April 4, 05; accepted May 3, 05. Date of publication July 9, 05; date of current version November 6, 05. This work was supported in part by the National Natural Science Foundation of China under Project and Project , in part by the Natural Science Foundation of Jiangsu Province under Grant BK0386 and Grant BK04084, in part by the Professional Research Foundation for Advanced Youth Talents of Nanjing University of Aeronautics and Astronautics under Project YAH40, and in part by the Funding of Jiangsu Innovation Program for Graduate Education under Grant KYLX_067. The authors are with the Key Laboratory of New Energy Generation and Power Conversion of Jiangsu Province, Nanjing University of Aeronautics and Astronautics, Nanjing 006, China ( jxf@nuaa.edu.cn; huangwx@nuaa.edu.cn; ruiwucao@nuaa.edu.cn; zhenyang_hao@nuaa.edu.cn; nuaa_jw@63.com). Color versions of one or more of the figures in this paper are available online at Digital Object Identifier 0.09/TIE electrical drive systems, which can reduce the running cost, aircraft s volume and weight, and fuel cost and improve the reliability and maintainability of the aircraft [4]. It is estimated that the weight and fuel cost of an aircraft can be reduced by 0% and 9% by adopting all-electrical aircraft, respectively [5]. More-electric aircraft is a transitional scheme from conventional aircraft to all-electric aircraft. At present, the mainly more-electric aircraft includes the European Airbus A380, American Boeing B787, and Lockheed Martin F-35 [5], [6]. The current research shows that the design of the electrical drive system is one of the key technologies of more-electric aircraft [6], [7]. Compared with the conventional hydraulic fuel pump system, the electrical drive system can not only improve the system efficiency and the flexibility of variable speed control but also reduce the weight and volume of the system in the aircraft fuel pump system [3] [5]. In the aircraft fuel pump system, where continuous operation must be ensured, reliability may be a critical requirement. As we know, the motor and its drive system are the core parts of the electrical drive system, particularly in an electrical drive aircraft fuel pump system. However, due to the problems of insulation aging of armature winding, characteristic change of electronic components, and electromagnetic interference, the motor drive system will inevitably occur faults in the motor or power converter. Generally, the motor faults mainly include open-circuit and short-circuit fault in the phase winding, and the power converter faults mainly include the open-circuit fault and short-circuit fault in the insulated-gate bipolar transistor (IGBT). Therefore, in addition to meet some specific functions, the motor drive system for aerospace applications must have high reliability and strong fault tolerance [8], [9]. The redundancy motor drives have been presented in many papers [0] []. Redundancy technology is a method to improve the reliability of the motor drive system by adding extra resources such as hardware or software [0]. The parallel dualredundancy motor is the mostly used redundancy motor, which consists of two sets of independent windings with 30 electrical degree shift in space, two sets of position sensors, and a mutual rotor. This system has some merits, such as simple principle, clear structure, and easy control strategy []. However, the common redundancy motor has a lethal defect in the electric drive systems for aerospace applications because the traditional redundancy motors all adopt permanent-magnet (PM) motor with distributed windings, which will lead to a great shortcircuit current and burn out the motor windings when the phase winding occurs short-circuit fault. Moreover, it will affect the IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See for more information.

