Three-Phase Inductive Power Transfer System for Roadways

3370 IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS, VOL. 54, NO. 6, DECEMBER 2007

A Three-Phase Inductive Power Transfer Systemfor Roadway-Powered Vehicles

Grant A. Covic, Senior Member, IEEE, John T. Boys,Michael L. G. Kissin, Student Member, IEEE, and Howard G. Lu

Abstract—The development of a new three-phase bipolar in-ductive power transfer system that provides power across theentire width of a roadway surface for automatic guided vehiclesand people mover systems is described. A prototype system wasconstructed to verify the feasibility of the design for a number ofmoving loads (toy cars). Here, 40 A/phase is supplied at 38.4 kHzto a 13-m-long test track. Flat pickups are used on the underside ofeach vehicle to couple power from the track to the vehicle. Finiteelement modeling software was used to design the geometricalposition of the track cables and to predict the power output. Thisdesign resulted in a considerably wider power delivery zone thanpossible using a single-phase track layout and has been experimen-tally verified. Mutual coupling effects between the various trackphases require additional compensation to be added to ensurebalanced three-phase currents.

Index Terms—Electromagnetic coupling, electromagnetic in-duction, energy conversion, road vehicle electric propulsion.

I. INTRODUCTION

INDUCTIVE power transfer (IPT) systems have found ap-plication where energy transfer to moving vehicles without

mechanical contact is required [1], [2]. Such systems includeclean rooms, monorail transportation, automatic guided vehi-cles (AGVs), and battery charging. In many of these, the vehicleis designed such that power is transferred continuously while itmoves along a track or rail [3], [4], requiring good horizontaland vertical alignments between the power pickup and the trackto ensure continuous power delivery. In monorail applications,this requirement is easily met as the magnetic power pickupis placed on a bogie whose movement relative to the trackis naturally constrained; consequently, there exist numerouswell-developed commercial systems [1], [2]. Roadway vehicleapplications using IPT have been proposed for more thantwo decades [5]–[7], but to date there has been only limitedcommercial development due to the difficulties in transfer-ring sufficient power to a moving vehicle without imposingconstraints on vehicle movement. Commercial AGVs requireautomatic steering control to ensure the power pickup is prop-erly aligned with the track [8], and battery charging systemsin people movers have been successfully employed but eitherrequire user plug-in or position alignment systems to ensureadequate charging [9], [10]. Roadway vehicles driven by humanoperators cannot meet the tolerance demanded by present sys-

Manuscript received November 29, 2006; revised May 23, 2007.The authors are with the Department of Electrical and Computer Engi-

neering, The University of Auckland, Auckland 1142, New Zealand (e-mail:[email protected]).

Digital Object Identifier 10.1109/TIE.2007.904025

Fig. 1. Typical IPT system.

tems, and consequently, system performance is compromised.Either the vehicle pickups have to be oversized or multiplepickups must be placed underneath the vehicle to compensatefor the inevitable misalignments of the pickup relative to thetrack [11]. Alternative suggestions to overcome this probleminclude sequential excitation of short track segments requiringmultiple switched primary coils and vehicle sensing [12], [13].Normally, on-board batteries are also required to facilitatestarting, manage power fluctuations, and enable unconstrainedmovement [5]–[8], [12], [14].

In this paper, a multiphase IPT system that removes theaforementioned restrictions while minimizing each vehicle’son-board power pickup weight is proposed. In AGVs andpeople mover systems, such a system will allow vehicles topass, which at present is impossible without on-board energystorage. Other advantages include the following: 1) providingshort power boosts to battery-powered electric vehicles whereincreased power demand is necessary (such as climbing particu-larly steep slopes) and 2) continuous power transfer without on-board batteries or battery charging along roadways without theneed for complicated pickup structures or driving restrictions.

