A 22 kW-85 kHz Three-phase Wireless Power
Transfer System with 12 coils
Keisuke Kusaka
Dept. of Electrical, Electronics, andInformation Engineering
Nagaoka University of Technology
Nagaoka, Japan
kusaka@vos.nagaokaut.ac.jp

Rintaro Kusui
Dept. of Electrical, Electronics, andInformation Engineering
Nagaoka University of Technology
Nagaoka, Japankusui@stn.nagaokaut.ac.jp

Jun-ichi Itoh
Dept. of Electrical, Electronics, andInformation Engineering
Nagaoka University of Technology
Nagaoka, Japan
itoh@vos.nagaokaut.ac.jp

Daisuke Sato
Nagaoka Motor Development Co., Ltd.,
Nagaoka, Japansato@nagaoka-md.co.jp

Shuichi Obayashi
Corporate Research & Development Center
Toshiba Corporation
Kawasaki, Japanshuichi.obayashi@toshiba.co.jp

Masaaki Ishida
Corporate Research & Development Center
Toshiba Corporation
Kawasaki, Japanmasaaki1.ishida@toshiba.co.jp

Abstract—A 22-kW three-phase wireless power transfer(WPT) system with six primary coils and six secondary coils areproposed is developed in this paper. For the rapid charging of
onboard batteries of electric vehicles (EVs), the output power ofthe WPT system should be increased. However, the increase of
the WPT system may cause an increase in leakage magnetic flux.
The leakage magnetic flux from the WPT system must comply
with guidelines. The proposed WPT system with 12 solenoid
coils, which are circularly placed, suppresses the leakage
magnetic flux with two opposite solenoid coils which are
differentially coupled. Moreover, the circularly placed coils
cancel out the magnetic interference, which is caused by the
coupling among the primary coils, and secondary coils. In thispaper, first, the 22-kW three-phase WPT system is designed.
Especially, the figure of coils, which achieves a minimum
volume of the core, are determined with the cancellation
capability of the magnetic interference. Then, the 22-kW WPTsystem is tested with a constant voltage load. The maximum
efficiency of the proposed WPT system from a primary DC to asecondary DC side is 91.1% at an output power of 22 kW with a
400-V constant voltage load.
Keywords—wireless power transfer, three-phase transmission,
leakage magnetic flux

I. INTRODUCTION
In recent years, the number of electric vehicles (EVs) has
been increasing for reducing the exhaust of NOx gas fromvehicles for saving the environment. The travel distance of
EVs is typically shorter than the conventional vehicles withgasoline or diesel because the power density of batteries isgreatly smaller than the one of fossil fuel. Due to thislimitation, EVs cannot be a substitute for conventional
vehicles in terms of convenience because users are required tohave a frequent charging.
One of the solutions for improving the convenience of theusers of EVs is a wireless power transfer system (WPT) withmagnetic induction [1–8]. The WPT system transmits power
without any cable between the primary coil and the secondary
This work is financially supported by Ministry of Internal Affairs andCommunications, Japan.

coil. Thus, the users will start the charging of the onboard
battery without an onerous connection of a charging cable.
Focusing on a wired battery charger, charging power ofEVs has been increasing. Owing to the increase of battery
capacity for achieving long travel distance, the battery
capacity of EVs tends to be increased [9]. Based on thisbackground, the charging power of the wired battery charger
had been increased to 50 kW as the rapid charging in thestandard published by CHAdeMO. Moreover, in the next
generation standards, the charging power will be increased to
350 kW or 500 kW as an ultra-fast charging. Considering theincrease of the charging power of the wired chargers, theoutput power of the WPT system should be increased forshortening charging time. In the standard, which will be
published by SAE and ISO, the WPT system will be classified
into five classes; WPT1–5, by the output power of the system.
In the largest output power class; WPT5, the output powershould be larger than 22 kW.
The increase of the output power may cause a problem of
the leakage magnetic flux. The leakage magnetic flux must notaffect any electronic devices, radio communications, and thehuman body. Thus the WPT system must comply with theguidelines, which are legislated in each area or country withconsidering the guidelines of ICNIRP and CISPR. As one of
the solutions for reducing the leakage magnetic flux, theshielding with metal or magnetic material has been used [10–13]. However, the shielding method causes a large power loss
due to the eddy current on the shield. In [14–16], the reduction
method of the leakage magnetic flux using spread spectrum
has been proposed. The radiation noise is reduced byrandomly changing the output frequency of the inverter. Bychanging the transmission frequency, the peak of the radiationnoise is suppressed in the frequency domain. However, theproposed method will not comply with the standard of the
WPT for EVs. In [17], two-channel WPT system has been