2 JIANG et al.: ELECTRIC DRIVE SYSTEM OF DFPM MOTOR FOR AEROSPACE APPLICATIONS 733 normal phase windings through the magnetic coupling and make the system output a pulsating electromagnetic torque, which makes the motor drive system work incorrectly [3] [5]. Because of the robust rotor structure and the inherent fault tolerance, the switched reluctance (SR) motor has been widespread concerned in the aerospace area, which can overcome the drawbacks of traditional redundant motor control technology [3]. However, compared with the PM motor, the SR motor suffers from drawbacks of lower power density, larger torque ripple, bigger noise, and lower efficiency [3] [6]. Hence, fault tolerance was introduced into PM motor drives to achieve high power density. In [3], a fault-tolerant motor having PMs in the rotor was presented, and a comparison of a fault-tolerant PM-rotor motor with an SR motor was made. Further researches on this machine show the high fault-tolerant and electromagnetic performance [6] [4]. Ever since 996, Prof. B. C. Mecrow has proposed nonbackup fault-tolerant PM motors with six-phase eight-pole and four-phase six-pole and completed their fault identification and fault-tolerant control by looking up the table of flux linkage current rotor position for aerospace applications [8], [8], [9]. In addition, Prof. Howe proposed a modular fault-tolerant PM brushless motor and an optimal torque control strategy based on the current hysteresis control method [6], [0], []. However, compared with the redundancy motor drive system, the existing faulttolerant PM motor drive systems suffer from some drawbacks, such as double power switches because of each phase winding driven by H-bridge circuit, lower power density and reliability, and complex control algorithms [4], [5], []. The purpose of this paper is to propose an electric drive system of a six-phase ten-pole dual-winding fault-tolerant PM (DFPM) motor for aerospace applications. The key is to propose a high fault-tolerant and electromagnetic performance DFPM motor and two sets of three-phase full-bridge drive circuits with fault-tolerant control strategy. In Section II, the structure, short-circuit current, overload ratio, cogging torque, and fault-tolerant characteristic are analyzed. In Section III, the mathematical model of the proposed DFPM motor is built and analyzed. In Section IV, the active active redundant faulttolerant control strategy based on vector control, the failure diagnosis, and the redundancy communication strategy are proposed and investigated. In Section V, the proposed fault strategies are adopted to analyze the behaviors of the proposed DFPM motor under open-circuit and short-circuit fault conditions. To examine the fault-tolerant performance, experimental results are presented in Section VI. Finally, conclusions are drawninsectionvii. II. DFPM MOTOR SYSTEM CONFIGURATION AND DESIGN A. Topology Configuration The topology configuration of the proposed DFPM motor system is shown in Fig., in which a high fault-tolerant and electromagnetic performance DFPM motor and two sets of three-phase full-bridge drive circuits are adopted in an electrical drive aircraft fuel pump system. Full rated torque can be Fig.. Topology of the proposed DFPM motor drive system. Fig.. Configuration of the proposed DFPM motor. provided by isolated abc windings or isolated xyz windings under fault conditions. Each redundancy consists of an electrically isolated three-phase full-bridge drive and a regenerative energy dump circuit. The drives are composed of many single 600-V IGBTs of Microsemi Corporation, and each redundancy is supplied by a separate 70 V, as shown in Fig.. Compared with the existing fault-tolerant PM motor drive systems whose each phase winding was driven by H-bridge circuit [3] [5], the drive system of the proposed DFPM motor can not only reduce the number of power switches and save the system cost but also improve the reliability and power density of the system. B. Design of Short-Circuit Current and Overload Ratio To ensure a limited value of the short-circuit current of the DFPM motor, we have to increase the self-inductance. Because the DFPM motor has the characteristics of magnetic isolation and the mutual inductance between adjacent phase windings is very small, the steady-state short-circuit current Īs is given by Ī s = Ē 0 (ω e L s ) + R s () L s = L sm + L sσ () where Ē0 is the no-load back electromotive force (EMF), R s and L s are the phase resistance and phase self-inductance, respectively. ω e is the electrical angular velocity of the rotor; L sm and L sσ are the excitation inductance and the leakage inductance of the motor, respectively. By adopting fractional slot, i.e., deep and narrow slot, as shown in Fig., the leakage