This paper begins by discussing the basic principles of IPT.Following this, a practical three-phase IPT resonant supplyis presented along with methods of controlling and balancingthe current in each phase. Finite element modeling (FEM)is then used to help determine the appropriate track spacingfor a prototype system, and results are compared with actualmeasurements on the track.

II. PRINCIPLES OF IPT

A typical IPT system is shown in Fig. 1. It is composed oftwo distinctly different electromagnetic systems as follows:1) a power supply takes (three-phase) power from a mainssupply and energizes an extended primary loop or track, and

0278-0046/$25.00 © 2007 IEEE

COVIC et al.: THREE-PHASE INDUCTIVE POWER TRANSFER SYSTEM FOR ROADWAY-POWERED VEHICLES 3371

2) pickup coils or secondaries along that track have powercoupled into them, which is then processed and used to drivemotors, lights, or other loads as required. As shown therefore,an IPT system is a magnetically coupled system, but one wherethe coupling coefficient is much lower than other familiarmagnetically coupled systems, such as induction motors ortransformers.

The power supply used drives a constant current in theprimary track inductance. Low coupling factors between thetrack and pickup controllers within IPT systems make powertransfer difficult and the cost per kilowatt high. To improve thepower transfer capability, it is therefore common practice to in-crease the operating frequency considerably above usual mainsfrequencies and to use resonance with either or both the primaryloop and the secondary pickup. Either fixed or variable fre-quency power supplies can be chosen; however, fixed frequencysupplies have the advantage that the pickup design is simplifiedgiven that the track frequency will not shift under all loadingconditions [1], [15], [16]. In both IPT and induction heatingapplications, LCL resonant topologies are commonly used asthe fixed-frequency current-sourced supply [17]–[19], using afull bridge with series output inductance followed by the par-allel tuned track. The tuning capacitor C is chosen to resonatewith the track inductance L at the operating frequency. If thesupply inductance is also chosen to match the track inductance,then the network transforms a constant voltage to a constanttrack current and converts load at unity power factor while alsofunctioning as a bandpass filter, reducing the generated noise.

The addition of resonance complicates the IPT system. Asstated, on the primary side, a resonant capacitor can provide anexcitation current to the track in the same way as a transformerhas a magnetizing current so that the power switches only needto supply the real power required by the load. At the pickup(s),resonance boosts the power output. The open-circuit voltageinduced in the pickup coil as a result of current I1 in the trackis given by

Voc = jωMI1 (1)

where M is the mutual inductance between the track andthe pickup coil (self-inductance L2). The short-circuit currentof this pickup coil is Voc limited by the impedance of thesecondary winding, i.e.,

Isc =MI1

L2. (2)

A measure of the uncompensated voltampere of the pickupis given as follows:

Su = VocIsc. (3)

If the secondary coil is tuned for resonance at the trackfrequency ω using an appropriate capacitor (C2, either in seriesor in parallel), the voltages and currents in both C2 and L2

increase by a nominal factor of Q2 and the power transferredfrom the track to the pickup coil is then expressed as [1]

P = SuQ2 = ωI21

M2

L2Q2. (4)

Fig. 2. Common pickup shapes. (a) E. (b) Flat-E. (c) Simple flat.

The output of the pickup is normally a regulated dc voltageusing a switched converter that causes the operating Q2 ofthe circuit to vary with load, within some maximum value(typically 10) because of limitations in control and deviceratings [1], [15]. Thus, for a given circuit Q2, a higher outputpower is normally possible by increasing the primary supplyfrequency and track current or by improving the magneticcoupling to the secondary pickup. In practice, however, nominaltrack frequencies of 10–20 kHz are used for systems above10 kW, as limited by available power switches; 40-kHz systemsare typical in the 100 W–10 kW region and so forth.