proposed. The two-channel WPT system has two solenoidtype primary coils and two secondary coils, which are placed
to cancel out the leakage magnetic flux, respectively. As the
drawback of the cancellation of leakage magnetic flux, themain flux, which contributes transmitting power, is alsointerfered.
In this paper, the WPT system with 12-solenoid coils is
proposed in order to reduce the radiation noise without any
magnetic interference among the coils. The six primary coils
are circularly placed to cancel out the magnetic flux due to thedifferential connection and magnetic interference among the
primary coils. By adjusting the relationship of the magnetic
coupling among the primary coils, the magnetic interference
can be canceled out. The secondary coils are placed in thesame manner. In the rest paper, first, the principle of theradiation noise cancellation is explained. Then, the 22-kW
WPT system is designed considering the cancellation
conditions. Finally, the 22-kW WPT system is performed.

iu1

u1
Cu1

VDC1

Cv1

w1

u2

Lu1A

Lu2A

Lu1B
Lv1B

Lw2B

Lv1A

v1

iu2

M

Cu2

Lu2B
Lw1B
Lw1A

Lw2A

Lv2B

Lv2A

Cv2

VDC2

v2
w2

Cw2

Cw1

Fig. 1. Proposed three-phase wireless power trasnfer system with six
primary coils and six secondary coils. The primary coils and the secondary
coils are connected in star-star winding.

II. THREE-PHASE WPT SYSTEM WITH 12 COILS
A. Circuit Configuration
Figure 1 shows the circuit configuration of the proposed
three-phase WPT system. The proposed WPT system is basedon a three-phase WPT system [18–20] with two coils on eachphase. The three-phase transmission coils are connected instar-star windings. It should be noted that the star-star windingis used in this paper however the star-delta, delta-star, or deltadelta windings can be used for the system. Power istransmitted through the magnetic couplings M between the
primary coils and the secondary coils, e.g., the coupling
between Lu1A and Lu2A. The WPT system has the three-phase
inverter, which is operated with a square wave operation, inthe primary side and the three-phase diode bridge rectifier inthe secondary side. In order to compensate the leakage
inductance due to weak magnetic coupling, the six resonantcapacitors Cu1, Cv1, Cw1, Cu2, Cv2, and Cw2 are connected into
the output of the three-phase inverter and the input of thethree-phase inverter. The capacitors are selected to resonate
with the self-inductance of each coil at the transmission
frequency. Owing to the resonance, the inverter is operated ina zero-voltage switching (ZVS) condition when the output
frequency is slightly higher than the resonant frequency
because the parasitic capacitors of the MOSFETs, which will
turn-on after the dead time, are discharged during the deadtime by the circulating current [21].
Figure 2 shows the schematic of the transmission coilswith the resonant capacitors. The solenoid coils are used for
power transmission. The six solenoid coils are circularly
placed. The pairs of two solenoid coils on the common phaseare placed opposite and connected in series. The amplitude of
the magnetic flux generated by two coils is the same because
the inductance value is common on the primary coils and thesecondary coils, respectively. The leakage magnetic flux iscanceled out at a measuring point; it is typically 10 m from thesystem because the direction of the magnetic flux caused by
the coils are opposite [17]. The coils on phase-v and phase-w
are placed 120 degrees. Using the three-phase WPT system,
the transmission power of each phase is divided into three.
Thus, it is effective to reduce the duty of each coil in term ofheat generation. However, the multiple coils cause a magnetic
interference such as Man, Mbn, and Mcn among the coils on thecommon side, i.e., primary side and the secondary side where(a) Primary coils(b) Secondary coils
Fig. 2. Placement of 12-solenoid coils. The solenoid coils in common
phase are placed opposite to each coil for noise reduction.

