Fault Tolerant In-Wheel Motor Topologies for High ... - Semantic Scholar

2011 IEEE International Electric Machines & Drives Conference (IEMDC)

Fault Tolerant In-Wheel Motor Topologies for High Performance Electric Vehicles. C.J. Ifedi, B.C. Mecrow, S.T.M. Brockway, G.S. Boast, G.J. Atkinson and D. Kostic-Perovic Abstract – The use of in-wheel motors, often referred to as hub motors, as a source of propulsion for pure electric or hybrid electric vehicles has recently received a lot of attention. Since the motor is housed in the limited space within the wheel rim, it must have a high torque density and efficiency, and survive the rigours of being in-wheel in terms of environmental cycling, ingress, shock and vibration and driver abuse. Finally, to ensure adequate levels of functional safety are met it is essential that failures do not lead to loss of control of the vehicle. This paper presents studies of a fault tolerant concept for the design of inwheel motors. The study focused on achieving a high torque density and the ability to sustain an adequate level of performance following a failure. A series of failures are simulated and then compared with experimental tests on a demonstrator motor. Keywords- drive train; electric vehicles (EVs); electric propulsion; Fault tolerance; sub-motor; drag torque, in-wheel, hub motor.

I.

INTRODUCTION.

The drive to reduce CO2 emissions through alternative energy sources has been the most talked about subject in recent times. Hence the need for vehicles to be powered by any energy source other than petroleum has received huge interest. Worldwide, researchers are working to develop more efficient drive systems for vehicles, in particular using electric motors. With numerous different vehicle applications and requirements, it is clear that no single motor design fits all. Motors designed for electric vehicle applications have to meet rigorous demands, with space limitations and the driving environment key factors. The following are typical design requirements for selection and design of the drive motor [1, 2, 3]. • • • • • • •

High instantaneous torque and continuous torque density. High torque at low speed for starting and climbing. Fast torque response for use in braking (ABS) Low cogging torque and torque harmonics for refined drivability. High efficiency over constant torque and power range High fault tolerance. Good overload capability for uphill climbing.

Due to the high torque density and high efficiency of the permanent magnet motor, there is an ever growing interest in the application of these motors to the propulsion of electric vehicles. By employing concentrated, single tooth windings and fractional numbers of slots/pole/phase [4], the torque density can be maximised, with high slot fill factors and short endwindings.

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Mounting the EV drive train in the wheel of the vehicle has received significant interest due to the following advantages: • • • • •

It releases vehicle space for passengers cabin, traction batteries etc. The torque at each wheel can be independently controlled, giving true 4 wheel drive. True wheel-slip based control allows improved vehicle response and shorter braking distances. Removal of drivetrain components gives vehicle designers greater freedom to design vibration free, more comfortable vehicles Elimination of gears and differentials, producing a simpler and more reliable overall mechanical design [5].

Problems often perceived with in-wheel motors are related to ride and handling, due to an increase in the vehicle un-sprung mass, and safety issues when a motor fault occurs, developing a dangerous torque disturbance. In [6, 7, 8] investigations show the un-sprung mass isn’t actually a serious issue with wheel mounted motors. This paper addresses the safety issues associated with inwheel motors and presents a fault tolerant concept for a direct drive in-wheel motor, with integrated drive electronics for use in passenger and light commercial vehicles. The motor is made up of series of three phase sub-motors, with each submotor operating independent of the other sub-motors, so that a fault on any one sub-motor introduces only a fraction of the torque disturbance that would occur in a full motor. The machine has also been designed with a relatively large phase inductance, which limits the fault current during a terminal short circuit. Since the motor is a direct drive wheel motor with no gearing, and requires a high power density and high efficiency [9], the motor has to be designed as a high torque, low speed machine. With no gearing, the size and weight of the motor is larger than a geared motor (excluding gears), as the motor has to provide full torque capability at zero and low speed as required for acceleration from rest and uphill climbing. However with a careful motor design the size and weight can be minimised. II. WHEEL MOTORS. A. In-wheel Motor Designs For EV Previous research has been carried out to develop different design concepts for high-torque, low speed motors, suitable for direct drive in-wheel applications. However most publications present work on motor designs used for hybrid