In monorail IPT systems, a common pickup configuration isthe E-pickup as illustrated in Fig. 2(a) [1], [20]. Because theIPT track wires are placed in close proximity to each other, thepickup is designed such that the field contributions from bothwires couple in a constructive manner into the secondary powerpickup, thereby improving the power transfer capability of thesystem. In vehicle applications, the track wires are normallyburied in the ground, and the permeable material within thepickup cannot extend into the track. In such cases, either“flat-E” or simple “flat” pickups are used as shown in Fig. 2(b)and (c), respectively.

When the vertical and horizontal power profiles of thesepickups are considered, taking into account potential horizontalmisalignment from the track center as shown in Fig. 3, thereare marked differences. In the case of a simple flat pickup,the ferrite and coil are configured to capture the horizontalflux component of the field above the road surface. The fieldcontribution is at a maximum when the pickup is centeredabove any one wire and is at a minimum in the center pointbetween the two wires [Fig. 3(b)]. For a given track width, thetolerance to horizontal misalignment is maximized by separat-ing the track wires and using two pickups. A flat-E pickup isconfigured to capture the vertical component of the flux, whichis a maximum between the two wires. The power capabilityof the flat-E is best when correctly aligned with the track;however, this profile drops considerably with small horizontalmisalignments [Fig. 3(a)]. Because simple flat pickups are ableto perform considerably better with misalignment, they aretherefore preferred where horizontal vehicle movement cannotbe easily constrained [5], [6], [10].

If the vehicle movement in a given system is sufficientlyconstrained, a pickup with tighter coupling to the track can beused. This allows the use of very tightly coupled pickups, suchas coaxial pickups [7], for applications where the vehicle movesalong a track or is stationary during the power transfer, such asa battery charging system [9], [10]. The disadvantage of these


Fig. 3. Comparison of flat pickup power profiles, with 60-mm track spacing.(a) Flat-E. (b) Simple flat.

pickups is of course that they inherently disallow movement ormisalignment of the vehicle while power is being transferred.

III. CONCEPTUAL THREE-PHASE SYSTEM

In the proposed system, the objective is to produce a tracklayout that provides increased power delivery with greater hor-izontal tolerance for a given pickup topology. A polyphase IPTsystem that can achieve this goal is defined as being any systemin which the currents in the track cables are electrically sepa-rated by angles other than an integer multiple of 180◦ such thata rotational magnetic field is produced around the track. Single-phase distributed-winding (where several wires are spread outto create any particular phase), sequential concentric-coil, andmeander-coil type tracks [12], [13] are naturally excluded, inthat these track topologies have currents that remain separatedby 180◦ and produce a stationary field across the track.

A given IPT track system can be labeled as either “unipolar”or “bipolar,” depending on whether the pickup is exposed toonly the forward currents or both the forward and reversecurrents of each phase. For example, a three-phase systemcan utilize either three track cables (unipolar) with a closed-delta output transformer [Fig. 4(a)] or six track cables (bipolar)with an open-delta output transformer [Fig. 4(b)]. While bothsystems have been studied, this paper considers the bipolarsystem in particular with the track phases laid out as shownin Fig. 4(c).

An N -phase IPT system is spaced in a manner similar toa linear induction motor such that each adjacent track cablecarries a current separated by 360◦/N . This produces a time-varying field that sweeps across the track horizontally, allowinga simple flat pickup to capture the horizontal component of thisfield across the entire width of the track. In the proposed three-

Fig. 4. (a) Unipolar 3Φ system with closed-delta output transformer.(b) Bipolar 3Φ system with open-delta output transformer. (c) 3Φ bipolar tracklayout as driven by a positive phase sequence inverter with ferrite pickup.

phase system, the phase displacement between the currents inadjacent track wires is either 120◦ or 60◦ depending on whethera unipolar or bipolar arrangement is used so that the horizontalfield component does not directly cancel between each wire.This can be compared to the horizontal component capturedfrom a simple flat pickup in a traditional single-phase system[as shown in Fig. 3(b)] that contains statically positioned nullsin the power transfer profile.