n represents the primary side as 1 and the secondary side as 2.
The magnetic interference will be a reason for the power loss
because the magnetic interference causes a circulating current
on the windings due to the unnecessary induced voltage. In the
proposed three-phase WPT system, the unnecessary induced
voltage is canceled out by adjusting these unnecessary
coupling as follow.
B. Cancellation of Magnetic Interference
In this section, the cancellation method of the magnetic
interference, which will degrade the transmission efficiency,
is explained. The induced voltage caused by the magnetic
interference is canceled out by adjusting the coupling among
the coils on the primary side and the secondary side.
The induced voltages on the 12 coils vunx, vvnx, and vwnx are
expressed as (1) considering the magnetic interference Man,
Mbn, and Mcn where x represents A or B. The variable L1Ys is
the self-inductance of each coil. Focusing on the first line of
the matrix in (1), the induced voltage is expressed as (2).

 vu1 A   L1Ys  vu1B   M c1
 v  M
 v1 A   b1
 vv1B   M a1
 
 vw1 A   M b1
 vw1B   M a1=
 vu 2 A   M
v   0
 u2B  
 vv 2 A   0
v   0
 v2B  
 vw 2 A   0  vw 2 B   0

M c1

M b1

M a1

M b1

M a1

M

0

0

0

0

L1Ys

M a1

M b1

M a1

M b1

0

M

0

0

0

M a1

L1Ys

M c1

M b1

M a1

0

0

M

0

0

M b1
M a1

M c1
M b1

L1Ys
M a1

M a1
L1Ys

M b1
M c1

0
0

0
0

0
0

M
0

0
M

M b1

M a1

M b1

M c1

L1Ys

0

0

0

0

0

0

0

0

0

0

L2Ys

M c2

M b2

M a2

M b2

M

0

0

0

0

M c2

L2Ys

M a2

M b2

M a2

0

M

0

0

0

M b2

M a2

L2Ys

M c2

M b2

0

0

M

0

0

M a2

M b2

M c2

L2Ys

M a2

0
0

0
0

0
0

M
0

0
M

M b2
M a2

M a2
M b2

M b2
M a2

M a2
M b2

L2Ys
M c2

diu1
di
+ M u1
dt
dt
di 
di
 div1 diw1  di− M a1 ++ M b1  v1 + w1  − M c1 u1

dt 
dt 
dt
 dt
 dt

vu1 A = L1Ys

(2)
The first term of (2) is the induced voltage by the selfinductance and the second term is by the mutual coupling
between the primary coil and the secondary coil, whichtransmits power. In contrast, the third to fourth terms are theinterferences, which do not contribute to transmitting power.
These terms should be canceled out. If the sum of the third to
fourth terms can be zero, the proposed WPT system is as sameas the single-phase WPT system with a primary coil and asecondary coil. In other words, the effect of other coils is notconsidered.
When the magnetic coupling between the primary coil andthe secondary coil is the same, the three-phase current isequilibrium. Thus, (2) can be simplified as
vu1 A = L1Ys

diu1di
+ M u2
dt
dt


2 

−ω I m1  M c1 cos ω t + ( M a1 − M b1 ) cos  ω t − π 
3 


4 

+ ( M a1 − M b1 ) cos  ω t − π  
3 

(3),

0   iu1  
0   − iu1 
0   iv1  
0   − iv1  
0   iw1 M  d  − iw1  
M a 2  dt  iu 2 
M b 2   − iu 2  
M a 2   iv 2 
M b 2   − iv 2  
M c 2   iw 2  
L2Ys   − iw 2 

The self-inductance of the primary coils are common.
Therefore the condition for the cancellation is expressed as

ka1 = kb1 + kc1

Equation (3) shows that the unnecessary induced voltage
can be canceled out by adjusting the relationship among Ma1,
Mb1, and Mc1 as shown by

M a1 = M b1 + M c1

(4).

(5)

with the magnetic coupling coefficient among the primarycoils.
The cancellation condition of the induced voltage on thesecondary side is introduced in the same manner as

ka 2 = kb 2 + kc 2

(6).

