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IEEJ Transactions on Industry ApplicationsVol.134 No.1 pp.1–8 DOI: 10.1541/ieejias.134.1
Paper
Radial Force Control of IPMSM
Considering Fundamental Magnetic Flux Distribution
Masato Kanematsu∗a) Student Member, Takayuki Miyajima∗ Student Member
Hiroshi Fujimoto∗ Senior Member, Yoichi Hori∗ Fellow
Toshio Enomoto∗∗ Non-member, Masahiko Kondou∗∗ Non-member
Hiroshi Komiya∗∗ Non-member, Kantaro Yoshimoto∗∗ Member
Takayuki Miyakawa∗∗ Non-member
(Manuscript received January 23, 2011, revised March 15, 2011)
This paper proposes an optimal d-axis current to suppress the 2nd radial force, which is caused by thefundamental permanent flux. Under the no-load condition, the flux distribution is approximated in order tocalculate the radial force. Considering the cyclic nature of 3-phase, the optimal d-axis current reference tosuppress the 2nd radial force is derived. Simulations and experiments under both load and no-load conditionsare performed to demonstrate the validity of the proposed optimal d-axis current reference.
Keywords: IPMSM, maxwell stress, radial force control, radial force, flux distribution, vibration suppression
1. Introduction
IPMSMs (Interior Permanent Magnet SynchronousMotors) are widely applied in many industrial applica-tions. In these applications, it is essential to reduce thenoise and vibration. Compared with other types of elec-trical machine, such as induction motor and switchedreluctance motor, IPMSMs are relatively quiet. How-ever, in applications, such as industrical servos, con-sumer products, and automotive drives, acoustic noiseand vibration are important issues. Quietness enhanceshigh commodity value of IPMSMs. Moreover the lessweight IPMSMs become because of efficient vehicle (1),the larger noise and vibration have been caused.
In IPMSMs, a strong electromagnetic force exists be-tween the rotor magnets and the stator teeth, havingcomponents in both the tangential and radial directions.The tangential force results in torque ripple. The ra-dial force which is stronger than the tangential force in-duces mechanical deformation and vibration of the sta-tor. The causes of noise and vibration in PMSMs areinvestigated in (2). In (3) the relationship between skewand radial force vibration is examined. In (4) and (5),vibration caused by radial force is examined with struc-tual analysis. In (6), with FEA (Finite Element Anal-ysis) and experiments, a detailed study of the vibra-tion is performed. Some relationships between vibration
a) Correspondence to: Masato Kanematsu. E-mail:[email protected]
∗ The University of Tokyo, 5-1-5 Kashiwanoha KashiwaChiba, Japan
∗∗ Nissan Motor Co., LTD., 1-1 Morinosato Aoyama AtsugiKanagawa, Japan
and structual characteristics are shown in (7) and (8).Some papers such as References (7) and (8) propose ra-dial force reduction using structural changes. On theother hand, few methods to suppress vibration usingcurrent control are proposed. Reference (9) proposes ra-dial force reduction with current control in condition ofno-tooth effect. However this method is not versatilebecause the Reference (9) does not consider tooth effectand the relationship between current reference and ra-dial force is obscure. Therefore, simpler methods whichpredict and reduce amplitude of radial force are required.
The radial force which is caused by the permanentmagnet is mainly electrical 2nd order. This paper pro-poses an optimal d-axis current to suppress the 2nd ra-dial force. The flux distribution on the no-load conditionis approximated as a series of rectangle. Based on theapproximate model, the radial forces on each tooth arecalculated. Considering cyclic nature of 3-phase, the op-timal d-axis current is derived. Second, q-axis influenceis investigated with FEA. Finally, simulations and exper-iments under load and no-load conditions are performedto demonstrate the validity of the proposed optimal d-axis current reference.
2. Model and Maxwell Stress
2.1 dq Model of IPMSM The voltage equa-tion of IPMSM in dq-axis and the motor torque T arerepresented by[
vd
vq
]=
[Lds + R −ωeLq
ωeLd Lqs + R
] [idiq
]+ ωeΨa
[01
], (1)
T = Kmtiq + Krtidiq, (2)
c© 2014 The Institute of Electrical Engineers of Japan. 1
(Masato Kanematsu et al.)