3 734 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 6, NO., DECEMBER 05 inductance is increased, and the short-circuit current will be reduced and limited. In addition, the DFPM motor has a big overload ratio, i.e., if the rated torque of the DFPM motor is T N, the maximum torque capability of the DFPM motor are T max and T max under normal and fault conditions, which can be expressed as follows: T max =3T N (3) T max = 3T N. (4) C. Optimization of Cogging Torque and Back EMF The traditional fault-tolerant PM motor consists of surfacemounted PM rotor and stator with concentrated armature windings on alternate teeth. Hence, the fault-tolerant PM motor can not only offer the advantages of high power density, high efficiency, small volume, and low torque ripple, which are the same as the general PM motor, but also have the characteristics of electrical isolation, magnetic isolation, thermal and physical isolation, and inhibition of short-circuit current. However, the existing fault-tolerant PM motors suffer from the drawback of nonsinusoidal no-load back EMF. Meanwhile, it has cogging torque ripple and the general PM motor, and it will affect the output performance of the motor [5], [0], []. In order to solve the aforementioned problems, the general optimization methods, namely, chutes, changing magnetic steel pole arc coefficient, and changing notch s size, have been adopted. However, it is difficult to adopt one method to reduce the harmonic component of no-load back EMF and the cogging torque ripple, simultaneously [5]. In this paper, the back-emf and cogging torque of the proposed DFPM motor shown in Fig. are optimized by changing the PM shape. To optimize the proposed motor, some parameters are defined as follows. The centrifugal height h is defined as the distance between the center of the rotor magnet steel inner arc circle and the center of the rotor magnet steel outer arc circle. The parameters o and R n are the center and the radius of the rotor magnet steel inner circular arc, respectively. In addition, o and R w are the center and the radius of the rotor magnet steel outer circular arc, respectively. When the radius of the rotor magnet steel inner circular arc R n is unchanged, the cogging torque and back-emf waveforms of the proposed DFPM motor at different centrifugal height h are shown in Fig. 3(a) and (b), respectively. It can be seen that the radial flux density of air gap is increasingly more sinusoidal with the increasing of the centrifugal height h, which will reduce the impact of the notch on the whole radial flux density of the proposed motor, simultaneously. Hence, by changing the centrifugal height h, the harmonic components of the noload back-emf and the cogging torque ripple of the proposed motor will be reduced simultaneously. It should be mentioned that, with the increase of the centrifugal height h, the leakage coefficient of the PM will be increased. Hence, the centrifugal height h =0mm is chosen to be the optimal height. Compared with the conventional fault-tolerant PM motor whose centrifugal height h =0 mm [8], [], the cogging torque ripple and radial flux density in the air gap of the proposed Fig. 3. Cogging torque and back-emf waveforms of the proposed DFPM motor at different centrifugal height. (a) Cogging torque. (b) Flux density. Fig. 4. FEM and measured no-load back-emf waveforms. DFPM motor have been reduced and improved by the mean of the centrifugal height optimization. To verify the theoretical analysis, the finite-element method (FEM) and the experimental results of the no-load back-emf of the proposed DFPM motor are shown in Fig. 4. It can be seen that the waveform of experimental no-load back EMF is very sinusoidal and agrees well with the FEM result. The key parameters of the proposed DFPM motor are listed in Table I. III. DFPM MOTOR SYSTEM MODELING It is shown in Figs. and that the two sets of independent three-phase concentrated armature windings wound

4 JIANG et al.: ELECTRIC DRIVE SYSTEM OF DFPM MOTOR FOR AEROSPACE APPLICATIONS 735 TABLE I KEY PARAMETERS OF THE PROPOSED DFPM MOTOR alternately around the stator teeth. Moreover, there are 60 electrical degrees apart between the abc windings and xyz windings. Hence, each phase winding of the proposed motor is independent, and the mutual inductance between adjacent phase windings is very small. Hence, the proposed motor has a strong fault tolerance. In addition, the mutual inductance of the proposed DFPM motor can be ignored, which will simplify the mathematical model of the proposed DFPM motor. The phase voltage and flux linkage of the abc and the xyz windings of the proposed DFPM motor in the stationary reference frame can be expressed as follows: [ Uabc ψ abc [ Uxyz ] = ψ xyz ] [ R = s pψ abc L ψ fabc ][ iabc [ ][ R s pψ xyz ixyz L ψ fxyz where U abc =[U a U b U c ] T and ψ abc =[ψ a ψ b ψ c ] T are the phase voltage vector and the stator flux vector of the abc windings, respectively. i abc =[i a i b i c ] T is the phase current vector of the abc windings. R s = diag[r s R s R s ] T is the stator resistance vector. p is the differential operator. U xyz = [U x U y U z ] T is the phase voltage vector of the xyz windings. ψ xyz =[ψ x ψ y ψ z ] T is the stator flux vector of the xyz windings. i xyz =[i x i y i z ] T is the phase current vector of the xyz windings. L and L are the stator inductance vectors of the abc windings and the xyz windings, respectively. ψ fabc and ψ fxyz are the PM flux vectors of the abc windings and the xyz windings, respectively. To get the two-phase d q-axis electromagnetic parameters, the traditional Park matrices are used in this paper as follows: T dq = 3 T dq = 3 cos θ sin θ cos(θ 0 ) sin(θ 0 ) cos(θ + 0 ) sin(θ + 0 ) cos(θ 60 ) sin(θ 60 ) cos(θ 80 ) sin(θ 80 ) cos(θ +60 ) sin(θ +60 ) ] ] (5) (6) T (7) T. (8) Fig. 5. Control block diagram of the proposed DFPM motor. The voltage vector of d q-axis can be described in the synchronous rotating reference frame as follows: [ ] [ ] Vd (Rs + pl = d ) ω e L q pψ f i d i V q ω e L d (R s + pl q ) ω e ψ q (9) f [ ] [ ] Vd (Rs + pl = d ) ω e L q pψ f i d i V q ω e L d (R s + pl q ) ω e ψ q f (0) where ω e is the electrical angular velocity of the rotor, i.e., ω e = nω r /, ω r is the mechanical angular velocity of the rotor; n is the number of pole pairs of the DFPM motor; V d and V q are