The physical separation of the wires and the presence offerrite in the pickup [Fig. 4(c)] distort the field vectors so thata simple field analysis will not accurately predict the receivedpower. Consequently, FEM modeling was undertaken to helpdetermine the output power for a given ferrite block. This willbe discussed in later sections.

IV. PROTOTYPE THREE-PHASE POWER SUPPLY

A three-phase inverter power supply was used as shown inFig. 5. This system is composed of a three-phase full-bridgeconverter, a three-phase isolating transformer, which is part ofthe LCL network in each phase, a control circuit, and open-circuit protection circuitry (not shown). Here, the three-phasetransformers provide galvanic isolation to the system. Using anopen-delta arrangement on the transformer secondary enablesall phases and their return paths to be available at the output.

In practice, the isolation transformer is constructed usingthree separate single-phase transformers placed directly be-tween the bridge converter and the track. These transformershave some leakage inductance and are designed to form the firstL of the LCL network in each phase. Since most of the reactivecurrent circulates in the last CL of this network, only realpower goes through the transformers and the voltampere rating


Fig. 5. Resonant inverter and track layout.

of each transformer is small. In order to minimize conductednoise generated by the inverter, the primary and secondary coilsare separated with a 2-mm gap and individually wound. Thisreduces the interwinding capacitance but produces a large valueof leakage inductance in each transformer.

If there is a volt-second imbalance due to unequal conductionvoltage drop or unequal switching times in the power switches,then a dc voltage component can be present at the inverteroutput that will saturate the transformer cores. Capacitors C1

to C3 are used to prevent such saturation but are also de-signed to compensate for variations in the transformer leakageinductances.

Capacitors C4 to C9 are designed to tune the various trackinductances L, which should all be made identical to ensurebalanced phase currents. The value of these capacitors dependson the transformer turns ratio.

A. Novel Voltage Controller

In normal operation, a bridge inverter is operated usinga standard six-step pulse enabling a fundamental line-to-linevoltage given in terms of the dc voltage Vd by

VLL =√

6π

Vd. (5)

A problem with such a topology is that the inverter cannotcontrol the magnitude of the output ac voltages, requiring thedc input voltage to be controlled in order to control the outputvoltage magnitude and, consequently, the magnitude of thecurrents in each phase.

A pulsewidth modulation (PWM) technique is presented hereto vary the output voltage magnitude removing the need toadjust the input dc voltage. Here, the voltage output of eachphase is essentially a square wave except for a notch and apulse to control the ratio of the third harmonic content in theoutput. The phases A and B output voltages VAN and VBN andthe line-to-line voltage between these phases VA−B are plottedin Fig. 6(a). A notch is introduced in the first half-cycle at60◦, and a pulse is introduced in the other half-cycle at 240◦

in each phase waveform. The width of each notch and pulseare made identical and can be adjusted from 0◦ to 60◦ usinga microprocessor. These variations are indicated by arrowsin Fig. 6(a), along with the effect on the line-to-line voltage

Fig. 6. (a) Variable output voltage control. (b) Measured PWM gatewaveforms.

VA−B . If the width of these notches and the pulses is set to 0◦,the output voltage will be identical to that created using thestandard six-step waveforms having a maximum fundamentalvalue given by (5).

If their width is adjusted to 60◦, then the output line-to-linevoltages will have in-phase common-mode third harmonics,and since the delta arrangement of the isolation transformerpresents infinite impedance to triplen harmonics, no current willflow in the track. In practice, the track currents can be measuredand adjusted by varying the width of these additional pulsesand notches giving full current control. Fig. 6(b) shows themeasured gate signals to the power switches Q1, Q2, and Q3

of Fig. 5 that produce the required output phase voltages.If the width of the notch is defined as θ and the voltages are

normalized to have a magnitude of 1, the harmonic content ofthe line-to-neutral output voltages VAN and VBN and the line-to-line output voltage VA−B can be obtained by a Fourier seriesexpansion [defined in (6)] of the waveforms in Fig. 6(a), i.e.,

V (x) = a0 +∞∑

n=1

[an cos

(2πnx

T

)+ bn sin

(2πnx

T

)].

