Three-phase Inductive Power Transfer System with12 coils for Radiation Noise Reduction
Keisuke Kusaka1* and Jun-ichi Itoh2
1 Department of Electrical, Electronics and Information Engineering, Nagaoka University of Technology, Nagaoka, Japan
2 Department of Science of Technology Innovation, Nagaoka University of Technology, Nagaoka, Japan*E-mail: kusaka@vos.nagaokaut.ac.jp

Abstract— A three-phase inductive power transfer
system with six transmission coils and six receiving coils is
proposed in this paper. The proposed system achieves areduction in radiation noise by canceling the noise using
multiple coils, which are placed opposite to each other.
However, in IPT systems with multiple coils, a magnetic
interference among the multiple coils decreases atransmission efficiency due to the occurrence of circulating
current. The proposed three-phase IPT system achieves
cancellation of the interference using six coils on each side.
First, the IPT system with the 12 coils is proposed. Then, acanceling condition of the magnetic interference among thecoils is mathematically introduced from the voltage
equation on the coils. Finally, the proposed IPT system isexperimentally demonstrated with the 3-kW prototype. The
experimental result shows that the radiation noise at thefundamental frequency is suppressed to 1/6 (from 12.2dBA
to 2.1dBA).
Keywords— inductive power transfer, wireless powertransfer, three-phase transmission

I.

INTRODUCTION

In recent years, inductive power transfer (IPT) systems
for electrical vehicles (EVs) are actively studied [1–11]
because IPT systems are capable of improving theusability of EV users. In order to put the IPT system forpractical use, a standardization of the IPT system for EVs
is ongoing [12–13]. According to the standards, which
will be published by IEC/ISO, the IPT systems will beclassified by the maximum input power into four class;the maximum input power is ≤ 3.7 kW for WPT1, themaximum input power is > 3.7 kW and ≤ 7.7 kW for
WPT2, the maximum input power is > 7.7 kW and ≤ 11
kW for WPT3, the maximum input power is > 11 kW and≤ 22 kW for WPT4, and the maximum input power is >22 kW for WPT5 [12–13].
The standardization of WPT1 and WPT2 are currently
in progress. However, the standardization of high-power
IPT systems such as WPT5 has not been discussed well
despite strong demand for higher power IPT system forheavy-duty vehicles because high-power IPT systems
may cause two problems. One of the problems is adecrease in transmission efficiency. Copper loss and ironloss on transmission coils may increase because thecurrent in the transmission coils increases [14]. The
second problem is radiation noise. The radiation noisecaused by the IPT systems must satisfy the regulations,which are published by CISPR or legislated in eachcountry or region. However, the satisfaction ofregulations becomes further difficult because theradiation noise caused by a loop coil is proportional tothe conduction current.
In previous studies, a radiation noise reduction
technique using a duplicated 44-kW IPT system forheavy-duty EVs has been proposed [14–16]. In [15], theradiation noise is canceled by duplicated transmission
coils. The two transmission coils with differential current
are placed to cancel out the leakage radiation noise.
However, the duplicated IPT system limits freedom of
coil positions because the duplicated coils cause amagnetic interference between the duplicated coils [15].
The magnetic interference decreases power factor
regarding an inverter output.
Meanwhile, three-phase IPT systems have been
proposed to increase the output power of the IPT system[17-20]. The three-phase IPT systems allow the reductionof the current on each transmission coil. Thus the threephase IPT is an effective method to reduce the copper
loss and the iron loss. However, in the conventional
three-phase IPT system, a magnetic interference among
the additional coils on each side occurs. The magnetic
interference causes a degrade the efficiency because themagnetic interference leads to circulating currents.
This paper proposes a three-phase IPT system with 12
solenoid coils, which consists of six pairs of coils placed
opposite to each other. The proposed IPT system
contributes reducing the current on the winding without abyproduct such as a magnetic interference. A
contribution of this paper is offering the newtransmission coil structure, which reduces the radiationnoise with a small effect on efficiency. In the proposedsystem, the radiation noise is canceled out by thecountercurrent in the opposite coil. Moreover, theprimary coils and the secondary coils are placed at anangle of 60 degrees to each other. Consequently, theinduced voltage caused by magnetic interference among