the d q-axis voltages of the abc windings; V d and V q are the d q-axis voltages of the xyz windings. The electromagnetic torque T e can be represented by the following equation: T e = 3 n ([ψ f i q +(L d L q ) i d i q ] +[ψ f i q +(L d L q )i d i q ]). () Because the PM of the DFPM motor is surface mounted, L d L q,l d L q, T e can be simplified as follows: T e = 3 n ψ f (i q + i q )= 3 n ψ f i q. () The motion equation can be expressed as follows: J dω r = T e T L Bω r (3) dt where T L is the load torque and its unit is N m, J is the inertia of the DFPM motor and its unit is kg m,andb is the damping coefficient. IV. CONTROL STRATEGY AND FAULT DIAGNOSIS A. Fault-Tolerant Control Strategy of the System The control block diagram of the proposed DFPM motor is shown in Fig. 5, in which the active active redundant faulttolerant control strategy based on the failure diagnosis and redundancy communication and the two sets of three-phase full-bridge drive circuits, namely, inverter and inverter, are adopted. In addition, the two sets of independent redundancy control strategy consist of the speed controller, the current

5 736 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 6, NO., DECEMBER 05 TABLE II FAULT DIAGNOSIS METHODS Fig. 6. Comparison of fault-tolerant topologies. controller, space vector pulsewidth modulation (SVPWM), abc/dq conversion, and the inverter. Furthermore, the electric drive system of the proposed DFPM motor has some merits as follows: It can offer higher overall kva/kw ratio for constant power output need, stronger power output capacity, and less power switches. Because the high reliability is a critical requirement for an electrical drive aircraft fuel pump system, the fault-tolerant topology with higher overall kva/kw ratio can offer stronger power output capacity, bigger overload ratio, and better radiating effect [3] [5]. Fig. 6 shows the comparison of the number of power switches and overall kva/kw ratio of two basic fault-tolerant topologies for multiple single-phase (n +) systems and multiple three-phase (3n +3)systems. For example, the six single-phase (5 + ) systems, as shown in Fig. 6, express that six single-phase systems with six H-bridge drive circuits adopt five single-phase systems operated under one single-phase fault condition. The two three-phase (3 + 3) systems, as shown in Fig. 6, namely, the electric drive system of the proposed DFPM motor, express that two three-phase systems with two sets of three-phase full-bridge drive circuits adopt the other three-phase operated under one three-phase fault condition. It is shown in Fig. 6 that the overall size of the fault-tolerant drive system tends to decrease with the increasing of the modules number n, but the number of power switches will increase. However, as the number of power switches increases, the drive systems will suffer from some drawbacks of lower power density, poor reliability, and complex control algorithms [4], [5], []. The electric drive system of the proposed DFPM motor (3 + 3) has stronger power output capacity, bigger overload ratio, and better radiating effect by using less power switches. It can offer fault-tolerant arrangement with the minimum number of power supplies, processors, filters, and simply control algorithms. In addition, they act as a balanced load on the supplies, reducing electromagnetic interference and total harmonic distortion. Each redundancy provides smooth torque; in general, loss of one redundancy does not bring in ripple or drag torque into the drive system: compensation for one fault redundancy simply by increasing the torque demand for the other normal redundancy. In addition, the proposed electric drive system has the functions of fault diagnosis and redundant communication between the two redundancies. At the same time, the active active redundant control strategy based on the i d =0vector control method was adopted in the proposed drive system. Hence, when the electric drive system of the proposed DFPM motor operated normally, each redundancy of the drive system worked simultaneously and outputted 50% power, respectively. In addition, the opencircuit or short-circuit fault can be detected, and the fault set of windings of the DFPM motor will be removed from the proposed system at once. Then, the fault signal will be sent to the normal redundancy by using the redundant communication function. Finally, the output power can be kept at 00% of the normal condition by changing the control state of the normal redundancy. B. Fault Diagnosis Motors and power converters have the highest failure rate in the aerospace control system [5]. The two main faults among the highest failure rate per flight hour are open-circuit fault and short-circuit fault in phase winding. Hence, the opencircuit fault and short-circuit fault in the phase winding of the proposed motor will be investigated in this paper. The fault diagnosis methods under open-circuit fault and short-circuit fault conditions are shown in Table II. V. S YSTEM SIMULATION VERIFICATION The control system model of the electric drive system of the proposed DFPM motor will be built by using MATLAB/ Simulink. The simulation model mainly contains the inverter, the DFPM motor, the speed PI controller, the current PI controller, and SVPWM module. The dc-link voltage is 70 V. Based on the control system of the simulation model and the proposed control strategy and fault diagnosis, the operational performance of the proposed DFPM motor under open-circuit and short-circuit fault conditions will be investigated here.