ωm =1
Js + DT, ωe = Pωm, (3)
where vd, q are the d-axis and q-axis voltages, id, q arethe d-axis and q-axis currents, Ld, q are the d-axis andq-axis inductances, R is the stator winding resistance, ωe
is the electric angular velocity, Ψa is the back EMF con-stant, Kmt := PΨa, Krt := P (Ld−Lq), P is the numberof pole pairs, and J is the rotor innertia. In this paper,2-phase/3-phase transform is absolute transformation.
2.2 The Electromagnetic Forces of IPMSMϕm refers to the stator position angle between the centerof a U-phase tooth and a point where maxwell stress isconsidered. Using the definition of Maxwell stress ten-sor, the radial and tangential electromagnetic forces arecalculated as
fr(ϕm) =B2
r (ϕm) − B2θ (ϕm)
2µ0, (4)
fθ(ϕm) =Br(ϕm)Bθ(ϕm)
µ0, (5)
where Br(ϕm) and Bθ(ϕm) are the radial and tangentialflux densities on ϕm, fr(ϕm) and fθ(ϕm) are the radialand tangential maxwell stress on ϕm. In (4), B2
θ (ϕm) ismuch smaller than B2
r (ϕm) and (4) can be approximatedas
fr(ϕm) =B2
r (ϕm)2µ0
. (6)
In this paper, the radial force FrU,rV,rW acting on thesurfaces of U, V and W-phase teeth are defined as
FrU,rV,rW =∫∫
fr(ϕm)dS, (7)
where S is the internal surface area on a tooth.
3. Approximate Model of Flux Distribu-tion
In this paper, the radial force is calculated using theproposed approximate flux distribution model. In thischapter, the radial flux distribution is derived. JMAG(electromagnetic field analysis software with the FEA)produced by JSOL Corporation is applied to this analy-sis. In FEA, it is supposed that ideal sinusoidal currentflows.
3.1 Hypotheses for Approximation This pa-per considers 12 poles 18 slots IPMSM. The windingpattern is concentrated winding.
The value of Br(ϕm) is a function of id, iq, ϕm, andθe, where θe is rotor angle. Moreover, it is assumed thatthe radial flux density Br(ϕm) is expressed as
Br(ϕm) = Brid(ϕm, id) + Briq (ϕm, iq) + Brm(ϕm),
(8)where, Brid
(ϕm, id) and Briq (ϕm, iq) are the radial fluxdensities due to id and iq on ϕm, Brm(ϕm) is the ra-dial flux density which is generated by the permanentmagnet on ϕm. Firstly, in order to consider no-loadcondition, this paper neglects Briq (ϕm, iq). The effectof q-axis current is described at Section 4.4.
In this paper, the linearity between id and Brid(ϕm)
is assumed.
S
Brm(ϕm)
γSγS
(a) Magnetic flux lines
-1.5
-1
-0.5
0
0.5
1
1.5
-30 -20 -10 0 10 20 30
flux density Br[T]
Stator Position ϕm[deg]
W U V
FEM analysisapproximate model
(b) Brm(ϕm)
Fig. 1. Flux distribution Brm(ϕm) by permanentmagnet
3.2 Approximate Model of Flux Distributionby Permanent Magnet The FEA result of the fluxdistribution of the permanent magnet at θe = 0 Deg. isshown in Fig. 1(b). As shown in Fig. 1(b), the fluxdistribution is nearly flat on the U-phase tooth, but un-equal on the V-phase and the W-phase teeth. Here,the flux distribution is approximated by a rectangle.The magnetic flux passes through the area γS and nomagnetic flux passes through the other area (1 − γ)S.γ (0 < γ ≤ 1) is flux interlinkage area. It is expectedthat γ depends on the rotor structure, but this relation-ship is inevident. In this paper, γ is determined fromFEA, such that γ = 1 on U-phase and γ = 0.5 on V-phase and W-phase.