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vehicle applications, with few publications on in-wheel motor designs for pure electric vehicle application. In [10-14] different types of in-wheel motors designed for the EV application are presented. B. Drive Specification and Topology Table 1 shows the specification for the motor drive used in this work, revealing a high demand, torque dense arrangement. The topology chosen is that of a 3-phase, radial field, permanent magnet, outer rotor, in-wheel motor, with each inwheel motor split into eight sub motors: figure 1 shows a cutaway cross section of the in-wheel motor. The motor is part of the Protean Electric in wheel propulsion system, with the power electronic drive and motor as one packaged unit. A series of these motors have been built and extensively tested in a variety of vehicles, ranging from a Volvo C30 to a Ford F150 SUV. In both of these vehicles the motors were placed in all four wheels, giving true four wheel drive. Fig. 1: In-wheel drive train

TABLE 1: WHEEL MOTOR DRIVE SPECIFICATION.

Peak Output Power

80 kW

Continuous Output Power

54 kW

Peak Output Torque

800 Nm

Continuous Output Torque

475 Nm

DC Supply Voltage

380V

Overall width

115 mm

Overall diameter

420 mm

Maximum speed

1400 r/min

System mass (inverter + motor)

31 kg

III. THE PROTEAN WHEEL MOTOR. The outer rotor design of the motor maximises the air-gap diameter and thereby minimises the shear force required at the air-gap for a given torque. Since the motor is direct drive, the maximum speed requirement of the motor is modest, hence it is possible to have a large number of poles, resulting in very shallow core backs and creating a “ring” type structure. The centre of the ring is used to house the power electronics. Windings are concentrated around individual teeth to minimise end-winding length: this simple winding structure also enables a high slot fill factor. The stator is water cooled, enabling high winding current density and a consequently high magnetic loading. The result is a very torque dense drive system, producing 26Nm/kg of total mass.

Fig. 2: Loaded motor flux distribution

In order to improve the two dimensional simulation, a single three dimensional model was constructed to determine the end-winding inductance of the machine. Figure 4 is a section of the motor, showing the model used for the calculation of the end-winding inductance. The endwinding inductance was then incorporated into the two-dimensional finite element model as an external impedance in series with the windings.

A finite element model of the machine was developed and validated against measurement. This model is two dimensional in nature and therefore does not accommodate fringing and leakage at the axial ends of the machine. A time stepping method is used so that harmonic fields are represented and variations in torque with position can be determined. Figure 2 shows a finite element plot of the machine under load, and figure 3 compares the variation of mean torque with current, whilst operating at rated speed. Two dimensional modelling tends to slightly overestimate the performance by up to 10%. This is primarilly a result of the failure to model end effects in a relatively short machine. At high currents there is a modest element of saturation, reducing the torque constant slightly. Fig. 3: Motor Torque Performance versus Electric Loading

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Fig. 4: 3D Element Size.

Fig. 5: Predicted Cogging torque of Prototype Motor.

The machine back emf has low harmonic content, primarilly due to the nature of the winding, and so torque ripple remains low when the motor is supplied with sinusoidal current. Cogging torque results from interaction between the magnet and the reluctance change with position of the airgap due to the slotting effect of the stator. It was found that for this machine the predicted cogging torque is exceptionally low (0.025% of rated torque), as illustrated in figure 5. In practice, manufacturing tolerances will significantly increase this value, but it still remains remarkably low in built prototypes. Plots of the phase back EMF and line back EMF are shown in figure 6 at a speed of 1200rpm. IV. RELIABLITY AND SAFETY REQUIREMENT. Reliability requirements are in line with those for conventional commercial internal combustion engine driven vehicles. The reliability of conventional industrial drives is of the order of 105 hours between failures, corresponding to 3,000,000 miles at 30 miles per hour. Clearly this is well beyond the life of most vehicles and offers the promise of increased reliability over the internal combustion engine. However, the environment of an in-wheel motor is far from benign: the motor is sealed against water and dust ingress and designed to withstand shock loads. Having independent inwheel motors gives a level of redundancy, so that there is also a “limp home” capability in the event of any one motor drive failure. Vehicle safety must be considered in the event of catastrophic failure of one motor drive: this is of most concern when travelling at high speed. System studies have revealed that a very large disturbance torque of 280Nm or more suddenly appearing at one wheel could cause loss of vehicle control, and so the system must be designed to virtually eliminate such a possibility. The next section discusses how this is achieved, using concepts similar to those adopted for permanent magnet fault tolerant aerospace drives [15- 18].