6 JIANG et al.: ELECTRIC DRIVE SYSTEM OF DFPM MOTOR FOR AEROSPACE APPLICATIONS 737 current of each phase is 6 A before 0.0 s. After the opencircuit fault occurs in the phase-a winding, the phase current of a, b, andc windings reduced to 0 A, while the phase current of x, y, andz windings are twice as the original one. Its peak value of current in each phase was 5 A, and it provided 00% power to ensure that the output power of the proposed electric drive system keeps constant. Fig. 7(c) shows the corresponding torque waveforms of xyz windings when the open-circuit fault occurs in phase-a winding. It can be seen that the output torque generated by xyz windings was 0 N m before 0.0 s, which is only half of the rated load. After the open-circuit fault occurs in the phase-a winding, the normal phase x, y,andz windings will output the whole rated load power, and the output torque was 0 N m. Fig. 7(d) and (e) shows the output torque and speed waveforms of the DFPM motor at pre- and post-opencircuit fault condition. It can be seen that the output torque and speed are all kept constant, which verify the proposed faulttolerant control strategy when the proposed electric drive system of the DFPM motor operates at open-circuit fault condition. Fig. 7. Simulation waveforms of the DFPM motor at pre- and postopen-circuit fault condition. (a) Current waveforms of x, y, and z windings. (b) Current waveforms of a, b, and c windings. (c) Torque waveforms of xyz windings. (d) Torque waveforms of the DFPM motor. (e) Speed waveforms of the DFPM motor. A. Simulation of Open-Circuit Fault Fig. 7(a) (e) shows the simulated results of the open-circuit fault in the phase-a winding, when the DFPM motor operated at the rated load is 0 N m and the speed is 000 r/min. When the open-circuit fault occurs in the phase-a winding at 0.0 s, the current waveforms of x, y, andz windings and the current waveforms of a, b, andc windings are shown in Fig. 7(a) and (b), respectively. It is shown in Fig. 7(a) and (b) that each set of windings provided 50% power and the peak B. Simulation of Short-Circuit Fault Fig. 8(a) (e) shows the simulated results of the short-circuit fault in phase-a winding, when the DFPM motor operated at the load is 5 N m and the speed is 000 r/min. When the shortcircuit fault occurs in the phase-a winding at 0.5 s, the current waveforms of x, y, andz windings and the current waveforms of a, b, andc windings are shown in Fig. 8(a) and (b), respectively. It is shown in Fig. 8(a) and (b) that the peak value of short-circuit current of a, b, andc fault windings was nearly 50 A, when the system comes to steady state. Because 50 A was close to the limited value of the DFPM motor s shortcircuit current, it can be seen that the system had the function of inhibiting the short-circuit current. Meanwhile, the peak current of x, y,andz windings also increased greatly and stable at 5 A. Fig. 8(c) shows the torque waveforms of xyz windings, when the short-circuit fault occurs in phase-a winding. It can be seen that the output torque of xyz windings was.5 N m before 0.5 s, which is only half of the load. After the short-circuit fault occurs in the phase-a winding, the normal phase x, y, andz windings will compensate for the absent torque of the removed fault phases and offset the pulsating torque of the short-circuit phases. Fig. 8(d) and (e) shows the output torque and speed waveforms of the DFPM motor at pre- and post-short-circuit fault condition. It can be seen that the proposed electric drive system can still operate steadily after a short pulsation, which verifies the proposed fault-tolerant control strategy when the proposed electric drive system of the DFPM motor operates at short-circuit fault condition. VI. EXPERIMENTAL RESULTS A. Experimental Platform To verify the proposed control strategy and fault diagnosis, the experiments have been carried out on the test platform of the electric drive system of the proposed DFPM motor, as shown in Fig. 9. Its main experimental equipment included the DFPM motor, two sets of controllers, torque and speed sensors,