The fluxes φmU,mV,mW on a tooth surface are calcu-lated as
φmU = φ, φmV,mW = −12φ, (9)
where N is the number of commutating turns per a tooth
and φ :=√
23
Ψa
PN . Here,√
23 is the coefficient to trans-
form 2-phase into 3-phase.The flux distribution of the permanent magnet is ap-
proximated as Brmj = φmj
Sj, where Brmj is the flux dis-
tribution of the j-phase teeth by the permanent magnet,φmj is whole magnetic flux of the j-phase teeth, Sj isinterlinkage flux area on the j-phase teeth. Fig. 1(b)shows the approximate model of the flux distribution.
3.3 Approximate Model of Flux Distributionby d-axis Current Under the assumption of Sec-tion 3.1, Bri(ϕm) is calculated by the diffenrence be-
2 IEEJ Trans. IA, Vol.134, No.1, 2014
(Masato Kanematsu et al.)
S
Bri(ϕm)
SS
(a) Magnetic flux lines
-25
-20
-15
-10
-5
0
5
10
15
-30 -20 -10 0 10 20 30flux density Br(
ϕ m)[mT]
stator position ϕm[deg]
W U V
FEM Analysisapproximate model
(b) Bri(ϕm)
Fig. 2. Flux distribution Bri(ϕm) by d-axiscurrent
-1.5
-1
-0.5
0
0.5
1
1.5
-30 -20 -10 0 10 20 30
flux density Br(
ϕ m)[T]
stator position ϕm[deg]
W U V
FEM analysisapproximate model
Fig. 3. Magnetic field analysis and flux approxi-mation of Br(ϕm) (id = −20 [A])
tween Br as id = −1[A] and by Brm(ϕm). Fig. 2 showsthe flux distribution of Bri(ϕm) at θe = 0 Deg. Theflux distribution on the U-phase tooth is nearly flat. Onthe other hand, the flux distribution on the V-phase andW-phase teeth concentrate in side of the U-phase teeth.In order to obtain simple approximation model, the fluxdistributions on the V-phase and the W-phase teeth areregarded as flat.
φiU = ldid, φiV,iW = −12ldid (10)
where φiU,iV,iW are fluxes on U-phase, V-phase and W-
phase teeth by d-axis current. Here, ld:=√
23
Ld
PN .Fig. 3 shows the flux distribution by FEA and pro-
posed approximate model.
Table 1. IPMSM parameter
Stator Configuration Concentrated winding
interlinkage area coefficient γ 0.5
d-axis 1-teeth inductance ld 53.1 [µH]
turn number N 20
teeth area S 4.13 × 10−4 [m2]
1-teeth magnetic flux φ 3.65 [mWb]
4. Radial Force Approximation UsingFlux Distribution
4.1 Radial Force on U-phase Tooth Fig. 4(a)shows the flux distibution image on the U-phase tooth.The flux distribution Br(ϕm) on the U-phase tooth iscalculated as follows:
Br(ϕm) =(
φ
S+
ldidS
)(11)
By substituting (6) and (11) into (7), (12) is obtaind.
FrU =∫ ∫
fr(ϕm)dS
=B2
r (ϕm)2µ0
· S
=(φ + ldid)2
2µ0S(12)
4.2 Radial Forces on V and W-phase TeethThe flux distribution of the permanent magnet is not flaton the V-phase and W-phase teeth. Therefore, approxi-mate model is derived in two areas. Fig. 4(b) shows theflux distibution image on the V-phase and the W-phaseteeth. The flux distribution Br(ϕm), B′
r(ϕm) and theradial force FrV , FrW on the V-phase and the W-phaseteeth at θe = 0 Deg. are calculated as
Br(ϕm) =(
φ
2γS+
ldid2S
)(13)
B′r(ϕm) =
ldid2S
(14)
FrV , FrW =∫ ∫
fr(ϕm)dS
=Br
2(ϕm)2µ0
· γS +B′
r2(ϕm)2µ0
· (1 − γ)S
=(φ + ldid)2 + (1−γ)
γ φ2
8µ0S(15)
4.3 Current Reference Method to SuppressRadial Force In this section, d-axis current refer-ence is derived to suppress the 2nd order radial force.FrU (θe), FrV (θe), and FrW (θe) refer to the radial forcesof U-phase, V-phase and W-phase teeth when rotor elec-trical angle is θe. When θe is 0 or π rad., the radialforce on U-phase tooth is maximum and equal to FrU (0),which is obtained in (12). The radial force on U-phasetooth is minimum when θe is 1
2π or 32π rad. At this
point, it is difficult to approximate radial force becausethe flux which does not interlink exists. Therefore, theradial force FrU (π
3 ) is used instead of FrU (π2 ). When
3 IEEJ Trans. IA, Vol.134, No.1, 2014
(Masato Kanematsu et al.)