the 12 coils are canceled out. First, a system
configuration of the proposed IPT system is unveiled.
Next, a design procedure of the proposed three-phase IPT
system with a star-star winding and a delta-delta winding
is explained. Then, the canceling method of the magnetic
interference among the multiple coils is mathematicallyintroduced. Finally, the proposed IPT system with starstar winding is tested with a 3-kW prototype. Then, theradiation noise is assessed and compared between theconventional three-phase IPT system and the proposed
three-phase IPT system with 12 coils.

iu1

u1
Cu1Y

v1
w1

Lu2A

Lu1B
Lv1B

Lw2B

A. System Configuration
Figure 1 shows the proposed three-phase IPT systemwith 12 coils. The three-phase coils can be connectedwith not only star-star winding but also star-delta, deltastar, and delta-delta windings. In this paper, only star-starand delta-delta windings are discussed due to the page
limitation. Power is inductively transmitted through thethree-phase coils. Each phase has two coils connected inseries. These coils connected in series cancel out theradiation noise because of the countercurrent flowing on
the coils [14].
Besides, the three-phase inverter in the primary sideand the three-phase rectifier in the secondary side areused. The primary inverter output square-wave voltage ineach phase. The phase difference of each phase are 120
degrees. Furthermore, resonant capacitors are connectedto the output of the primary inverter and the input of therectifier in series. The series-series compensation
technique [21] is applied to the proposed system in orderto cancel out the leakage inductance due to the weakmagnetic coupling. Using three-phase IPT and the seriesseries compensation technique, copper loss of thewindings can be reduced because the current is shunted
into six coils. Therefore, the proposed IPT system is
effective to improve the heat dissipation of thetransmission coils.
Figure 2 shows the position and connection of thetransmission coils. In the proposed IPT system, six
solenoid coils are used in each of the primary side and thesecondary side. Fig. 3 shows the mechanism of the
reduction in radiation noise. Two transmission coils
connected in series (e.g., Luv1A and Luv1B) are
differentially connected and are placed in an opposite
place. These opposite coils cancel out radiation noise at ameasurement point, which is typically 10 m from the IPT
system [22]. Other pairs of coils are placed in an angle of
120 degrees to the coils on phase-u each other. As adrawback of the reduction in radiation noise using
opposite coils, magnetic interference occur among the six
coils on each the primary side and the secondary side. In

R

Lw2A

Lw1B

v2

Lv2B

VDC2

Cv2Y

Lv2A

Lw1A

Cv1Y

w2
Cw2Y

Cw1Y

(a) Star-star winding

Cu1D

v1

Lwu1B

Luv1A
iuv1

Luv1B

Lwu2B Cu2D
vuv2

Luv2A

Lwu1A
Luv2B
Lvw2B Lvw2A

Lvw1A Lvw1B

Cv1D

VDC1

iu2

M

u1

THREE-PHASE IPT SYSTEM WITH 12 COILSIn this chapter, the three-phase IPT system is proposed.
The proposed IPT system achieves radiation noise
reduction using 12 coils without magnetic interference
among multiple coils. In other words, the IPT systemtransmits power with unity power factor regarding
inverter output.

Cu2Y

Lu2B

iu1

II.

u2

Lu1A

Lv1A

VDC1

iu2

M

u2

u2

R

Lwu2A

v2

Cv2D
w2

w1
Cw1D

VDC2

Cw2D

(b) Delta-delta winding
Fig. 1. Proposed three-phase IPT system with 12 coils.
iw1

iw2

w1

Lv1B

Lv2B

Lw1A

Lw2A

Ma1

iu1

Lu1B
Mc1
Lv1A

Ma2

iu2

Lu2B

u1

Mc2

Lu1A

Lw2B

Mb2

iv1

u2

Lu2A

Lv2A

Lw1B

Mb1

w2

iv2

v1

v2

(a) Star-star winding
w1

w2

Lvw1B
Lwu1A

Ma1

Luv1B
Luv1A
Lwu1B

iuv2

Mb1
Lvw1A

v1
ivw1

Ma2

Luv2B

u1
Mc1

iuv1

Lvw2B
Lwu2A

v2

u2
Mc2

Luv2A
Lwu2B

Mb2
Lvw2A

iwu1
ivw2
iwu2
(b) Delta-delta winding
Fig. 2. Placement and connection of coils.