By satisfying both (5) and (6), the interference among themultiple coils is canceled out.

III. DESIGN OF 22-KW PROTOTYPE
A. Parameter Design
The transmission coils for designed in this section under
the assumption that the magnetic interference is canceled out
by satisfying (5–6).
The required inductance for each coil is calculated from animpedance matching condition. Thus, first, the equivalent AC
resistance is calculated from a rated output power Pm and theDC voltage VDC2 on the secondary side. The AC resistance Req
is expressed as (7) considering a derivation of an equivalent
AC resistance of a single-phase rectifier [22]. The equivalent
resistance is for one phase.

2)
8 (V
6 V2
Req = 2 DC 2
= 2 DC 2
π
Pm 3
π Pm
2

where Im1 is the amplitude of the primary current.

(1)

(7)

The proposed system assumes to use a battery with avoltage of 200 to 400 V. Then, the system is designed totransmit 22 kW when the secondary DC voltage is 400 V. It
means that the output power is derated when the secondary
DC voltage is lower than 400 V.

Besides, the system is designed with an input power of 25
kW with a margin to ensure the input power of 22 kW
considering the power loss of the system. Thus, the equivalent
resistance is calculated with VDC2 = 400 V and Pm = 25 kW.
The secondary inductance is designed from the impedance
matching condition. The impedance of the self-inductanceshould be equal to the equivalent AC resistance expressed by(7). In this consideration, parasitic resistances of the windings
are ignored for simplicity. The secondary inductances for eachphase are calculated by
2
6 VDC
2
L2Y = Lu 2 = Lv 2 = Lw2 = 2
π kω Pm(8),

where ω (ω = 2πf) is the angular transmission frequency andk is the coupling coefficient between the primary coil and thesecondary coil at the nominal position. The inductancecalculated by (8) is for each phase. Thus, the inductance ofeach coil L2Ys is the half of L2Y because the proposed system
has two coils in series in the phase. Note that, the transmission
frequency f is the 85 kHz with considering the standardizationof the WPT system by SAE and ISO.
The primary inductance is calculated from the voltage gain.

VDC 2
L
= 2Y
VDC1
L1Y

(9)

cancellation and the magnetic interference cancellation
introduced in the previous chapter.
Figure 3 shows the schematic of the transmission coils forelectromagnetic analysis. In this chapter, the transmission coilfor a 22-kW wireless charger is analyzed and developed. First,
the size of each coil is determined to obtain the required main
coupling k for the power transmission. In this paper, the sizeof each transmission coils are x = 500, y = 250 mm for themain coupling k of 0.55 when the nominal transmission
distance is d = 90 mm.
Then, the radius of the transmission coils is analyzed inorder to satisfy the cancellation condition of the magnetic
interference shown by (3). Figure 3 shows the analyzed
model. The magnetic coupling among the coils on thecommon side mostly depends on the size of the ferrite core,
and the placement of cores such as the radius r. The size of theferrite core has been determined to obtain the required main
coupling k. Thus, the cancellation conditions have to besatisfied by adjusting the distance r. Table I shows the effectof the radius r analyzed using an electromagnetic analysis. In
this study, the radius r is designed to let the remaining
magnetic interference ka - kb - kc ten times smaller than themain coupling k. By increasing the radius r up to 750 mm, theremaining coupling is reducing because the effect of ka isdominant. The transmission coils should be as small as
possible as the cancellation condition is satisfied. Thus, in thispaper, a radius r of 750 mm is selected from this consideration.
Table II shows the results of the design consideration of thecoils.

From (8–9), the primary inductance is calculated by

L1Y = Lu1 = Lv1 = Lw1 =

2
6 VDC
1
π kω Pm
2(10).

The inductance for each coil L1Ys is half of L1Y in the samemanner to the secondary side.
Next, the resonant capacitors are selected. The resonantcapacitors play a role as a compensation of the leakage
inductances due to weak magnetic coupling k. In this paper, aseries-series compensation [23] is used. In the series-series
compensation, selecting the resonant capacitors to resonate
with a self-inductance of each coil at the transmission
frequency allows the efficient power transmission. Thus, theresonant capacitors for the primary side and the secondaryside is calculated as
Cu1Y = Cv1Y = C w1Y =C u 2Y = C v 2 Y = C w 2Y =1

ω 2 L1Y
1

ω 2 L2Y

(11)

(12)

Fig. 3. Analytical model of transmission coils.

