Radiative Noise Reduction Techniqueusing 12 Coils Suitable forHigh-power Inductive Power Transfer
Keisuke Kusaka
Dept. of Electrical, Electronics andInformation Engineering
Nagaoka University of Technology
Nagaoka, Niigata, Japan
kusaka@vos.nagaokaut.ac.jp

Keita Furukawa
Dept. of Electrical, Electronics andInformation Engineering
Nagaoka University of Technology
Nagaoka, Niigata, Japanarcher_FK@stn.nagaokaut.ac.jp

Abstract—High power inductive power transfer system forelectrical vehicles is highly requested today in order to shorten
a charging time of an onboard battery. One of the difficulties
of increasing power is radiation noise. The radiation noise may
greatly increase as transmission power increase. The radiationnoise must satisfy a regulation, which is/will be legislated. In
this paper, an inductive power transfer system with 12 solenoid
coils is proposed to suppress the radiation noise. The proposed
12-coils IPT system has six coils on a primary side and six coils
on a secondary side. The six coils on the primary side arecircularly placed in every 60 degrees in order to cancel out theradiation noise and magnetic interference among the six coils.
The two opposed solenoid coils cancel out the radiative noise
because these coils generate flux in the opposite direction.
Moreover, the magnetic coupling among six coils with threephase equilibrium current cancels out the magnetic
interference, which degrades a power factor of the inductive
power transfer system. The six coils on the secondary side areplaced in the same manner. The proposed IPT system is tested
with a 2.9-kW prototype. The radiative noise on thefundamental frequency is suppressed by 1/6 using the proposed
12-coil system. Moreover, the calculated power loss shows good
agreement with the measured power loss with an error of
7.3%.
Keywords—inductive power transfer, wireless power transfer,
magnetic interference, three-phase transmission

I. INTRODUCTION
In recent years, inductive power transfer (IPT) systems
for electrical vehicles (EVs) have been actively studied [1-4].
The IPT systems transmit power using magnetic coupling
between coils without an electric cable. Thus, the IPT
systems will improve the usability of the EVs because usersof EVs do not need to connect the wire when they charge an
onboard battery. Currently, the development of the IPT
system with an input power of 3.3 kW or 7.7 kW are ongoingtowards a commercial realization for the light-duty vehicles
with a standard, which will be published by ISO/IEC [5–6].
In the standard, the IPT system is classified by the input
power. The IPT system with an input power of less than 3.3
kW and 7.7kW are WPT1 and WPT2, respectively.
Meanwhile, the amount of the onboard battery of EVs
tends to increase in order to extend a travel distance. As a

Jun-ichi Itoh
Dept. of Science of Technology
Innovation
Nagaoka University of Technology
Nagaoka, Niigata, Japan
itoh@vos.nagaokaut.ac.jp

wired fast charging system, an output power of 50 kW or
more is required. Additionally, the power of the wired
charger is planned to be increased to 150 kW or 350 kW in
CHAdeMO. The output power of the IPT system should be
increased for the quick charge with considering the trend ofincreasing power of the EV charger. As high power IPT
systems, WPT3, WPT4, and WPT5 will be standardized [5–6]. The maximum input power Pin of WPT 3 is 7.7 kW < Pin≤ 11 kW, WPT4 is 11 kW < Pin ≤ 22 kW, and WPT5 is Pin >22 kW will be standardized.
One of the challenges of increasing power is a radiationnoise, which is emitted by the transmission coils. The
magnetic flux from the coil is in proportional to the current
on the coils. Then, the radiation noise will increase as theoutput power increase. The radiation noise may cause themalfunction of electrical equipment and jam a radio
communication. Thus, the radiation noise from the IPT
system must be suppressed to below the allowable level,
which will be legislated in each area or countries. Note that,
this allowable level is determined based on the CISPR’s
guideline [7].
In [8], radiation noise reduction technique using a spread
spectrum is proposed. The peak of the radiation noise on thefundamental frequency and low-order harmonic frequencies
are suppressed by changing the transmission frequency at
every period. Especially, a new spread spectrum method,which selects the transmission frequency using a biased
probability distribution, is proposed in [8]. The radiationnoise is suppressed by only changing the control of the
primary inverter. However, the effect of the suppression
performance is not enough for the high power IPT system.
Meanwhile, the radiation noise reduction technique withfour solenoid coils has been proposed [9]. The pair of the
primary coils is placed to cancel out the radiation noise. The
secondary coils are in the same manner. From the theoreticalanalysis, the maximum performance on the noise reduction is
obtained when the two coils are placed to face each other.
However, the magnetic coupling, which is also called asmagnetic interference, between the two coils causes acirculating current. It limits the freedom of the coilplacement because the circulating current increases a copper
loss.