x
Lu2B or Luv2B
Lu1B or Luv1B

fuvB

fuvA

Lu2A or Luv2A
Lu1A or Luv1A
ferritewindings

Fig. 3. Mechanism of reduction in radiation noise using solenoid coils.

order to cancel out the magnetic interference, the position
relationship and figure of the coils have to be adjusted.
The design method of the coils will be explained in
section C.
B. Mathematical Expression of Three-phase Coils
Equations (1) and (2) represents the induced voltage of
12 coils on the star-star winding system and the delta- 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  vv 2 B   0
 vw 2 A   0  vw 2 B   0

M c1
L1Ys
M a1
M b1
M a1
M b1
0
M
0
0
0
0

M b1
M a1
L1Ys
M c1
M b1
M a1
0
0
M
0
0
0

M a1
M b1
M c1
L1Ys
M a1
M b1
0
0
0
M
0
0

M b1
M a1
M b1
M a1
L1Ys
M c1
0
0
0
0
M
0

M a1
M b1
M a1
M b1
M c1
L1Ys
0
0
0
0
0
M

M
0
0
0
0
0
L2Ys
M c2
M b2
M a2
M b2
M a2

 vuv1 A   L1Ds  vuv1B   M c1
v
 M
 vw1 A   b1
 vvw1B   M a1 
 v wu1 A   M b1
 v wu1B   M a1
 vuv 2 A   Mv
  0
 uv 2 B   vvw2 A   0  vvw2 B   0
 v wu 2 A   0 
 v wu 2 B   0

M c1
L1Ds
M a1
M b1
M a1
M b1
0
M
0
0
0
0

M b1
M a1
L1Ds
M c1
M b1
M a1
0
0
M
0
0
0

M a1
M b1
M c1
L1Ds
M a1
M b1
0
0
0
M
0
0

M b1
M a1
M b1
M a1
L1Ds
M c1
0
0
0
0
M
0

M a1
M b1
M a1
M b1
M c1
L1Ds
0
0
0
0
0
M

M
0
0
0
0
0
L2 Ds
M c2
M b2
M a2
M b2
M a2

0
M
0
0
0
0
M c2
L2Ys
M a2
M b2
M a2
M b2
0
M
0
0
0
0
M c2
L2 Ds
M a2
M b2
M a2
M b2

0
0
M
0
0
0
M b2
M a2
L2Ys
M c2
M b2
M a2
0
0
M
0
0
0
M b2
M a2
L2 Ds
M c2
M b2
M a2

delta winding system, respectively where L1Ys is the selfinductance of each coil on the star-star winding, L1Ds isthe self-inductance of each coil on the delta-delta
winding.
The mutual inductance M represents the mutual
inductances between the primary side and the secondaryside, which contributes to the power transmission, e.g.,
the mutual inductance between Lu1A and Lu2A. Ma
represents the mutual inductance between the adjacentcoils, e.g., the mutual inductance between Lu1A and Lv1B.
The mutual inductance Mb represents the mutual
inductance between the coils, which are placed apartfrom 120 degrees, e.g., the mutual inductance between
Lu1A and Lw1A. The mutual inductance Mc represents themutual inductance between the opposite coils, e.g., themutual inductance between Lu1A and Lu1B. The suffix “1”indicates inductances on the primary side, and “2” means
inductances on the secondary side. Note that otherunwanted mutual inductances, e.g., the mutual inductance
between Lu1A and Lu2B are ignored in Eq. (1).
It is clear from Eq. (1) that the unwanted mutual
inductances cause an unwanted induce voltage on each of
the transmission coils. This induced voltage decreases theefficiency due to a circulating current. The cancelingmethod of these unwanted induced voltages will be
explained in the next section.
C. Cancellation of Magnetic Interference Coupling
In this section, the canceling method of unwanted
induced voltage caused by the magnetic interference is
explained. The cancellation method of the magnetic
interference coupling is explaind with the star-star