7 738 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 6, NO., DECEMBER 05 Fig. 9. Test platform of the electric drive system of the DFPM motor. Fig. 8. Simulation waveforms of the DFPM motor at pre- and postshort-circuit fault condition. (a) Current waveforms of x, y, and z windings. (b) Current waveforms of a, b, and c windings. (c) Torque waveforms of xyz windings. (d) Torque waveforms of the DFPM motor. (e) Speed waveforms of the DFPM motor. hysteresis dynamometer, voltage regulator, 8-V control power supply, oscilloscope, industrial computer, and load controller. A TMS30F8 digital signal processor of Texas Instruments was adopted, and its basic frequency is 50 MHz. The logic processing chip adopted M4A5-9/96 of Lattice, and its basic frequency is 0 MHz. B. Experiment of Open-Circuit Fault Fig. 0 show the experimental results of the system under phase-a open-circuit fault condition, when the DFPM motor operated at the rated load is 0 N m and the speed is 000 r/min. Fig. 0. Measured waveforms under phase-a open-circuit fault condition with the rated load of 0 N m at the speed of 000 r/min. (a) Measured waveforms before phase-a open-circuit fault. (b) Measured waveforms at pre- and post-phase-a open-circuit fault. (c) Measured waveforms in phase-a open-circuit fault steady state. When the open-circuit fault occurs in the phase-a winding at the rated load of 0 N m and the speed of 000 r/min, the current waveforms of the phase-a and phase-x winding, the torque waveforms outputted by xyz windings, and the speed waveforms of the DFPM motor are shown in Fig. 0(a) (c),

8 JIANG et al.: ELECTRIC DRIVE SYSTEM OF DFPM MOTOR FOR AEROSPACE APPLICATIONS 739 Fig.. Measured waveforms before the phase-a short-circuit fault. respectively. It is shown in Fig. 0(a) that each set of windings provided 50% power and the peak current of the phase-x and phase-a winding are all 6 A before the open-circuit fault occurs in phase A. Fig. 0(b) shows that the value of the current in the fault phase-a winding is 0 A after the open-circuit fault occurs in phase A. In addition, the peak value of current in the healthy phase-x winding stabilized at 5 A after 600 ms, and it is twice as the original peak current. At this moment, the torque produced by xyz windings increases from 0 to 0 N m, namely, which provides 00% of power. At the same time, the speed of the DFPM motor keeps constant at 000 r/min. It is shown in Fig. 0(b) that the speed decreases slightly and the torque produced by xyz windings increases transiently, when the fault occurs. Fig. 0(b) and (c) shows the output torque and the motor speed kept constant in phase-a open-circuit fault steady state. It can be seen that the experimental results agree well with the simulated results. Therefore, it can be concluded that the proposed electric drive system can still operate steadily after a short pulsation, which verifies the proposed fault-tolerant control strategy when the proposed electric drive system of the DFPM motor operates at open-circuit fault condition. C. Experiment of Short-Circuit Fault Fig. shows the measured waveforms of current, torque, and speed of the proposed DFPM before phase-a winding short-circuit fault, when the load is 5 N m and the given speed is 000 r/min. It can be seen that each set of windings offers 50% of the total output power. The peak value of current in the phase-x and phase-a winding are the same and equal to 3 A. Moreover, the torque of abc windings keeps constant at.5 N m, and the speed of the DFPM motor keeps constant at 000 r/min. Fig. shows the waveforms of current in the phase X and phase A, the output torque, and the speed of the DFPM motor after the short-circuit fault occurs in phase A. After the short-circuit fault occurs in phase A, it can be observed that the peak value of short-circuit current of phase-a winding is about 45 A, when the system comes to steady state. Because 45 A was close to the limited value of the experimental DFPM motor s short-circuit current, it can be seen that the system had the function of inhibiting the short-circuit current. Compared with the state without short-circuit fault, the peak current of phase-x winding with short-circuit fault has a great increase and stabilizes at 5 A. This was to compensate the absent torque of the removed fault phases and offset the pulsating torque of the short-circuit phases. When