TABLE I.

INTERFERENCE OF ELECTROMAGNETIC ANALYSIS.
Interference coupling

Cancellation

Note that, the output frequency is adjusted to be slightly
higher than the resonant frequency, which is determined by (8,10–12) for the ZVS of the inverter.

Radius
[mm]

ka

kb

kc

ka - kb - kc

700

0.100

0.02550.01750.0570

B. Size of coils
In this chapter, the shape of the transmission core isanalyzed with the electromagnetic analysis. The figure of thecore and the placement of coils are optimized for the minimum
volume of the core with the capability of the radiation noise

750

0.08930.02360.01630.0494800

0.08240.02200.01520.0452

C. Power Loss
The power loss in the proposed system is calculated in thissection.
1) Transmission coils
Assuming the unity power factor of the inverter output, thecurrent on the primary coil at the rated power is calculated by

I1 =

πP
P
=
3V1 3 2VDC1(13),

where VDC1 is the primary DC voltage, V1 is the phase voltageof the inverter output with the square wave operation
expressed as (14).

V1 =

2

π

(14)

VDC1

SPECIFICATION OF PROTOTYPE.

Parameters

Symbols

Value

Input power

Pm

25

kW

Primary voltage

VDC1

650

V

Secondary voltage

VDC2

200-400

V

Transmission frequnecy

f

85

kHz

Air gapg

90mm

Core width

x

500

mm

Core length

y

250

mm

Core thickness

z

10

mm

Radiusr

750mm

Number of turns
Self-inductance

The secondary current at the rated power is (15) in the samemanner.

I2 =TABLE II.

πP

(15)

3 2VDC 2

The copper loss of each primary coils and secondary coils
is calculated by

P1_ copper = r1I12(16),

P2 _ copper = r2 I 22

(17),

Primary

N1

8

turn

Secondary

N2

5

turn

Primary

L1Y

21.8

μH

Secondary

L2Y

8.52

μH

k

0.55

Main couplingCore

PC95, TDK

Litz wire

2UEWLZ 7×130 (φ0.1)
BSM300D12P2E001,
ROHM
IDW40G120C5BFKSA1,Infineon Technologies
Td
500
ns

MOSFETs
DiodesDead time

under the ZVS conditions. In contrast, turn-off loss occurs
whereas the output current is sinusoidal. Assuming thesinusoidal current on the output of the inverter, thesimultaneous current at the turn-off is approximately
expressed bywhere r1 and r2 are the AC resistance of the transmission coils.
Besides, the iron loss is calculated from the magnetic flux
density and the shape of coils. The iron loss is partiallycalculated with assuming the magnetic flux density is unity
entire the ferrite core.

B1 =LY 1s I1 + MI 2
N1S1

(18),

isw _ off = 2 I1 sin ωTd(21),

where Td is the dead time of the inverter. Thus the turn-off lossof the MOSFETs is expressed by

P1_ swoff = 6Eoff (isw _ off ,VDC1 ) × f sw

(22),

where Eoff is the turn-off energy of MOSFETs.

L I + MI1
B2 = Y 2 s 2
N 2 S2(19),

where Nn (n = 1, 2) is the number of turns and Sn is the crosssection of the ferrite core. The iron loss is read out from adatasheet of the magnetic material.
2) Inverter on the primary side
In the primary inverter, the switching loss and the
conduction loss occur. The conduction loss of the MOSFETs
in the primary inverter is calculated with assuming theconstant equivalent resistance rds(on).
2
1

Pinv _ cond = 6rds ( on) I(20)

The switching loss of the MOSFETs is calculated. The
turn-on loss does not occur because the inverter is operated

3) Rectifier on the secondary side
In the rectifier on the secondary side, the conduction loss
due to the diodes occurs. The current I2 flows through two
diodes during a half-period. Thus the sum of the conduction
loss of the six diodes is calculated by (23).
PD _ cond = 6