maxwell stressdistribution
linearsuperposition
φ ldid
S SBri(ϕm)Brm(ϕm)
Br(ϕm)
fr(ϕm)
S
(a) Flux distribution image on U-phase
Bri(ϕm)
1
2φ
B′
r(ϕm)
S
−
1
2ldid
Br(ϕm)
γS
maxwell stressdistribution
Brm(ϕm)
fr(ϕm)f ′
r(ϕm)
linearsuperposition
γS(1 − γ)S
(b) Flux distribution image on V and W-phase
Fig. 4. Radial force calculation
FrU(0)FrV (0)
(a) θe = 0
FrU(2
3π)
(b) θe = 23 π
Fig. 5. The positions of teeth and magnet
three-phase equilibrium is correct in IPMSM, FrU ( 23π)
equals FrV (0). In Fig. 5, the physical relationships be-tween teeth and magnet when θe changes are shown.FrV (0) is approximated in (15). Moreover, by cyclicnature, FrU (θe) has the equal values at electrical angleθe = 1
3π, 23π, 4
3π, 53π[rad]. Therefore, when the following
equation is true, it is predicted that 2nd radial force issuppressed largely.
FrU (0) = FrU
(23π
)(16)
As the equations (12) and (15) are functions for id, wecan select a d-axis current reference to satisfy the equa-tion (16). From (12), (15), and (16), the optimal d-axiscurrent which achieves minimum 2nd order radial forceis represented by
id:opt =(−1 ±
√1 − γ
3γ
)φ
ld. (17)
By taking into the consideration of 0 < γ ≤ 1, id:opt isreal number. Here, plus sign in (17) is selected in or-der to use minimum current amplitude. Using Table. 1,id:opt is calculated by
id:opt = −29.1A. (18)
4.4 The Influence of Flux Distribution by q-axis Current In this section, the influence of fluxdistribution attributed by q-axis current is considered.The FEA result of the flux distribution by q-axis currentis shown in Fig. 6. Fig. 6 shows that the flux distribu-tion on U-phase tooth is 0. The dq/3-phase transformmatrix Cdq
UV W at θe = 0 is given by
CdqUV W =
√23
1 0−1
2
√3
2
−12 −
√3
2
. (19)
This means the flux by q-axis current passes onlythrough the V-phase and the W-phase teeth. In Fig. 6,the magnetic flux and the q-axis flux affect each other,and this leads to asymmetry of the flux distribution onthe V-phase and W-phase teeth.
4.5 Verification of Radial Force EstimateEquation In this section, the accuracy of the op-
4 IEEJ Trans. IA, Vol.134, No.1, 2014
(Masato Kanematsu et al.)