This paper proposes a three-phase inductive powertransfer system with 12 coils. The proposed system has six
solenoid coils for the primary side and six solenoid coils forthe secondary side. The six pairs of the coils on each side arecircularly placed in every 60 degrees. The opposed coils
cancel out the radiation noise because countercurrent flows
on the two coils. The opposed coils are the best placement tocancel out the radiation noise [9]. Moreover, the six coils
cancel out the magnetic interference among these coils whenthe equilibrium current flow on the coils. Due to this coilstructure, the three-phase IPT system without magnetic
interference among the coils is obtained. In other words, six
IPT systems, which are magnetically independent of eachother, is achieved. In this paper, first, a system configurationof the three-phase IPT system is explained. Next, a design
procedure to cancel out the magnetic interference is
described. Then, the experimental results about the radiationnoise and power loss distribution with a 2.9-kW prototype isshown. Finally, a 22-kW IPT system is designed for WPT5.
II. PROPOSED IPT SYSTEM WITH 12 COILSIn this chapter, the proposed three-phase IPT system is
introduced. The proposed system achieves cancellation of theradiation noise and magnetic interference among 12 coils bya special coil structure and three-phase equilibrium current.
A. System Configuration
Figure 1 shows the three-phase inductive power transfer
system. The proposed system consists of the three-phase
inverter, resonant capacitors, six primary coils, six secondary
coils, and three-phase rectifier. The three-phase inverter isoperated in square wave operation at around 85 kHz. The
phase difference among the phase are 120 degrees.
The resonant capacitors are inserted to the output of theinverter and the input of the rectifier on each phase in series
for compensation of a reactive power, which is caused byweak magnetic coupling between the primary coil and thesecondary coil. This topology is as same as the series-series
compensation [11], which is typically used for a single-phase
IPT. The design of the capacitors is explained in the next
section.
The six primary and six secondary coils are connected instar-star connection. Power is transmitted through the threephase windings. Note that, the coils can be connected in stardelta, delta-star, and delta-delta connection. However, theexplanation of these connections is omitted in this paper due
to the page limitation. In this system, the two coils LxnA and
LxnB are connected in series in the phase, where suffix “x”
represent “u,” “v,” or “w,” and n is “1” or “2” forrepresenting the primary or secondary side. These seriesconnected coils cancel out the magnetic raditaion because
these coils are placed in the opposite. However, the coils LxnA
and LxnB, which are magnetically coupled, causes themagnetic interference. The magnetic interference should becanceled out with the coil structure because the magnetic
interference degrades the power factor reflected from theinverter output. The cancellation technique will be explainedin the next section.

Fig. 1. Three-phase inductive power transfer system with six primarycoils and six secondary coils. The six coils in each side are connected instar winding.
iw1

w1

Lv1B
Lw1A

Ma1

iu1

Lu1B
Mc1
Lv1A

Mb1

u1

Lu1A
Lw1B

iv1

v1

(a) Primary coils(b) Secondary coils

Fig. 2. Coil placement and connection.