(6)

The coefficients of the harmonics in the line-to-neutral volt-age are given by

a0 =0.5

an =1

nπ

(sin

(nπ

3

)− sin

(nπ

3+ nθ

)+ sin

(4nπ

3+ nθ

)

− sin(

4nπ

3

))

b0 =−1nπ

(cos

(nπ

3

)− cos

(nπ

3+ nθ

)+ cos

(4nπ

3+ nθ

)

− cos(

4nπ

3

)+ cos (nπ)−1

). (7)


Fig. 7. Harmonic content of the output voltage with changing pulsewidth.(a) Line-to-neutral voltage. (b) Line-to-line voltage.

The line-to-line voltage is an even function so that all bn

coefficients are zero. The an coefficients can be shown to begiven by

a0 = 0

an =2

nπ

(sin

(n

6(2π − 3θ)

)− sin

(nθ

2

)− sin

(n

(π− θ

2

))

+ sin(n

6(4π − 3θ)

)). (8)

The first five nonzero harmonics for the line-to-neutral andline-to-line voltages are shown graphically in Fig. 7(a) and (b),respectively. Note that in both graphs, the magnitude of the fun-damental component decreases approximately linearly from itsinitial value to zero as the pulsewidth increases from 0◦ to 60◦.It is this fundamental component that controls the track current.From Fig. 7(a), it can be seen that as the pulsewidth increases,the fundamental component is replaced entirely by triplenharmonics, which are blocked by the delta-connected outputtransformer. The remaining harmonic components, which arenot blocked by the output transformer, are attenuated by thefiltering properties of the LCL network [21] so that the trackcurrent is essentially purely sinusoidal.

In an IPT system, all of the track wires are fully insulated (insome cases buried in concrete), and consequently, accidentalshort or open circuit faults, while possible, are extremely un-

likely. If either condition were to occur, however, the current inthe bridge would slowly increase with a rate of rise limited bythe relatively large inductances at the output of the bridge (cor-responding to the large leakage inductances of the transformersas aforementioned). If the currents are not interrupted, they willeventually damage the power switches. The required protectioncircuit as shown in Fig. 5 is extremely simple. The slow rateof rise is easy to detect using three current transformers in theoutput phases of the inverter. Over current in any one phasecauses a trigger signal that shuts down the inverter if a faultoccurs. Upon receiving the signal, the microcontroller attemptsto restart the inverter after a short delay. If the fault persists, theinverter will shut down again.

In practice, the three-phase track currents within the con-structed prototype were found not to be exactly balanced, asshown in Fig. 8(a). This is largely due to an imbalance in themutual coupling between the relative phases. During operation,each phase couples voltage into the adjacent phases, but theasymmetries in the cable positions mean that these couplingfactors vary. The effect of the magnetic coupling is to changethe effective inductance (apparent length) of each phase, andthis effect is more noticeable in the two outer phases. In orderto balance all phases in the system, additional inductance canbe added into the A and B phases, while phase C may require aseries capacitor to reduce its overall inductance to match. Theresulting balanced currents are shown in Fig. 8(b).

V. PICKUP AND TRACK DESIGN

For the purpose of constructing a prototype pickup, one ormore ferrite blocks with similar characteristics to 3C85 wereused as they were readily available. The dimensions of eachferrite block are given as follows: length = 33 mm, width =50 mm, and thickness = 3 mm. The initial pickup windingcomprised 20 turns of 0.7-mm enameled wire as illustrated inFig. 9(a). Alternative pickup configurations used two ferriteblocks with a single identical core [Fig. 9(b)] or dual 10-turnseries-connected windings [Fig. 9(c)].