S
Brq(ϕm)
SS
(a) Magnetic flux lines
-20
-10
0
10
20
30
-30 -20 -10 0 10 20 30flux density Br(
ϕ m)[mT]
stator position ϕm[deg]
W U V
(b) Briq (ϕm)
Fig. 6. Flux distribution Briq (ϕm) by q-axiscurrent
0
10
20
30
40
50
60
70
80
-60 -50 -40 -30 -20 -10 0
Radial Force [N]
d-axis current id [A]
|FrU - FrV|(appro.)amplitude of 2nd order(iq = 0[A])amplitude of 2nd order(iq = 10[A])
Fig. 7. Comparison with approximate radial forceand analysis data
timal d-axis current reference is validated. The dif-ference between FrU (0) and FrU ( 2
3π) is proportionalto the amplitude of 2nd radial force. In Fig. 7,∣∣FrU (0) − FrU
(23π
)∣∣ and 2nd radial force analyzed byFEA are shown on no-load and load conditions. The2nd radial force is minimized by the optimal d-axis cur-rent reference on each condition. On load condition, 2ndradial force at id:opt cannot be reduced to 0 by d-axiscurrent because the remaining radial force is caused byq-axis current. Fig. 7 shows the d-axis current referenceid:opt which is calculated on no-load condition is helpfulto reduce the 2nd radial force on load condition. Fig. 8shows the time waveforms when id:opt is 0 A and -29.1A on each condition. Fig. 8(a) and 8(b) show that (17)is an optimal reference to suppress the radial force vi-bration. 2nd radial forces are suppressed by over 95 %
-140
-120
-100
-80
-60
-40
-20
0
20
0 0.02 0.04 0.06 0.08 0.1 0.12
Radial Force Fr[N]
Time [s]
id = 0.0 [A]
id = -29.1 [A]
(a) iq = 0 A(no load condi-
tion)
-140
-120
-100
-80
-60
-40
-20
0
20
0 0.02 0.04 0.06 0.08 0.1 0.12
Radial Force Fr[N]
Time [s]
id = 0.0 [A]
id = -29.1 [A]
(b) iq = 10 A(load condi-
tion)
-70
-65
-60
-55
-50
-45
-40
0 20 40 60 80 100 120
Radial Force Fr[N]
time[ms]
id = -26.0 [A]
id = -29.1 [A]
id = -32.0 [A]
(c) iq = 0 A(detail)
0
10
20
30
40
50
60
no-load load
FrU [N](spectrum)
id = 0[A]id = -29.1[A]
(d) 2nd order spectrum of
radial force
Fig. 8. U-phase radial force FrU suppression withd-axis current
in no-load condition, by over 60 % in load condition inFig. 8(d).
5. Experiment
It is difficult to measure the radial force directly. Inthis paper, the radial accerelation outside the stator isevaluated instead of the radial force.
The experimental results at 1500 rpm(fe = 150 Hz)are shown in Fig. 9. Here, horizontal axis is the fre-quency normalized by electric angle frequency fe. Inthis paper, the negative d-axis current reference is lim-ited within -25 A because of the motor rating. Fig. 9(a)and 9(b) show the 2nd order acceleration on load con-dition is slightly larger than that on no-load condition.This result corresponds to FEA result as shown in Fig.8(d). 2nd order acceleration on load condition is less re-ductive by d-axis current than that on no-load conditionin Fig. 9(g) and 9(h). In no-load and load conditions,the radial forces of 2nd order are largely suppressed bythe d-axis current.
Fig 10 shows the change of amplitude of 2nd acceler-ation. In each condition, 2nd accelerations are largelydecreased. Amplitude of 2nd acceleration on load con-dition is larger than that on no-load condtion.
6. Conclusion
In this paper, an optimal d-axis current referencewhich reduces 2nd radial force is proposed. The relation-ship between d-axis current and 2nd radial force is ex-amined. The approximate model of the 2nd radial forcepredicts and controls the amplitude of 2nd radial accel-eration. Experiments are performed to demonstrate thevalidity of this modeling. In automotive applications,IPMSMs use wide range of rotation speed, and the noiseand vibration cause serious problem when the frequencyof 2nd radial force corresponds to the frequency of res-
5 IEEJ Trans. IA, Vol.134, No.1, 2014
(Masato Kanematsu et al.)
0
20
40
60
80
100
120
140
0 2 4 6 8 10
ar[mm/s2]
elec. Order[]
(a) id = 0 A(iq = 0A)
0
20
40
60
80
100
120
140
0 2 4 6 8 10
ar[mm/s2]
elec. Order[]
(b) id = 0 A(iq = 10A)
0
20
40
60
80
100
120
140
0 2 4 6 8 10
ar[mm/s2]
elec. Order[]
(c) id = −5 A(iq = 0A)
0
20
40
60
80
100
120
140
0 2 4 6 8 10
ar[mm/s2]
elec. Order[]
(d) id = −5 A(iq = 10A)
0
20
40
60
80
100
120
140
0 2 4 6 8 10
ar[mm/s2]
elec. Order[]
(e) id = −10 A(iq = 0A)
0
20
40
60
80
100
120
140
0 2 4 6 8 10
ar[mm/s2]
elec. Order[]
(f) id = −10 A(iq = 10A)
0
20
40
60
80
100
120
140
0 2 4 6 8 10
ar[mm/s2]
elec. Order[]
(g) id = −25 A(iq = 0A)
0
20
40
60
80
100
120
140
0 2 4 6 8 10
ar[mm/s2]
elec. Order[]
(h) id = −25 A(iq = 10A)
Fig. 9. Experimental Results ar spectrum
onant frequency of the stator. In this rotation speed,it is useful to reduce motor vibration with injecting theoptimal d-axis current although the motor efficiency de-creases temporarily.