Fig. 3. Mechanism of radiation noise reduction with two pairs ofsolenoid coils.

Figure 2 shows the placement of the 12 coils and theconnection. In the proposed system, the solenoid-coils are
circularly placed. The secondary coils are put onto theprimary coils across the air gap. The power is transmitted
from the bottom coils to the upper coils, e.g., Lu1A to Lu2A
with a mutual coupling M.

 vu1 A   L1Ys  vu1B   M c1
 vv1 A   M b1
 
 vv1B   M a1
v  M w1 A   b1
 vw1B   M a1
v  =  M
 u2A  
 vu 2 B   0  vv 2 A   0
 vv 2 B   0  vw 2 A   0
v   0
 w2 B  

M c1

M b1

M a1

M b1

M a1

M

0

0

0

0

L1Ys
M a1

M a1
L1Ys

M b1
M c1

M a1
M b1

M b1
M a1

0
0

M
0

0
M

0
0

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
0

M a1
0

M b1
0

M c1
0

L1Ys
0

0
L2Ys

0
M c2

0
M b2

0
M a2

0
M b2

M
0

0
M

0
0

0
0

0
0

M c2
M b2

L2Ys
M a2

M a2
L2Ys

M b2
M c2

M a2
M b2

0
0
0

0
0
0

M
0
0

0
M
0

0
0
M

M a2
M b2
M a2

M b2
M a2
M b2

M c2
M b2
M a2

L2Ys
M a2
M b2

M a2
L2Ys
M c2

Fig. 3 shows the principle of the radiation noise
reduction with two coils. The two coils on common phasesuch as Lu1A and Lu1B are differentially connected and placedin the opposite. For this reason, the flux, which is generated
from the coils, is the opposite direction and the same flux
amplitude. Thus, the radiation noise is canceled out at themeasuring point, which is typically 10 m from the system [7].
The principle of the cancellation is explained in detail in [9–10].
The remaining two pairs of the coils on phase-v andphase-w are placed in every 120 degrees to the coils onphase-u. In the IPT system with multiple coils, the magnetic
interference occurs. In this system, there are three kind ofmagnetic interference Ma, Mb, and Mc due to the symmetricalplacement of coils. Ma is the mutual coupling between theadjacent coils, e.g., Lu1A and Lv1B. Mb is the mutual coupling
between the coils apart from 120 degrees, e.g., Lu1A and Lv1A.
Mc is the mutual coupling between the opposed coils, e.g.,
Lu1A and Lu1B. By adjusting these magnetic interference, theeffect of the magnetic interference is canceled out. In otherwords, six single-phase IPT systems, which are magneticallyindependent of each other, with radiation noise reduction
capability are achieved.
B. Cancellation of Effect by Magnetic Interference
In this subsection, the cancellation method of the effectby the magnetic interference among the coils Ma, Mb, and Mc
is explained. As shown in the previous subsection, theproposed system has three kinds of the magnetic interference.
These magnetic interference increase circulating current
among the coils. The effect of the magnetic interference
should be suppressed since the circulating current decrease
the transmission efficiency.
Equation (1) represents the induced voltage vxn on eachcoil, where LnYs are the self-inductance of each coil, M is themain mutual coupling, and ixn are the current on each coil.
The mutual coupling contributes to transmits power from theprimary side to the secondary side. Focusing on the first lineof (1), the induced voltage on the coil Lu1A is expressed by

diu1
di
di
+ M u 2 − M c1 u1
dt
dt
dt
di
di
di 


 di
− M a1  v1 + w1  + M b1  v1 + w1 
dt 
dt 
 dt
 dt

vu1 A = L1Ys

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)

In (1), the first term is the induced voltage by the selfinductance, the second term is the induced voltage by themutual inductance, which contributes transmitting power.
Although, the third to fifth terms in (1) is the induced voltage
by the magnetic interference. In other words, the effect of themagnetic interference can be zero if the sum of the third to
fifth terms is zero.
The primary current can be expressed as (3) when thethree-phase current is an equilibrium where ω is the angulartransmission frequency, Im1 is the amplitude of the primary
current.