The prototype vehicles used here are radio-controlledTAMIYA 1 : 10 scale quick drive sports car models. In order tosupply the car with the equivalent power as the usual battery, thepickup should be capable of delivering 30 W. The dimensionsof the car are given as follows: length = 440 mm, width =180 mm, and height = 140 mm, with 10-mm ground clearance.The track width is 2.5 times that of the car. Notably, the positionof the pickup above the cables, the magnitude of the trackcurrent, and the cable spacing are three major factors that affectthe track design. As the height of the car is fixed, then theposition of the ferrite above the cables is also fixed at 6 mm.The pickup could become skewed relative to the track cable,but here it is assumed to be perfectly aligned (i.e., the vehicleis traveling in a straight line). The car should nominally drivein the center of the track; however, it is also desirable thatsufficient power is available at the extremes to enable the carto return to the center of the track.

A 3-D FEM analysis was undertaken using JMAG-Studiodeveloped by The Japan Research Institute Limited. This workwas undertaken prior to the actual construction of the track or


Fig. 8. Measured track currents. (a) With current imbalances. (b) After balancing.

Fig. 9. (a) Single ferrite. (b) Single winding double ferrite block. (c) Dualwinding double ferrite block.

TABLE ITRACK CONFIGURATIONS

power supply. Three track designs were investigated, and theircable positions are given in Table I. Simulation results wereundertaken using a similar approach to that discussed in [22]and later verified experimentally using the single pickup ofFig. 9(a). In both simulation and experiment, the pickup wasshifted horizontally in 10-mm intervals from the track center.The track current was maintained at 40 A/phase in all cases,and the frequency of the supply was 38.4 kHz. The open-circuitvoltage and short-circuit current of the pickup as defined in (1)and (2) were obtained at each position. The uncompensatedoutput power of the pickup was then calculated using (3).Results of this comparison are shown in Fig. 10.

Fig. 10(a) presents the simulated results, whereas Fig. 10(b)shows the measured results. As noted, there is a close agree-ment, enabling further design and investigation using simula-tion. Small discrepancies result from slight variations in track

Fig. 10. (a) Simulated Su versus distance off center. (b) Measured Su versusdistance off center.

positioning in the experimental system. As noted, Track 1 is notsymmetrical. The phases closer to the track center are spacedcloser together. As expected, the calculated uncompensatedpower is higher here and the power drops off rapidly. Tracks 2and 3 are designed with equal spacing between phases, resultingin a more continuous power profile across the width of the track.A comparison of power profiles from these two configurationsshows that the pickup width should be large enough to captureflux from more than one phase. Ideally, it should be at leasttwice the spacing between any two phases.

In general terms, the power profile can be widened furtherif the spacing between the track phases is increased. However,comparing the results of Tracks 2 and 3, for a given pickup size,any increase in track spacing causes a corresponding (almostlinear) reduction in the power density of the system, which


Fig. 11. (a) Simulated Su versus distance for the single-coil pickup on athree-phase and a two-phase bipolar track. (b) Measured Su versus distanceoff center.

reduces the uncompensated power available to the pickup(but without substantially affecting the system efficiency). Thepower peaks and troughs (notable as the pickup is movedacross the complete width of the track) will also becomemore pronounced in a system with wider spaced tracks as thepickup is unable to capture all of the available field lines fromneighboring track phases.

A. Finalizing Track and Pickup Designs

As shown, Track 3 has a power profile that extends nearly135 mm but requires a wider pickup to ensure a greater propor-tion of the field flux is captured at all points across the track,thereby boosting power and providing a much flatter powercurve over the center of the track.

Consequently, Track 3 was chosen for the final design.JMAG was used to simulate the performance of a new improveddouble-width single-coil pickup, which was constructed asshown in Fig. 9(b) using two of the original ferrite blocks,on the chosen three-phase bipolar Track 3 layout. For thepurpose of comparison, this new pickup was also simulatedon a conventional single-phase bipolar track. Both tracks wereassumed to have identical currents of 40 A/cable at 38.4 kHz.The results are presented in Fig. 11(a).