6 2VF I 2
1 T2
VF 2 I 2 sin ω tdt =
T 0
π

(23)

IV. EVALUATION OF 22-KW PROTOTYPE
A. System Configuration
The 22-kW WPT system is tested in this chapter. Figure 4
shows the system configuration for the test. The input poweris supplied through the variable transformer to adjust the input

Fig. 4. System configuration of 22-kW test of the proposed WPT system. The output power ciculates through the buck and boost chopper. The primary
DC voltage VDC1 is adjustable to control the power by the adjustable transformer at the input. The secondary DC voltage is also adjustable by the buckand boost chopper to emulate the constant voltage load such as batteries.

DC voltage. In the actual system, the primary DC voltage isadjustable by a PWM rectifier and a DC/DC converter for an
output power control. However, it is omitted in this test for
simplicity. The buck and boost chopper plays a role tocirculate power from the output of the WPT system to the
input. The secondary DC voltage VDC2 is adjustable from 200
to 400 V to emulate the constant voltage load such as onboard
batteries. Due to this configuration, the input power source
supplies only the power loss of the WPT system and the buckand boost chopper.
The transmission coils are developed according to theanalysis in chapter III. The magnetic interference is ka1 =0.0043, kb1 = 0.0026, kc1 = 0.0026, ka2 = 0.0090, kb2 = 0.0013,
and kc = 0.0006. The difference between the primary side andthe secondary side is caused by a difference in the number ofturns nevertheless the size of the ferrite core is the same.
B. Operation Waveforms
Figure 5 shows the operation waveforms when the primary
DC voltage VDC1 is 650 V and the DC voltage on the secondaryside is 400 V. The secondary DC voltage is controlled toemulate the constant voltage load by the buck and boost
chopper. The output power is 22 kW. From Fig. 5 (a), it isshown that the proposed WPT system is operated with theresonant condition because the zero-cross of the primary
current is synchronized with the voltage rise and fall of theinverter output voltage. From Fig. 5 (b), it is confirmed thatthe three-phase primary current is equilibrium. Fig. 5 (c)shows the three-phase output voltage of the inverter. The
output voltage slowly step up or down by the ZVS operation.
C. Efficinecy Evaluation
Figure 6 shows the efficiency characteristic of the
proposed WPT system. The efficiency from the primary DC
VDC1 to the secondary DC side VDC2 is measured. It means thatthe DC-to-DC efficiency take into account the power losses of
the inverter and the rectifier. The primary DC voltage VDC1 isadjusted to control the output power. By contrast, the(b) Inverter output voltage, and output currents on each phase(c) Inverter output current, and inverter output voltages on each phase(a) Inverter output voltage, output current, rectifier input current, andoutput voltage

Fig. 5. Operation waveforms of the proposed three-phase WPT system.
The output power is 22 kW, primary DC voltage VDC1 is 650 V, secondary
DC voltage is 400 V.

secondary voltage is changed to 200, 300, and 400 V toemulate the constant voltage load. The maximum output
power is derated by the secondary current on the coils and thediode bridge rectifier. The maximum efficiency is 91.1% at an
output power of 22 kW, 91.3% at 10 kW, and 91.3% at 5.0
kW when the secondary DC voltage is 400, 300, and 200V,
respectively.
Figure 7 shows a breakdown of the power loss when the
primary DC voltage is 650 V, the secondary DC voltage is 400
V, the output power is 22 kW. The power loss is calculatedwith the equations in chapter III. The power loss measured
through the experiment is 2.16 kW. On the other hand, the sumof the calculated power loss is 1.90 kW. The error between theexperiment and the calculation is 262 W where it is 1.2% of
the output power. In other words, the estimated efficiency by
the calculation is 92.0% as against a measured efficiency of
91.1%.
In the calculation, the power loss is dominated by the
conduction loss of the MOSFETs on the primary inverter. It
accounts for 26.9% of the total loss. The second largest power
loss is the copper loss of the secondary coil because the large
current flows on the secondary coil because the secondary DC
voltage is lower than the primary DC voltage. It is 22.3% of
the total loss.