0
0
0
M
0
0
M a2
M b2
M c2
L2Ys
M a2
M b2
0
0
0
M
0
0
M a2
M b2
M c2
L2 Ds
M a2
M b2

0
0
0
0
M
0
M b2
M a2
M b2
M a2
L2Ys
M c2

0
0
0
0
M
0
M b2
M a2
M b2
M a2
L2 Ds
M c2

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 

(1)

0   iuv1  
0    iuv1 
0   ivw1  
0    ivw1  
0   iwu1 M  d   iwu1  
M a 2  dt  iuv 2 
M b 2    iuv 2  
M a 2   ivw2  
M b 2    ivw2 M c 2   iwu 2  
L2 Ds    iwu 2 

(2)winding. However, the cancellation technique can beused for the delta-delta winding.
The induced voltage on Lu1 is derived from the firstcolumn of Eq. (1),

diu1
di
di
 M u 2  M c1 u1
dt
dt
dt
(3)
 div1 diw1  div1 diw1  M a1  M b1 

dt 
dt 
 dt
 dt

vu1 A  L1Ys

The first term of the right side in Eq. (3) is the induced
voltage by self-inductance, the second term is the mutual
inductance caused by the coil on the secondary side. The
second term contributes to the power transmission. The
third to fifth terms are the induced voltage caused by theinterference coupling. If the third to fifth terms of Eq. (3)is zero, the relationship between the primary coil L1Ys andthe secondary coil L2Ys is seen as the same as atransmission with one-by-one.
Assuming the three-phase equilibrium, Eq. (3) istransformed as

vu1 A  L1Ys

diu1di
 M u2
dt
dt


2 

 I m M c1 cos t  M a1  M b1  cos t    (4),3 M a1  M b1 cos t  4  

3 

where Im is the maximum value of the primary current iu1,
which flows in each of the primary coils.

From Eq. (4), it is found that the unwanted induced
voltage does not occur when(5)

M a1  M b1  M c1

2) Secondary inductance
Maximizing a transmission efficiency, an impedanceof the secondary inductance should be equal to theimpedance of the equivalent AC resistance. By ignoring a

is satisfied.
Because the self-inductance of the primary coils L1Ys isequal to each other, the canceling condition of theunwanted induced voltage is

0.35

Coupling coefficient ka

Therefore, the unwanted coupling can be canceled bydesigning the magnetic coupling according to Eq. (6) on
the primary side.
Similarly, the unwanted coupling on the secondaryside can be canceled by designing the magnetic couplingaccording to Eq. (7).

cos
0.3 q =
0.95–10.3

0.25

ka = kb + kc

0.2

0.2

0.9–0.95
0.85–0.9
0.15
0.8–0.85
0.75–0.8
0.1

0.1

(7)

ka 2  kb 2  kc 2

0.7–0.75
0.05
0

00

0.05

0.7-0.75

00.40.4

0.1
0.2
0.1
0.15
0.2
0.25 0.3
0.3
Coupling coefficient kc

0.35

0.75-0.8

0.9-0.95

0.8-0.85

0.85-0.9(a) kb = 0

Power factor (kb = 0.1)

cos q =
0.35
0.95–1
0.3

0.3
0.9–0.95
0.85–0.9
0.25
0.8–0.85
0.75–0.8
0.2

ka = kb + kc
0.2

0.15
0.1
0.7–0.75

0.1

DESIGN OF THREE-PHASE IPT SYSTEM

0.05

In this chapter, the design method of proposed threephase IPT system with the star-tar winding and the deltadelta winding is described. In this design procedure, theinduced voltage generated by the magnetic interference
among the multiple coils can be ignored with assuming
Eq. (6–7).

2

(8).