Maximum power is transferred when a vehicle’s powerpickup is perfectly aligned with the track cable. In practice,some misalignment from this ideal driving position is to beexpected. In order to provide this tolerance in a single-phasetrack system, it is assumed that the target minimum poweroccurs at 50% of maximum so that a usable driving region ispossible. Fig. 11(a) shows that a vehicle with a single pickuphas two 30-mm-wide zones within which it can operate. Insideeach driving zone the untuned pickup delivers a minimumof 3.4 VA. In comparison, the three-phase track delivers anuntuned power of at least 7.5 VA over a 220-mm driving zone.As such, the three-phase track delivers a continuous powerprofile more than twice as large as the single-phase system,and that extends almost the entire width of the constructedtrack—the lateral tolerance is at least three times better. Inpractice if the three-phase system is operated with slightlyless than two thirds of the required track current/phase ofthe single-phase system, both systems will provide a similaraverage uncompensated power Su to a pickup operating in theirrespective usable operating regions. Thus, for a desired outputpower, the operational efficiency of both systems will be almostidentical (approaching 80%–85% near rated operation) as thelosses in the three-phase system running under this reducedtrack current are similar to the single-phase system despite theadditional switches, transformers, capacitors, and track phases.The star–delta transformer also tends to balance these lossesacross the inverter bridge irrespective of the position of thepickup.

The designed prototype toy car system requires 30 W to bedelivered from the power pickup. Using the three-phase tracksystem with the vehicle driving near the track center, the un-compensated power achieved from Fig. 11(a) is approximately9.5 VA so that the required Q2 using (4) is around 3. This lowcircuit Q ensures a low voltampere rating for the pickup andhigh tolerance to tuning errors [15]. At 135-mm off track center,the measured uncompensated power is around 1.5 W. AssumingQ2 = 10 is allowed by the vehicle power controller, this wouldenable sufficient power to drive the vehicle back toward thetrack center.

Fig. 11(b) shows the resulting measured power profiles usingtwo pickup designs constructed as shown in Fig. 9(b) and (c).Of note is that the power profiles of each are similar, withoutputs close to that predicted using JMAG. The single-coilpickup is found in practice to have a “smoother” power profileover the center of the track and was consequently chosenin the final system. The profile of the single-coil system issurprisingly flatter than predicted by simulation [Fig. 11(a)].This arises from slight practical variations in the constructedsystem compared with that simulated. The physical construc-tion of the pickup of Fig. 9(b) required two ferrite blocks tobe glued together. This process is not exact, and there existsa nonuniform air gap in construction. For the purpose of thesimulation, this is assumed to be a uniform air gap of 0.4 mm.Furthermore, in practice, the coil width of the pickup is slightlylarger (JMAG assumes a compact coil), and the track currents,while close, are not exactly 40 A/cable. A larger coil acts toreduce the secondary leakage and therefore improves couplingand power, while the presence of the air gap effectively lowers


the relative permeability of the ferrite and therefore reducespower. Simulations have shown that if this air gap were notpresent, the power output would increase by more than 35%,which further enhances the benefit of the three-phase systemover a traditional single-phase track system.

VI. CONCLUSION

A new three-phase IPT system has been developed. Thepower supply comprises a standard six-switch inverter withresonant LCL topology incorporating full isolation and protec-tion means. A novel PWM generation scheme that enables fullcontrol of the track currents without modifying the dc voltage aswould normally be required was presented. The effect of mutualinductance between parallel long cables results in imbalancesbetween the track phases. However, by finely tuning the trackinductance, the track can be balanced.

The track geometry and pickup design were undertaken usingFEM modeling and experimentally verified. A pickup havinga width that couples flux from all three phases is desirablebecause of the time-varying nature of the field. The width ofthis pickup should be more than double the spacing betweenphases.