Fig. 6. Efficiency characteristic with different secondary DC voltage. The
secondary DC voltage is changed from 200 V to 400 V. The DC-to-DC
efficiency consists of the power loss of the primary inverter, the 12 coils,
and the rectifier.

V. CONCLUSION
In this paper, the 22-kW three-phase WPT system with 12
coils has been developed for a rapid charger of EVs. The
proposed WPT system transmits power through 12-solenoid
coils, which are circularly placed. The pairs of two coils,
which are connected in series, in the common phase are placed
opposite. Thus, the radiation noise will be canceled out at themeasuring point of emission. However, not only the proposedsystem has a problem, but also the WPT system with multiple
coils. The magnetic interference among the coils degrades thetransmission efficiency. In the proposed WPT system, themagnetic interference is canceled out by using three-phasetransmission and adjusting the relation between the coils.
Specifically, the coupling among the coils is adjusted tosatisfy kan = kbn + kcn. The proposed 22-kW WPT system istested with a constant voltage load. The maximum efficiencyof the proposed system from the primary DC side to thesecondary DC side is 91.1% with the constant output voltage
of 400 V when the output power is 22 kW. The power loss at
the rated power is dominated by the conduction loss of the
primary inverter. It accounts for 26.9% of the total loss.
In future work, radiation noise emission from the proposed
three-phase WPT system will be measured. Moreover, systemefficiency will be improved by the optimization of the system.

Fig. 7. Power loss separation of 22-kW prototype.

[3]

[4]

[5]

[6]

REFERENCES
[1]

[2]

R. Bosshard, J. W. Kolar, J. Mühlethaler, I. Stevanović, B. Wunsch and
F. Canales, “Modeling and η-α−Pareto Optimization of Inductive
Power Transfer Coils for Electric Vehicles,” IEEE Journal of Emerging
and Selected Topics in Power Electronics, Vol. 3, No. 1, pp. 50-64(2015)
F. Musavi, M. Edington and W. Eberle, “Wireless power transfer: A
survey of EV battery charging technologies,” 2012 IEEE Energy
Conversion Congress and Exposition (ECCE), pp. 1804-1810 (2012)

[7]

[8]

S. Li, C. C. Mi, “Wireless Power Transfer for Electric Vehicle
Applications,” IEEE Journal of Emerging and Selected Topics inPower Electronics, Vol. 3, No. 1, pp. 4-17 (2015)
G. Lovison, T. Imura, H. Fujimoto, Y. Hori, “Secondary-side-only
Phase-shifting Voltage Stabilization Control with a Single Converter
for WPT Systems with Constant Power Load,” IEEJ Journal ofIndustry Applications, Vol. 8, No. 1, pp. 66-74 (2019)
V. Doan, H. Fujimoto, T. Koseki, T. Yasuda, H. Kishi, T. Fujita,“Simultaneous Optimization of Speed Profile and Allocation ofWireless Power Transfer System for Autonomous Driving Electric
Vehicles,” IEEJ Journal of Industry Applications, Vol. 7, No. 2, pp.
189-201 (2018)
K. Kusaka, J. Itoh, “Development Trends of Inductive Power Transfer
Systems Utilizing Electromagnetic Induction with Focus on
Transmission Frequency and Transmission Power,” IEEJ trans. on
Industry Applications, Vol. 6, No. 5, pp. 328-339 (2017)
T. Koyama, K. Umetani, E. Hiraki, “Design Optimization Method for
the Load Impedance to Maximize the Output Power in Dual
Transmitting Resonator Wireless Power Transfer System,” Vol. 7, No.
1, pp. 49-55 (2019)
G. Lovinson, D. Kobayashi, M. Sato, T. Imura, Y. Hori, “Secondaryside-only Control for High Efficiency and Desired Power with Two
Converters in Wireless Power Transfer Systems,” Vol. 6, No. 6, pp.
473-481 (2017)