the electric drive system of the Fig.. Measured waveforms in phase-a short-circuit fault steady state. proposed DFPM motor is in steady state after the short-circuit fault in phase-a winding, it is shown in Fig. that the drive system can still operate steadily. Hence, it can be concluded that the output performance of the DFPM motor remained nearly the same at pre- and post-short-circuit fault condition, which verifies the proposed fault-tolerant control strategy when the proposed electric drive system of the DFPM motor operates at short-circuit fault condition. These experimental results agree well with the simulated results. Therefore, it can be concluded that the high performance of fault tolerance is achieved in the electric drive system of the proposed DFPM motor. VII. CONCLUSION In this paper, a new fault-tolerant electric drive system based on a six-phase ten-pole DFPM motor for aerospace applications has been proposed and investigated. To reduce the cogging torque and harmonic components of the proposed DFPM motor, a new method by optimizing the PM shape has been adopted and verified by means of FEM and experiments. To further improve the fault-tolerant performance, the limited value of the short-circuit current and the big overload ratio have been designed. Then, to investigate the fault-tolerance performance of the proposed electric drive system, the mathematical model of the proposed DFPM motor in d q axis has been derived. Furthermore, the fault-tolerant control strategy and the failure diagnosis strategy of the drive system have been proposed and introduced. Moreover, the operational performance of the drive system under open-circuit and short-circuit fault has been verified by simulation and experiments. The simulation and experimental results show good fault-tolerant characteristics of the proposed electric drive system, i.e., the output torque and speed of the proposed electric drive system can keep constant when open-circuit and short-circuit fault occurs in the phase winding. The proposed fault-tolerant electric drive system and control strategy can be expected to have a bright future in high-reliability and high-power-density applications such as aerospace. REFERENCES [] D. P. Rubertus, L. D. Hunter, and G. J. Cecere, Electromechanical actuation technology for the all-electric aircraft, IEEE Trans. Aerosp. Electron. Syst., vol. AES-0, no. 3, pp , May 984. [] R. I. Jones, The more electric aircraft: The past and the future? in Proc. Inst. Elect. Eng. Colloq. Elect. Mach. Syst. More Elect. Aircraft, 999, pp. / /4.

9 7330 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 6, NO., DECEMBER 05 [3] J. W. Bennett, G. J. Atkinson, B. C. Mecrow, and D. J. Atkinson, Faulttolerant design considerations and control strategies for aerospace drives, IEEE Trans. Ind. Electron., vol. 59, no. 5, pp , May 0. [4] G. J. Atkinson et al., Fault tolerant drives for aerospace applications, in Proc. 6th Int. Conf. Integr. Power Elect. Syst., 00, pp. 7. [5] W. Cao, B. C. Mecrow, G. J. Atkinson, J. W. Bennett, and D. J. Atkinson, Overview of electric motor technologies used for more electric aircraft, IEEE Trans. Ind. Electron., vol. 59, no. 9, pp , Sep. 0. [6] E. D. Ganev, High-performance electric drives for aerospace more electric architectures Part I Electric machines, in Proc. IEEE Power Eng. Soc. Gen. Meet., 007, pp. 8. [7] J. A. Weimer, The role of electric machines and drives in the more electric aircraft, in Proc. IEEE IEMDC, 003, vol., pp. 5. [8] B. C. 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Haylock, A comparative study of permanent magnet and switched reluctance motors for high-performance fault-tolerant applications, IEEE Trans. Ind. Appl., vol. 3, no. 4, pp , Jul./Aug [4] R. Cao, M. Cheng, C. Mi, and W. Hua, Influence of leading design parameters on the force performance of a complementary and modular linear flux-switching permanent magnet motor, IEEE Trans. Ind. Electron., vol. 6, no. 5, pp , May 04. [5] Z. Hao, Y. Hu, and W. Huang, The research on key technologies of faulttolerant permanent magnet motor in the drive system of electric vehicle, in Proc. IEEE VPPC, 008, pp. 4. [6] K. Atallah, J. B. Wang, and D. Howe, Torque-ripple minimization in modular