iu1 = I m1 sin ωt
2  
 iv1 = I m1 sin  ωt − π  
3 



4 

 iw1 = I m1 sin  ω t − π  
3  


(3)

Substituting (3) into (2), (2) is transformed to

vu1 A = L1Ys

diu1di
+ M u2
dt
dt


2 

− ωI m M c1 cos ωt + (M a1 − M b1 ) cos ωt − π 
3 



(4).



(M a1 − M b1 )cos ωt − 4 π 

3 

From (4), the condition for the cancellation of themagnetic interference is derived. The induced voltage causedby the magnetic interference is zero when

M a1 = M b1 + M c1

(5)

is satisfied. The self-inductance of the primary coils arecommon. Thus, (5) can be expressed as(2).

k a1 = k b1 + k c1

(6).

Therefore, the unwanted induced voltages, which arecaused by the magnetic interference is canceled by adjusting
the magnetic coupling among the primary coils.
In the same manner, the cancellation condition on thesecondary side can be introduced from the seventh line of (1)as(7)

k a 2 = kb 2 + kc 2
7).

The transmission coils should be designed to satisfy (6–Figure 4 shows the simulated effect of the magnetic
interference on the power factor reflected from the inverter
output. The dotted line in these graphs is the cancellation
condition of the magnetic interference calculated by (6–7).
When the three magnetic interferences are negligible, thepower factor is unity. However, the power factor greatly
degrades when only one or two magnetic interference
increase. However, if the three magnetic interference
increases along the dotted line, the power factor remain high.
From these graphs, the validity of the cancellation conditionof the magnetic interference is confirmed.
III. DESIGN AND ANALYSIS OF THREE-PHASE IPT SYSTEM

L2Ys = Lx 2 A = Lx 2 B =

2
3 VDC
2
π 2 kω P(10).

Then, the primary inductance is determined. The voltage
gain between the primary DC voltage and the secondary DC
voltage at the resonant frequency is obtained as(11).

 VDC 2 
L2Y

=
L1Y
 VDC1 

From the voltage gain, the primary inductance should bedesigned by

L1Y = Lu1 = Lv1 = Lw1 =

2
6 VDC
1
π 2 kω P(12).

In the same manner to the secondary side, because theinductance of each coil is half, the inductance is

L1Ys = Lx1A = Lx1B =

2
3 VDC
1
π 2 kω P(13).

A. Parameter Design
The proposed IPT system is designed in this chapter
under the assumption that the magnetic interference iscanceled out by satisfying (6–7).
First, the equivalent AC resistance should be calculated.
The AC resistance Req is derived by expanding a derivation
method of an equivalent AC resistance of a single-phaserectifier described in [12]. The equivalent resistance of thecapacitor-input three-phase rectifier is calculated by

8 (VDC 2 2 )6 V2Req = 2= 2 DC 2
π
P3
π P
2

(8),

where VDC2 is the secondary voltage at rated power and P is
the rated output power. Note that the resistance is for one
phase.
Next, the secondary inductance is designed in order tomaximize transmission efficiency. From the impedance
matching perspective, the impedance of the secondary
inductance should be equal to the load, i.e., AC equivalent
resistance derived by (8). Note that the winding resistance isignored for the simplicity. Thus, the secondary inductancefor each phase is calculated by

L2Y = Lu 2 = Lv 2 = Lw2 =

2
6 VDC
2
π kω P2(9),where k is the coupling coefficient at the nominal position, ωis the angular transmission frequency. In the proposedsystem, two transmission coils are connected in series ineach phase. Thus, the inductance of each coil is expressed by(a) kb = 0

Power factor (kb = 0.1)

0.4

cos =
0.95–1
0.3

0.2

0.9–0.95
0.85–0.9
0.8–0.85
0.75–0.8

0.1

0.7–0.75

ka = kb + kc

0

0

0.1
0.2
0.3
Coupling coefficient kc

0.4(b) kb = 0.1

Fig. 4. Effect of the magnetic interference on power factor.