Fig. 6: Phase and Line Back EMF at 1200rpm. Bearing loadings are virtually identical to those in conventional cars: bearing failures are managed successfully in mass produced vehicles and are therefore not considered further here. Open-circuit faults result in high torque ripple and loss of torque, but do not introduce disturbance torque, hence they are relatively benign. Short-circuits can result in large circulating currents, irrespective of whether they are caused by winding, power device, capacitor of power supply failure. These currents continue to flow, even if the power supply is removed, because they continue to be driven by the rotor magnets. Under certain circumstances these circulating currents can result in large disturbance torques and so the drive system has two features to prevent this. (i)

V. REDUNDANCY AND FAULT CURRENT LIMITATION. Potential failures in the drive system include: a) Winding open-circuit b) Motor winding short-circuit to earth, winding terminal short-circuit, or internal turn-turn shortcircuit. c) Power device failure – either open or short-circuit d) Capacitor failure – open or short circuit. e) Dc. bus open or short-circuit f) Mechanical bearing failure

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Each motor is split into a series of sub-motors and each sub-motor is a three phase motor in its own right, with its own three phase inverter and windings, which operate independently of the other sub-motors. The motor of figure 1 has eight sub-motors, with some of the power converter modules visible on the side of the motor. Failure of any one sub-motor only introduces one eighth of the braking torque of a full motor, whilst the other sub-motors can continue to produce motoring torque. Where there are four wheels driving the vehicle there are in effect 32 submotors in total, and so it is anticipated that the braking torque will be low.

(ii)

The sub-motors are designed to have h a significant per-unit inductance. It is this induuctance that limits the maximum fault current inn the motor, as summarised in the next section. Thhis inductance also allows for significant field weakeninng.

Figure 8 shows the finitee element predictions and the measured mean drag torque as a a function of speed, whilst figure 9 shows the variation of peak p fault current with speed.

VI. EXPERIMENTAL ARRANG GEMENTS The test rig set up is shown in figures 7 and 8. The test motor is driven from a dynamometer, withh an inline torque transducer. A toggle switch via a contactor is i used in order to implement a range of winding failures.

Fig. 8: Mean Drag Torqque versus Speed for a single phase short-circuit.

Fig. 7: Test Rig Setup and measuring devices.

VII. FAULTED PERFORMAN ANCE. In this section simulated and measured reesults for different fault scenarios are presented. All simulated results are derived from two dimensional finite element time stepping models, with the endwinding inductance includeed as a lumped parameter. The faults on the motor were appllied with the motor initially running open circuit (no-load) and thhe fault current and drag torque were recorded at different speeeds for the various fault scenarios.

Fig. 9: Fault current versus v Speed for a single phase short-circuit. Peak drag torque is 17Nm and occurs at a rotational speed of 80revs/min. This is an order of magnitude less than the permitted torque disturbance and a is likely to pass completely unobserved by the driver. Figuure 10 shows the measured fault current that flows for a shortedd phase at 1200rpm. The steady state current is limited by the phase inductance and is within the thermal limit of the motor.

Fig. 8: Schematic of test rig and shorting arrangement. a

A. Single Phase Fault. This fault is between the terminal of one phase and the star point of the sub-motor. It could conceivvably result from degradation of the winding insulation. The drag d torque for this fault scenario has a large degree of rippple, due to the unbalanced nature of the fault, but the frequenncy of the ripple is several hundred Herz or more and will not be b observed by the driver. The measurement bandwidth off drag torque is generally much less than the torque ripple freequency and so the ripple is not observed in the measurements.

Fig. 10: Measured Shorrt circuit current for a single phase short at 1200 1 rpm.

B. Three – Phase Fault. A three phase symmetrical fault is most likely to be caused by a failure in the power elecctronic converter resulting from shorting of the DC link due to either capacitor or device shortcircuit failure. This scenario prroduces a larger drag torque, as shown in figure 11, which is prredictably about three times that of a single phase short. Once more m peak drag torque occurs at

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low speed: the peak drag torque of 50Nm is less than one fifth of the permitted disturbance and does not pose a risk to vehicle safety. Figure 12 shows the variation of peak fault current with speed, which once more is within the thermal capability of the motor. Because the fault is symmetrical the drag torque has no significant oscillations, as shown in figure 13.

figure 14. This current is an order of magnitude greater than the thermal limit of the machine and will cause thermal overheating. The reason that such large currents are created is that the single turn has a very low inductance, so that the current is limited solely by the resistance of the turn.