permanent-magnet brushless machines, IEEE Trans. Ind. Appl., vol. 39, no. 6, pp , Nov./Dec [7] A. Stabile, J. O. Estima, C. Boccaletti, and A. J. Marques Cardoso, Converter power loss analysis in a fault-tolerant permanent-magnet synchronous motor drive, IEEE Trans. Ind. Electron., vol. 6, no. 3, pp , Mar. 05. [8] B. C. Mecrow, A. G. Jack, and J. A. Haylock, Fault tolerant permanent magnet machine drives, Proc. Inst. Elect. Eng. Elect. Power Appl., vol. 43, no. 6, pp , Nov [9] J. A. Haylock, B. C. Mecrow, and A. G. Jack, Enhanced current control of high-speed PM machine drives through the use of flux controllers, IEEE Trans. Ind. Appl., vol. 35, no. 5, pp , Sep./Oct [0] D. E. Jason, K. Atallah, J. B. Wang, and D. Howe, Effect of optimal torque control on rotor loss of fault-tolerant permanent magnet brushless machines, IEEE Trans. Magn., vol. 38, no. 5, pp. 3 30, Sep. 00. [] J. B. Wang, K. Atallah, and D. Howe, Optimal torque control of faulttolerant permanent magnet brushless machines, IEEE Trans. Magn., vol. 39, no. 5, pp , Sep [] W. Zhao et al., Stator-flux-oriented fault-tolerant control of fluxswitching permanent-magnet motors, IEEE Trans. Magn., vol. 47, no. 0, pp , Oct. 0. [3] W. Zhao, M. Cheng, K. T. Chau, R. Cao, and J. Ji, Remedial injectedharmonic-current operation of redundant flux-switching permanentmagnet motor drives, IEEE Trans. Ind. Electron., vol. 60, no., pp. 5 59, Jan. 03. [4] A. Mohammadpour, S. Sadeghi, and L. Parsa, A generalized faulttolerant control strategy for five-phase PM motor drives considering star, pentagon, and pentacle connections of stator windings, IEEE Trans. Ind. Electron., vol. 6, no., pp , Jan. 04. Xuefeng Jiang (S 5) was born in Chongqing, China, in 987. He received the B.S. degree in electrical engineering from Southwest Jiaotong University, Chengdu, China, in 0. He is currently working toward the Ph.D. degree in the Department of Electrical Engineering, Nanjing University of Aeronautics and Astronautics, Nanjing, China. His main research interests include control and optimal design for electric drive systems of fault-tolerant permanent-magnet motors, fault analysis, and fault diagnosis. Wenxin Huang (M 09) was born in Dongtai, China, in 966. He received the B.S. degree from Southeast University, Nanjing, China, in 988 and the M.S. and Ph.D. degrees from Nanjing University of Aeronautics and Astronautics (NUAA), Nanjing, China, in 994 and 00, respectively. In 003, he joined the Faculty of the College of Automation Engineering, NUAA, where he is currently a Professor. His research interests include stand-alone power systems, power electronics, and design and control for electrical machine systems. Ruiwu Cao (S 0 M 3) received the B.S. degree from Yancheng Institute of Technology, Yancheng, China, in 004 and the M.S. and Ph.D. degrees in electrical engineering from Southeast University, Nanjing, China, in 007 and 03, respectively. From August 00 to November 0, he was a joint Ph.D. student with the College of Electrical and Computer Science, University of Michigan, Dearborn, MI, USA, funded by the China Scholarship Council. Since 03, he has been with Nanjing University of Aeronautics and Astronautics, Nanjing, China, where he is a Lecturer in the Department of Electrical Engineering. His teaching and research interests include linear motors for rail transit and electromagnetic launch systems, motor drives for electric vehicles, and renewable energy generation. Zhenyang Hao received the B.S. degree from Nanjing Normal University, Nanjing, China, in 004 and the Ph.D. degree in electrical engineering from Nanjing University of Aeronautics and Astronautics, Nanjing, China, in 00. In 0, he joined the Faculty of the College of Automation Engineering, Nanjing University of Aeronautics and Astronautics. His research interests include design and control for electrical machine systems. Wen Jiang receivedtheb.s.degreeinelectrical engineering in 04 from Nanjing University of Aeronautics and Astronautics, Nanjing, China, where he is currently working toward the M.S. degree in electrical engineering. His research interests include fault-tolerant permanent-magnet motors and their fault diagnosis.

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