After that, the resonance capacitors are designed. The
resonance capacitors are connected in series to the output of
the inverter and the input of the rectifier. The resonancecapacitors are necessary to compensate the reactive power,
which is caused by the weak magnetic coupling between the
primary and the secondary side. The resonance capacitors aredesigned to resonate with the inductance on each phase at thetransmission frequency. Consequently, the resonancecapacitors on the primary side and the secondary side arecalculated by

Cu1Y = Cv1Y = C w1Y =

1
ω L1Y
2

(14),

isw _ off = 2 I1 sin ωTd(19),

where Td is the dead time.
by

The turn-off loss of the primary MOSFETs are calculated

P1_ swoff = 6Eoff (isw _ off )VDC1
× f swVn(20),

where Eoff is the turn-off energy, and Vn is the drain-source
voltage when the turn-off energy is measured.
2) Secondary Rectifier
The secondary current is calculated by

C u 2Y = C v 2Y = C w 2 Y =

1
ω L2Y
2

(15),
I2 =

respectively. For zero-voltage switching, the transmission
frequency is adjusted to be slightly higher than the resonancefrequency.
B. Loss Analysis
The power loss of the proposed system including the
primary inverter and the rectifier on the secondary side iscalculated in this section.
1) Primary Inverter
The primary inverter is operated in square wave
operation. The output phase-voltage of the inverter isexpressed by

V1 =

2

π

VDC1

(16).

When the power loss can be ignored, the output power isequal to the input power. The primary current I1 can bederived

I1 =

P
πP
=
3V1 3 2VDC1(21),

in the same manner to the primary side. The secondary
current flows on the diodes during half the period. Thus, theconduction loss of the secondary diodes are calculated by

P2 _ Dcond = 6

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

(22),

where VF is the forward voltage drop of the each diode.
3) Transmission Coils
The copper loss of the primary coils and the secondary coils
are calculated by

P1_ copper = r1 I12(23),

P2 _ copper = r2 I 22

(24),

respectively with the winding resistances r1 and r2.
(17),with the output power P.
The conduction loss of the MOSFETs on the primary
inverter is calculated by
Pp _ cond = 3rds _ on I12πP
3 2VDC 2

Then, iron loss of the transmission coils is derived fromthe magnetic flux density using the datasheet of the ferrite
core because the current on the transmission coils are
sinusoidal. In this consideration, it is assumed that themagnetic flux density on the core is uniform. The magnetic
flux density of the primary coils and the secondary coils is
calculated by(18),

where rds_on is the on-resistance of the MOSFETs.
Besides, switching loss is calculated. In IPT systems, the
MOSFETs turn-on with zero-voltage switching because theinverter output frequency is adjusted to be slightly higher
than the resonant frequency. Therefore, turn-on loss does notoccur. By contrast, turn-off loss occurs whereas the output
current is sinusoidal. Assuming that sinusoidal current flows
on the MOSFETs because the switching frequency is closed
to the resonant frequency, the simultaneous current at the
turn-off is approximately expressed by