[9]A. Kurokawa, M. Yano, Y. Matsuoka, “Hybrid vehicles, electricvehicles, fuel cell electric vehicles, traction motors,” Journal of Societyof Automotive Engineers of Japan, Vol. 72, pp. 90-100 (2018)[10] M. Jo, Y. Sato, Y. Kaneko, S. Abe, “Methods for Reducing Leakage
Electric Field of a Wireless Power Transfer System for Electric
Vehicles,” IEEE Energy Conversion Congress and Exposition (ECCE)
2014, pp. 1762-1769 (2014)[11] H. Kim, J. Cho, S. Ahn, J. Kim, J. Kim, “Suppression of Leakage
Magnetic Field from a Wireless Power Transfer System using
Ferrimagnetic Material and Metallic Shielding,” IEEE International
Symposium on EMC, 978-1-4673-2061-0, pp.640-645 (2012)[12] T. Campi, S. Cruciani, M. Feliziani, “Magnetic Shielding of WirelessPower Transfer Systems,” IEEE International Symposium on EMC,
15A-H1, pp.422-425 (2014)[13] K. Maikawa, K. Imai, Y. Minagawa, M. Arimitsu, H. Iwao, “Magnetic
Field Reduction Technology of Wireless Charging System,” Society ofAutomotive Engineers of Japan 2013, No. 110-13 (2013)[14] K. Inoue, K. Kusaka, J. Itoh, “Reduction in Radiation Noise Level for
Inductive Power Transfer Systems Using Spread Spectrum
Techniques,” IEEE Trans. on Power Electronics, Vol. 33, No. 4 pp.
3076-3085 (2017)[15] K. Kusaka, K. Inoue and J. Itoh, “Radiation noise reduction using
spread spectrum for inductive power transfer systems considering
misalignment of coils,” 2017 IEEE Energy Conversion Congress andExposition (ECCE) 2017, pp. 5507-5514 (2017)[16] K. Inoue, K. Kusaka and J. Itoh, “Reduction on radiation noise level
for inductive power transfer systems with spread spectrum focusing on
combined impedance of coils and capacitors,” 2016 IEEE Energy
Conversion Congress and Exposition (ECCE), pp. 1-8 (2016)[17] T. Shijo, K. Ogawa, M. Suzuki, Y. Kanekiyo, M. Ishida, S. Obayashi,“EMI Reduction Technology in 85 kHz Band 44 kW Wireless Power
Transfer System for Rapid Contactless Charging of Electric Bus”,
IEEE Energy Conversion Congress & Expo 2016, No. EC-0641 (2016)[18] Y. Tanikawa, M. Kato, T. Imura, Y. Hori, “Experiment of magnetic
resonant coupling three-phase wireless power transfer,” 2013 WorldElectric Vehicle Symposium and Exhibition (EVS27), pp. 1-8 (2013)[19] M. Kim, H. Kim, D. Kim, Y. Jeong, H. Park and S. Ahn, “A ThreePhase Wireless-Power-Transfer System for Online Electric Vehicles
With Reduction of Leakage Magnetic Fields,” IEEE Transactions on
Microwave Theory and Techniques, Vol. 63, No. 11, pp. 3806-3813(2015)[20] C. Song, H. Kim, D. H. Jung, K. Yoon, Y. Cho, S. Kong, Y. Kwack, J.
Kim, “Three-phase magnetic field design for low EMI and EMF
automated resonant wireless power transfer charger for UAV,” 2015
IEEE Wireless Power Transfer Conference (WPTC), pp. 1-4 (2015)[21] K. Liu, F. C. Y. Lee, “Zero-voltage switching technique in DC/DC
converters,” IEEE Transactions on Power Electronics, Vol. 5, No. 3,
pp. 293-304 (1990)[22] R. L. Steigerwald, “A Comparison of Half-Bridge Resonant Converter
Topologies,” IEEE Trans. on Power Electronics, Vol. 3, No. 2, pp. 174182 (1988)[23] Y. Sohn, B. H. Choi, E. S. Lee, G. C. Lim, G. Cho, C. T. Rim, “General
Unified Analyses of Two-Capacitor Inductive Power Transfer
Systems: Equivalence of Current-Source SS and SP Compensations,”
IEEE Transactions on Power Electronics, Vol. 30, No. 11, pp. 60306045 (2015)