The amplitude of high order radial force is smallerthan that of 2nd radial force. However, high order ra-dial force, specially 6th order, can easily transmit sta-tor and generate large acceleration when rotating speedequals stator natural frequency. In our future work, highorder radial force modeling and control method will berealized.
References
0
40
80
120
no-load load
ar [mm/s2](spectrum)
id = 0[A]id = -25[A]
Fig. 10. 2nd order spectrum of radial acceleration
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(5) M.Islam, R.Islam, T.Sebastian : ”Noise and vibrationcharacteristics of permanent magnet synchronous mo-tors using electromagnetic and structural analyses”,IEEE International Conference on Energy Conver-sion Congress and Exposition (ECCE), Vol.17, No.22,pp.3399-3405(2011)
(6) M.Boesing, Rik W.De Doncker : ”Exploring a vi-bration synthesis process for the acoustic character-ization of electric drives”, IEEE International Con-ference on Electrical Machines (ICEM), Vol.6, No.8,pp.1-6(2010)
(7) Tao Sun, Ji-Min Kim, Geun-Ho Lee, Jung-Pyo Hong,Myung-Ryul Choi: ”Effect of Pole and Slot Combi-nation on Noise and Vibration in Permanent MagnetSynchronous Motor”, IEEE Trans. Magnetics, Vol.47,No.5, pp.1038-1041(2011)
(8) Z.Q.Zhu, Z.P.Xia, L.J.Wu, G.W.Jewell : ”Analyti-cal modelling and finite element computation of ra-dial vibration force in fractional-slot permanent mag-net brushless machines”, International Conference onElectric Machines and Drives Conference(IEMDC),Vol.3, No.6, pp.144-151(2009)
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6 IEEJ Trans. IA, Vol.134, No.1, 2014
(Masato Kanematsu et al.)
Masato Kanematsu (Student Member) received the B.S.degree in Department of Electrical and Elec-tronic Engineering from The University ofTokyo, Tokyo, Japan, in 2012. He is currentlyworking towards the M.S. degree in the De-partment of Advanced Energy, The Universityof Tokyo, Chiba, Japan. His research interestsinclude motor drive and design for electric ve-hicle.
Mr. Kanematsu is a student member of the Institute of ElectricalEngineers of Japan, the Japan Society of Mechanical Engineers,and the Society of Automotive Engineers of Japan.
Takayuki Miyajima (Student Member) received the B.S.and M.S. degrees in electrical and computerengineering from Yokohama National Univer-sity, Yokohama, Japan, in 2009 and 2011,respectively. He is currently working to-wards the Ph.D. degree in the Department ofAdvanced Energy, The University of Tokyo,Chiba, Japan. His current work concerns mo-tor drive for electric and hybrid vehicles andelectric vehicle motion control.
Mr. Miyajima is a student member of the Institute of Electricaland Electronics Engineers and the Society of Automotive Engi-neers of Japan.
Hiroshi Fujimoto (Senior Member) received the Ph.D. de-gree in the Department of Electrical Engineer-ing from the University of Tokyo in 2001. In2001, he joined the Department of ElectricalEngineering, Nagaoka University of Technol-ogy, Niigata, Japan, as a research associate.From 2002 to 2003, he was a visiting scholarin the School of Mechanical Engineering, Pur-due University, U.S.A. In 2004, he joined theDepartment of Electrical and Computer Engi-
neering, Yokohama National University, Yokohama, Japan, as alecturer and he became an associate professor in 2005. He is cur-rently an associate professor of the University of Tokyo since 2010.He received the Best Paper Award from the IEEE Transactions onIndustrial Electronics in 2001, Isao Takahashi Power ElectronicsAward in 2010, and Best Author Prize of SICE in 2010. His in-terests are in control engineering, motion control, nano-scale servosystems, electric vehicle control, and motor drive. Dr. Fujimotois a member of IEEE, the Society of Instrument and Control En-gineers, the Robotics Society of Japan, and the Society of Auto-motive Engineers of Japan.