B1 =LY 1s I1 + MI 2
N1 S1

(25),

B2 =

LY 2 s I 2 + MI1
N 2 S2(26),

where N1 and N2 are the numbers of turns of the windings,
and S1 and S2 are the cross section of the ferrite core. The
iron loss can be calculated from the datasheet based on themagnetic flux density.
IV. EXPERIMENTAL VERIFICATION
A. Experimental Setup
The proposed three-phase IPT system is tested with a 2.9kW prototype. Fig. 5 shows the transmission coils. The
transmission coils are circularly placed and adjusted the
distance among them to satisfy the cancellation condition of
the magnetic interference derived as (6–7). The magnetic
interference among the coils are ka = 0.048, kb = 0.011, andkc = 0.003. Other parameters are shown in Table I.
Figure 6 shows the definition of the rotational coil
misalignment. Fig. 7 shows the parameter change by coil
misalignment. Due to the rotation misalignment, theinductance of the coils and the coupling coefficient ischanged. In Fig. 6, the inductance and the coupling(a) Three-phase SiC inverter

coefficient are normalized based on the value without amisalignment. Note that the coupling coefficient between
primary and secondary coils, which are closest, are indicated.
When the rotational angle is 30 degrees, the inductance andthe coupling coefficient are minimal. The inductance and thecoupling coefficient are reduced by 3.4% and 52.4% at a
maximum. However, the effect of the coupling coefficient issymmetry owing to the coil placement. If the rotational
misalignment is larger than 30 degrees, the primary coiltransmits power to another phase, e.g., from Lu1A to Lv2B.
B. Experimental Results
Figure 8 shows the operation waveforms. The output
current is 30 degrees lagging from the output line voltage vuv1.
Thus it is confirmed that the proposed IPT system is operated
at the resonant frequency without the effect of the magnetic
interference. Moreover, the inverter output current isequilibrium.
Figure 9 shows the radiation noise measured at 2 m fromthe edge of the coils. The radiation noise is compared withthe conventional three-phase six coil IPT system, i.e., thecoils LxnB are removed. The noise is measured with common
conduction current on the coils. The radiation noise at thefundamental frequency is suppressed from 12.2 dBμA to 2.1
dBμA.
Figure 10 shows the DC-to-DC efficiency characteristic.
The efficiency is evaluated with changing load resistance.
Thus, the output voltage is also changed. The maximum DCto-Dc efficiency is 90.6% at an output power of 2.9 kW(input power is 3.2 kW). Note that the output power ischanged by changing the load resistance in this experiment.
It means that the output voltage is not constant.

(b) Transmission coils

Fig. 5. 2.9-kW prototype.

TABLE I. SPECIFICATION OF 2.9-KW PROTOTYPEParameters
Output power
DC voltage

Primary
SecondarySwitching frequency
Coupling coefficient

Primary

Inductance(at nominal position)

Secondary

Capacitance (average)

Wire
MOSFETs
Diodes

kW
V
V
kHz

Lu1A
Lu1B

2.9
400
400
85.5
0.25
107.1
107.9

Lv1A
Lv1B
Lw1A
Lw1B

106.9
104.2
106.5
109.7

Lu2A

108.0
107.0
109.4
107.6
106.3
105.6
16.4
16.4
250
17
17

μH
μH
μH
μH
μH
μH

VDC2
fsw
k

Lu2B
Lv2A
Lv2B
Lw2A
Lw2B

Primary
Secondary

C1
C2
Td

Primary
Secondary

N1
N2

Dead time
Number of turnP
VDC1

Litz 2UEWSTC 784 / φ0.1
SCT3030AL (ROHM)
VS-20ETF0 (Vishay)

μH
μH

μH
μH
μH
μH
nF
nF
ns
turn
turn

Figure 11 shows the comparison of the power loss
between the calculation, which is calculated using theequation in chapter III, and the measured loss. The measured
power loss is 304 W whereas the calculation is 282 W. The
calculation result shows a good agreement with theexperimental value with an error of 7.3%. The dominating
power loss is the copper loss of the primary coil and thesecondary coil. The sum of the copper loss accounts forabout 80% of the total power loss. The system efficiency will
be improved by reducing an AC resistance of the wire using
more thick litz wire.
Figure 12 shows the calculation result on the power lossof the 22-kW proposed IPT system. Because the IPT systemwith an input power of more than 22 kW will bestandardized [5–6], 22-kW IPT system is evaluated. The
Lw2B

Lv1B

Lw1B
Lv2B
Lu2B

Lu1A

Lu1B
Lu2A

Lv2A

Lv1A

Lw1A
Lw2A

Fig. 6. Assumed rotational coil misalignment.

calculation conditions are shown in Table II. The majordifference between the 2.9-kW system is the input and output
DC voltage. The input voltage is increased from 400 V to800 V. Moreover, 1200-V SiC-MOSFETs are used for highvoltage input DC. A DC-to-DC efficiency of 93.7% will beobtained.