0

0.4 0
0
0.1
0.2
0.3
0 0.05 0.10.7-0.75
0.15 0.2
0.75-0.8
0.25 0.3
0.8-0.85
0.35 0.4
Coupling coefficient kc
0.85-0.9(b) k0.9-0.95= 0.1 0.95-1
bPower factor (kb = 0.2)

0.4

0.9–0.95

cos q =
0.95–1

0.85–0.9
0.8–0.85
0.75–0.8

ka = kb + kc

0.3

Coupling coefficient ka

A. Star-star winding
This section represents the design method of the
proposed three-phase IPT system with the star-starwinding shown in Fig. 1 (a) with.
1) Equivalent AC resistance
The equivalent AC resistance is calculated from therated output voltage VDC2, and rated output power P. The
equivalent AC resistance Req can be calculated byexpanding a derivation of an equivalent AC resistance ofa single-phase rectifier to the three-phase rectifier. The
equivalent AC resistance for a certain phase is expressedby2
8 VDC 2 2 
6 VDC
2

P3
2
2 P

0.95-1

0.4

0.4

Coupling coefficient ka

Figure 4 shows the simulated relationship between the
inverter output power factor and the magnetic
interference couplings ka, kb, and kc. The dotted line on
the graph represents Eq. (6). Fig. 4 (a), (b), and (c) show
the relation when the magnetic interference coupling kb iszero, 0.1 and 0.2, respectively. When the magnetic
interference coupling is small enough to be negligible,
the power factor closes to unity. The power factor
decreases when one of the interference coupling increases.
However, the power factor remains at high even when theinterference couplings increase under as long as Eq. (6) issatisfied. On the other words, these results show that themagnetic interferences among the transmission coils do
not affect the power factor when the relation of theinterference coupling fits on Eq. (6).

Req 

0.4

(6).

k a1  k b1  k c1

III.

Power factor (kb = 0)

0.4

0.2
0.7–0.75
0.1

0
0

0.1
0.2
0.3
Coupling coefficient kc

0.4(c) kb = 0.2
Fig. 4. Effect of magnetic interference coupling on power factor.

winding resistance and assuming the following resonancecondition, the secondary self-inductance for a certain
phase should be
L2Y  Lu 2  Lv 2  Lw 2 

2
6 VDC
2
 2 k P(9),

where  is the transmission angular frequency, k is thecoupling coefficient.
The secondary inductances have to be half because
two transmission coils are connected in each phase in the
proposed IPT system. Thus, inductance for a certain coilis L2Ys = L2Y / 2.
3) Primary inductance
From the voltage ratio between the input DC voltage
VDC1 and the output DC voltage VDC2 under the resonancecondition, a primary self-inductance for obtaining thedesired rated voltage is expressed by
2

2
V 
6 VDC
1
L1Y  Lu1  Lv1  Lw1  L2Y  DC1   2
 k P VDC 2 

(10),

where the primary inverter is operated with a squarewave operation. In the proposed IPT system, thetransmission coils are divided and connected in series.
The inductances represented by Eq. (10) have to be half.
Thus, the inductance of primary coil is L1Ys = L1Y / 2.
4) Resonance capacitors
Besides, the resonance capacitors are connected to theprimary inverter and the rectifier in series. The resonancecapacitors are designed to resonate with the selfinductance at the transmission angular frequency . Thus,
the resonant capacitors can be designed as
(11),

1
 L2Y

(12).

2

Note that, the operating frequency is slightly adjusted
to achieve a zero-voltage switching of the MOSFETs in
the inverter.
B. Delta-delta winding
This section describes the design method of the
proposed three-phase IPT system with the delta-delta
winding shown in Fig. 1 (b).
The equivalent resistance for the star connection is assame as the star-star winding.
1) Secondary inductance
The secondary inductance L2D should be equal to theimpedance of the equivalent AC resistance. Thus thesecondary inductance with a star winding is expressed by
L2Y 

Req
k



2
6 VDC
2
 k P
2

L2 D  Luv 2  Lvw2  Lwu 2  3L2Y (13)with assuming small winding resistance under the
resonance condition.