The prototype system has shown that continuous powertransfer to moving vehicles is possible without placing sig-nificant restrictions on vehicle operators. The prototype couldeasily be scaled to deliver the necessary power demands forpractical electric vehicles.

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[13] F. Sato, H. Matsuki, S. Kikuchi, T. Seto, T. Satoh, H. Osada, and K. Seki,“A new meander type contactless power transmission system—Active ex-citation with a characteristics of coil shape,” IEEE Trans. Magn., vol. 34,pt. 1, no. 4, pp. 2069–2071, Jul. 1998.

[14] C.-S. Wang, O. H. Stielau, and G. A. Covic, “Design considerations for acontactless electric vehicle battery charger,” IEEE Trans. Ind. Electron.,vol. 52, no. 5, pp. 1308–1314, Oct. 2005.

[15] O. H. Stielau and G. A. Covic, “Design of loosely coupled induc-tive power transfer systems,” in Proc. Int. Conf. Power Syst. Technol.(Powercon), Perth, Australia, Dec. 4–7, 2000, vol. 2, pp. 85–90.

[16] C.-S. Wang, G. A. Covic, and O. H. Stielau, “Power transfer capabilityand bifurcation phenomena of loosely coupled inductive power trans-fer systems,” IEEE Trans. Ind. Electron., vol. 51, no. 1, pp. 148–157,Feb. 2004.

[17] C.-S. Wang, G. A. Covic, and O. H. Stielau, “Investigating an LCL loadresonant inverter for inductive power transfer applications,” IEEE Trans.Power Electron., vol. 19, no. 4, pp. 995–1002, Jul. 2004.

[18] M. Borage, S. Tiwari, and S. Kotaiah, “Analysis and design of an LCL-Tresonant converter as a constant-current power supply,” IEEE Trans. Ind.Electron., vol. 52, no. 6, pp. 1547–1554, Dec. 2005.

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[22] D. Kacprzak, G. A. Covic, and J. T. Boys, “An improved magnetic designfor inductively coupled power transfer,” in Proc. Conf. IPEC, Singapore,Nov. 1–Dec. 2005, pp. 1–4.

Grant A. Covic (S’89–M’91–SM’04) received theB.E.(Hons.) and Ph.D. degrees from The Universityof Auckland, Auckland, New Zealand, in 1986 and1993, respectively.

He is a full-time Associate Professor with theDepartment of Electrical and Computer Engineering,The University of Auckland. His current researchinterests include power electronics, ac motor control,electric vehicle battery charging, and inductive (con-tactless) power transfer.

John T. Boys received the Ph.D. degree from TheUniversity of Auckland, Auckland, New Zealand,in 1962.

After gaining the Ph.D. degree, he worked for SPSTechnologies, USA, for five years before returning tothe academe. He is currently a Professor of electron-ics and the Head of the Department of Electrical andComputer Engineering, The University of Auckland.He is the holder of more than 20 patents. His fieldsof interests are motor control and inductive powertransfer.

Dr. Boys is a Fellow of the Institution of Professional EngineersNew Zealand.


Michael L. G. Kissin (S’02) received the B.E. de-gree in electrical and electronic engineering fromThe University of Auckland, Auckland, NewZealand, in 2005. He is currently working toward thePh.D. degree at The University of Auckland.

His research interests include inductive powertransfer and roadway-powered electric vehicles.

Howard G. Lu received the B.Eng. degree fromGuangdong University of Technology, Guangdong,China, in 1986 and the M.Sc.(Hons.) degree fromthe University of Waikato, Hamilton, New Zealand,in 1996.

He then spent ten years as an Electrical Engineerat Guangdong Power Grid, China. He is currently aTechnician of power electronics with the Departmentof Electrical and Computer Engineering, The Univer-sity of Auckland, Auckland, New Zealand.

Documents

Three-Phase Inductive Power Transfer System for Roadways