330
320
310
300

0

340 350

10 20

30

40
50
60

290

70

280

80

270

90

260

100

250

200 190180

170 160

120
130
140
150

0
Inverter output current iu1 [20A/div]0(a) Self-inductance of primary coils

0

340 350

Inverter output current iv1 [20A/div]

10 20

30

0.95

0
40

Inverter output current iw1 [20A/div]

50
60

0.90

0

70

L / Ln = 0.85

270

90

260

100

250

110

200 190180

170 160

120
130
140
150

(b) Self-inductance of secondary coils

330
320
310
300

0

340 350

10 20

0.8

30

40
50

0.6
0.4

70

280

0.2
k / kn = 0

80
90

260

(a) Conventional system (six coils)

100

250
240
230
220
210

Fig. 8. Operation waveforms.

60

290
270

(b) Inverter output voltage, output current on the primary coil

Radiation noise dB A

240
230
220
210

4[ sec/div]

80

110

200 190180

170 160

120
130
140
150

(c) Coupling coefficientFig. 7. Indutance and coupling coefficient change due to the rotationl coil
misalignment. The indutance and coupling coefficient are normalized by
the value without a coil misalignment.

Radiation noise dB A

330
320
310
300
280

Inverter output voltage vuv1 [500V/div]

110

240
230
220
210

290

(a) Inverter output voltage, output current, rectifier input current, and output voltage(b) Proposed system (12 coils)

Fig. 9. Radiation noise at 2m from the coil edge.

DC-to-DC efficiency%

V. CONCLUSION

100
90
80
70
60
50
1000

1500

2000

2500

3000

Fig. 10. Efficiency characteristic.

In this paper, the three-phase IPT system with 12 coils
for radiation noise reduction is proposed. The three-phase
IPT system cancels out the radiation noise with two seriesconnected solenoid-coils. Moreover, six primary coils andsix secondary coils capable of canceling out the magnetic
interference among these coils. Thus, the proposed system
can be dealt as six independent single-phase IPT system. The
2.9-kW IPT system is demonstrated. The radiation noise at
the fundamental frequency is reduced from 12.2 dBμA to 2.1
dBμA. Moreover the power loss is analyzed. The calculation
result shows good agreement with the experimental value
with an error of 7.3%. Finally, the 22-kW IPT system withthe proposed topology is estimated. The estimated DC-to-DC
efficiency is 93.7%.
REFERENCES

Power loss W[1]Fig. 11. Power loss separation of 2.9-kW prototype.
1600
1400

Power loss W

1200
1000
800
600
400
200
0

Fig. 12. Power loss calculation of a 22-kW IPT system.

TABLE II. SPECIFICATION OF 22-KW IPT SYSTEM.
Parameters
Output power
DC voltage

PrimaryP
VDC1

Secondary

VDC2
fsw
k

Inductance(at nominal position)

Primary
Secondary

L1s
L2s

Capacitance (average)

Primary
Secondary

C1
C2

Switching frequency
Coupling coefficient

Dead time
Number of turn
MOSFETs
Diodes

kW
V
V
kHz
μH
μH
nF
nF

250
ns
6
turn
Secondary
N2
6
turn
BSM080D12P2C008 (Rohm)
DSEP2x31-12A (IXYS)
Primary

Td
N1

22
800
800
85.5
0.25
16.5
16.5
14.4
14.4

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