2
18 VDC2
 2 k P(14).

The inductance expressed in Eq. (14) have to be halfin the proposed IPT system because the transmission
coils are divided and connected in series.
2) Primary inductance
The primary inductance is calculated by Eq. (15) fromthe voltage ratio between the primary DC voltage and thesecondary DC voltage when the primary inverter isoperated in square-wave operation mode.
2

2
V 
18 VDC
1
L1D  Luv1  Lvw1  Lwu1  L2 D  DC1   2
V

k

P
DC
2



(15)

The primary inductances also have to be half.
3) Resonance capacitors
The resonance capacitors are connected in series to the
output of the inverter and the input of the rectifier. The
resonance capacitors are designed to resonate with theprimary and the secondary inductance at the transmission
angular frequency .
Cu1D  Cv1D  Cw1D 

3

 2 L1D

Cu 2 D  Cv 2 D  Cw 2 D IV.

1Cu1Y  Cv1Y  Cw1Y  2
 L1Y
Cu 2Y  Cv 2Y  Cw 2Y 

The secondary inductance expressed by Eq. (13) havebe transformed into the inductance for the star-starwinding. The secondary inductance is expressed by3
 L2 D
2

(16)

(17)

EXPERIMENTAL VERIFICATION

Figure 5 and Table I show the prototype and its
specifications. In this paper, the proposed IPT systemwith the star-star winding is experimentally demonstrated.
The magnetic interference couplings are ka = 0.048, kb =0.011, and kc = 0.003. The each transmission coils haveferrite plates of 245 × 215 × 10 mm. The ferrite plates are
put into boxes made from acrylic. As windings for thetransmission coils, a litz-wire is used.
Figure 6 shows the operation waveforms with the 3kW prototype. Fig. 6 (a) shows the inverter output
voltage vuv1, output current iu1, rectifier input current iu2
and output voltage VDC2. Note that the line voltage vuv1 as
the inverter output voltage is observed. Thus, the inverter
output current lags 30 degrees with respect to the inverter
output line-volage. It means that the power factor
correction using resonance is achieved despite themagnetic interference coupling. Moreover, it isconfirmed that power is transmitted from the primaryside to the secondary side.
Figure 7 shows the voltage gain between the primary
DC voltage and the secondary DC voltage. The dottedline is the theoretical curve, which is calculated byignoring the magnetic interference among the coils. The
experimental results almost agree with the theoreticalcurve except for low load region. In the low-load area,

Inverter output voltage vuv1 [500V/div]0
Inverter output current iu1 [20A/div]0
Rectifier input current iu2 [20A/div]0
Output voltage VDC2 [250V/div]0(a) Three-phase inverter
Lw2A

Lv2B

4[sec/div](a) Inverter output voltage, output current, rectifier input current andoutput voltage
Inverter output voltage vuv1 [500V/div]

Lu2B

Lw1A

Lu2A

Lv1B

0
Inverter output current iu1 [20A/div]

Lu1A

Lu1B

0
Inverter output current iv1 [20A/div]

Lw2B

0

Lv2A

Inverter output current iw1 [20A/div]0(b) Transmission coils
Fig. 5. 3-kW prototype.
Table 1. Specifications of the prototype.
Item
SymbolValue
Primary DC voltage
VDC1
400
Secondary DC voltage
VDC2400
Rated powerP
3.0
Transmission frequencyf85.6
Coupling coefficient
k
0.26
ka
0.048
Interference coupling
kb
0.011
kc
0.003
Primary inductance
L1Ys
106
Secondary inductance
L2Ys
106
Primary capacitance
Cu1Y, Cv1Y, Cw1Y
16.5
Secondary capacitance
Cu2Y, Cv2Y, Cw2Y
16.5
MOSFETs
SCT3030AL (ROHM)
DiodesVS-20ETF06-M3 (Vishay)

4[sec/div](a) Inverter output voltage and output currents on the primary side
Fig. 6. Operation waveforms.