Fig. 14: Simulated turn-turn short circuit at 1200rpm. Fig. 11: Drag torque versus Speed for a three phase symmetrical short-circuit.

Although the shorted turn carries a very large current it was found that the fault creates a very small braking torque of only 2Nm. This is because the net MMF of the fault remains very small. D.

Discussion

The above results firstly show that

Fig. 12: Short Circuit current versus Speed for a gthree phase symmetrical short-circuit.

The drag torque in this case can produce pure braking torque, with little oscillation, as there is a balanced fault. The graph of drag torque is shown for this fault scenario, at 80rpm, where maximum drag torque is obtained.

•

There are no faults which result in a serious torque disturbance. Indeed, it is unlikely that the driver will even notice the occurance of any of the above failures.

•

The finite element simulation provides a remarkably accurate prediction of the torque and current resulting from various fault scenarios.

•

The largest drag torque occurs with a symmetrical short-circuit of a complete sub-motor.

•

The largest fault current occurs with a single turn short-circuit, but the resulting drag torque is very low.

•

It is noticeable from different fault scenarios that the drag torque variation with speed is rather like an induction motor characteristic. Peak distubance torque occurs at low speed, where its effect is not dangerous to the vehicle control: at high speed the disturbance torque of a single sub-motor short circuit can be as low as 10Nm in the machine, and is therefore relatively benign: it may not even be noticed by the driver. It can also be seen that the short circuit current increases with increasing speed, but converges at a certain speed as the current is no longer dependent on the speed of the rotor.

The only source of excitation for the winding fault current is the magnet flux, so the short circuit current is given by the equation: ∧

Fig. 13: Drag torque as a function of time with a symmetrical three phase short-circuit occuring at 80 rpm.

C. Inter–Turn Fault. Turn-turn faults can result from degradation of the winding insulation. No measurements were undertaken for this condition, but a finite element simulation was undertaken. The fault creates a large current in the shorted turn, as shown in

jωN φ isc = R + jωL

(1)

The short circuit current is limited by the winding resistance at lower speed and rises as speed rises. At high speeds the short circuit current is limited by just the winding

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inductance and becomes independent of speed, which can be obtained by the short circuit current equation (1).

REFERENCES. [1]

The drag torque is derived from the power loss equation (P = isc2R) and is obtained from the equation below: [2]

2

Tdrag

∧ ⎞ ⎛ ⎜ jN φ ⎟ ωR = p⎜ ⎜ R + jωL ⎟⎟ ⎠ ⎝

(2)

Where p is the number of pole pairs, φ is the magnet flux, N is the number of turns, ω is the electrical angular speed, and L and R are the winding inductance and resistance of the shorted winding section.

[3] [4] [5]

[6]

VIII. PREVENTING FAULT PROPAGATION To prevent fault propagation it is imperative that a fault in one sub-motor has little or no impact on the performance of an adjoining sub-motor. This requires that the magnetic coupling between two adjacent sub-motors must be as low as possible. Figure 15 below shows the phase back EMF of a sub-motor, before and after a symmetrical short-circuit fault was applied to the sub-motor adjacent to it. The phase back EMF drops by 3.3%, with no significant change in phase, which indicates that the healthy sub-motors can continue to operate, even during a fault on an adjacent sub-motor.

[7] [8] [9] [10]

[11] [12] [13] [14]

[15]

[16]

Fig. 15: Measured Pre-fault and Post-fault Phase back EMF in an adjacent unfaulted sub-motor IX. CONCLUSION

[17]

This paper has demonstrated that a high performance inwheel motor can be produced with intriniscally safe characteristics. A demonstrator motor has been extensively simulated and tested with a number of different fault scenarios. It has been shown through both simulation and experiment that the peak torque disturbance following a fault is less than 20% of the value which is likely to cause control problems for the driver. This very low value is achieved through a combination of design for an appropriate inductance to limit fault currents and through dividing the motor into a series of eight independent sub-motors. Without this division the fault torque is predicted to be eight times higher i.e. 400 Nm, which is 40% greater than safety permits. Experiment has shown that the sub-motors are only very weakly coupled magnetically, so that failure of one sub-motor should not significantly affect the operation of others.

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[18]

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