Yoichi Hori (Fellow) received his B.S., M.S., and Ph.D. de-grees in Electrical Engineering from the Uni-versity of Tokyo, Tokyo, Japan, in 1978, 1980,and 1983, respectively. In 1983, he joinedthe Department of Electrical Engineering, TheUniversity of Tokyo, as a Research Associate.He later became an Assistant Professor, an As-sociate Professor, and, in 2000, a Professor atthe same university. In 2002, he moved to theInstitute of Industrial Science as a Professor in
the Information and System Division, and in 2008, to the Depart-ment of Advanced Energy, Graduate School of Frontier Sciences,the University of Tokyo. From 1991-1992, he was a Visiting Re-searcher at the University of California at Berkeley.His research fields are control theory and its industrial applica-tions to motion control, mechatronics, robotics, electric vehicles,etc.Prof. Hori is the winner of the Best Transactions Paper Awardfrom the IEEE Transactions on Industrial Electronics in 1993 and2001, of the 2000 Best Transactions Paper Award from the Insti-tute of Electrical Engineers of Japan (IEEJ), and 2011 Achieve-ment Award of IEE-Japan.He is an IEEE Fellow and a past AdCom member of IES. He hasbeen the Treasurer of the IEEE Japan Council and Tokyo Sectionsince 2001. He is also a member of the Society of Instrument andControl Engineers; Robotics Society of Japan; Japan Society ofMechanical Engineers; and the Society of Automotive Engineersof Japan.He is the past- President of the Industry Applications Society ofthe IEEJ, the President of Capacitors Forum, and the Chairman ofMotor Technology Symposium of Japan Management Association(JMA) , the Director on Technological Development of SAE-Japan(JSAE) and the Director of Japan Automobile Research Institute(JARI).
Toshio Enomoto (Non-member) was born in 1965 inIbaraki, Japan. Since completing B.S., in Ap-plied Physics in 1988 at Tohoku Univ., from1988 to 1991, he worked for MAZDA MotorCo., Ltd. Since 1991, he has worked for NIS-SAN Motor Co., Ltd. Currently he is manag-ing groups of noise and vibration performance,engine mounting and exhaust system.
Masahiko Kondou (Non-member) was born in 1957 inTokyo, Japan. He received the B.S., degreefrom Chuo University, Tokyo, Japan, in 1981.Since 1981, he has worked for NISSAN MotorCo., Ltd.. Currently he has been developing agood noise and vibration performance for Elec-tric Vehicles.
Hiroshi Komiya (Non-member) received the B.E, degreesin mechanical engineering from Tokyo Univer-sity of Science, Japan, in 1989. Since 1989, Heworks as NVH development engineer in Nis-san Motor Company, and a particularly, expe-rience about Powertrain technology develop-ment are abundant. He is a member of Societyof Automotive Engineer of Japan.
7 IEEJ Trans. IA, Vol.134, No.1, 2014
(Masato Kanematsu et al.)
Kantaro Yoshimoto (Member) was born in 1975 in Yoko-hama, Japan. He received the B.S., M.S., andPh.D. from Yokohama National University,Yokohama, Japan, in 1997, 1999, and 2010 re-spectively. From 1999 to 2000, he worked forEast Japan Railway Co.,. Since 2001, he hasworked for NISSAN Motor Co., Ltd.. He isperforming the development of Electric Vehi-cles.
Takayuki Miyakawa (Non-member) was born in 1980 inKawasaki, Japan. He received the B.E., M.E.,degree from Chuo University, Tokyo, Japan,in 2003, 2005. Since 2005, he has worked forNISSAN Motor Co., Ltd.. Currently he hasbeen developing N.V. performance for ElectricVehicles and N.V. performance of chassis.
8 IEEJ Trans. IA, Vol.134, No.1, 2014