1.2
V
V
kW
kHz

uH
uH
nF
nF

the hard switching occurs because the current during thedead time is not enough to achieve zero-voltage
switching. Due to the hard switching, the voltage gain isslightly different to the theoretical curve.
Figure 8 shows the measured efficiency characteristicfrom the primary DC voltage to the secondary DC
voltage. In fact, the efficiency takes into account thepower loss in the inverter and the rectifier. The maximum
efficiency reaches 91%. The reason why the efficiency is
not so high because the transmission coils of the
prototype are a scaled-down model. Thus, the efficiency
is expected to be improved in an actual application.
Figure 9 shows the measuring point of the radiationnoise. The radiation noise is measured at 2 m from the1
Voltage gain VDC2 / VDC1

Lv1A

Lw1B

0.8

Experiment

0.6
Theoretical curve(magnetic interference isunconsidered)

0.4
0.2
0

10

20

30

40

50

60

Load R [W]

Fig. 7. Voltage gain between primary voltage and secondary voltage.

coil edge. The radiation noise is measured in a shieldedroom. As the conventional system, three-phase IPT
system with three coils is tested for the comparison on
the radiation noise. When the conventional method istested, the six coils (coil B in Fig. 2) are removed. The
transmission coils of phase-v and phase-w are placed in
an angle of 120 degrees to the coils on phase-u.
Moreover, the radiation noise is compared under thesame current condition. In order to remove the effect of
the radiation noise from the inverter, the inverter isplaced out of the shield room.
Figure 10 shows the radiation noise. Fig. 10 (a) is theradiation noise generated by the conventional three-phase
IPT system with six coils where the coils B in Fig. 2 are

58

100
90

38

Radiation noise [dBA]

DC-to-DC efficiency h [%]

48

80
70
60

28
18

12.2dBA

8
-2
-12
-22

50
1000

1500

2000

2500

3000

-32

Output power Pout [W]100Fig. 8. Efficiency characteristic.

200

300

400 500 600 700
Frequency [kHz]

800

900

(a) Conventional three-phase IPT system with six coils.
58
48
38

Radiation noise [dBA]

Probe

2m

28
18
8

2.1dBA

-2
-12
-22

(a) For conventional three-phase IPT system

-32
100
Probe

200

300

400 500 600 700
Frequency [kHz]

800

900

(a) Proposed three-phase IPT system with 12 coils.
Fig. 10. Radiation noise at 2 m from the coil edge.

2m(b) For proposed three-phase IPT system
Fig. 9. Measurement of radiation noise.

removed. Fig. 10 (b) is the radiation noise generated by
the proposed IPT system with 12 coils. The measurement
distance and conditions are same as those in Fig. 10 (a).
In both the experiments, common transmission coils areused. In order to have a fair comparison, the conduction
currents on each coil are adjusted to be same by adjusting
the input DC voltage. Note that the inverter and therectifier are put outside of the shielded room in order tomeasure only the radiation noise from the transmission
coils.
The radiation noise at the fundamental frequency issuppressed from 12.2dBA to 2.1dBA by the proposed
IPT system. It is confirmed that the proposed three-phase
IPT system is effective to reduce the fundamental
radiation noise. However, the radiation noise on the thirdorder harmonics of the proposed system slightly increases.
V.

CONCLUSION

This paper proposed three-phase inductive powertransfer system. The proposed system has six primarycoils and six secondary coils. The opposite coils, which
are in series and are differentially connected, cancel out
the radiation noise. Moreover, the six coils on each side
allow to cancel out the magnetic interference among themultiple coils. Due to the cancel of the magnetic
interference, the IPT system can be operated with anunity power factor regarding the inverter output.
In this paper, the effect of the magnetic interference
and its cancellation method is mathematically introduced.
Then, the proposed scaled-mode was experimentallytested with the 3-kW prototype. The experimental resultsshows that the maximum DC-to-DC efficiency is 91%.
Finally, the radiation noise is evaluated and compared tothe conventional three-phase IPT system with six coils.
The proposed method suppresses the radiation noise,
which is measured at 2 m from the coil edge, from
12.2dBA to 2.1dBA on the fundamental frequency.
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