1.5-kW, 150,000-r/min Ultra High-Speed PM Motor
Fed by 12-V Power Supply for Automotive Supercharger
Toshihiko Noguchi*, and Tetsuro Wada*** Shizuoka University
3-5-1 Johoku, Naka-Ku, Hamamatsu 432-8561, Japan
Tel.: +81-53-478-1128
Fax: +81-53-478-1128
E-Mail: ttnogut@ipc.shizuoka.ac.jp** Nagaoka University of Technology
1603-1 Kamitomioka, Nagaoka 940-2188, Japan
Tel.: +81-258-47-6000
Fax: +81-258-47-9500
E-Mail: twada@stn.nagaokaut.ac.jp

Acknowledgements
The authors wish to express their gratitude to Mr. Y. Sakai and Mr. K. Nakamura, Fukuda Corporation,
Tokyo for offering a valuable technical support on mechanical bearings, and to Mr. H. Akazawa andMr. H. Ninomiya, JFE Steel Corporation, Tokyo for providing electromagnetic steel plates.

Keywords«high speed drive», «permanent magnet motor», «adjustable speed drive», «automotive application»,«automotive component».

Abstract
This paper presents a 1.5-kW, 150,000-r/min PM motor fed by a 12-V power supply, which isapplicable to an automotive supercharger. The motor is specially designed to improve its efficiency
over 97 % (excluding a mechanical loss) and to raise its power density to 13 W/cm3 at the same time.
Feasibility of the design is confirmed through experimental tests, using a prototype motor.

Introduction
A supercharger is an auxiliary machine to enhance output power of a combustion engine and to reduce
the engine displacement at the same time. Employing the supercharger is also effective to improve the
combustion efficiency, the exhaust gas quality and the engine torque response. The conventional
supercharger has a mechanical linkage with the engine that uses a timing belt, and it compresses inlet
air into the engine cylinders by means of the mechanical power provided by the engine. Figure 1 (a)Displacement
compressor

InverterEngineBelt & Pulley

12 V PS
12-V
Battery

UltraUHS
high-speed
motor
motor

EngineCentrifugal
compressor(a) Conventional system.
(b) Electrically operated system.
Fig. 1: System configurations of conventional and electrically operated supercharging systems.

illustrates a mechanical configuration of the conventional supercharging system. As shown in the
figure, many of the superchargers employ a positive displacement compressor because its operation
speed is limited by a low revolution speed of the engine. However, the positive displacement
compressor has drawbacks such as lower efficiency and lower boost pressure than those of acentrifugal compressor, which prevents further performance improvement of the supercharging system.
In order to solve these problems, electrification of the supercharger is significantly important and isvery promising approach as a next-generation auxiliary machine system for future automobiles.
Figure 1 (b) shows an outline of the electrically operated supercharging system where an ultra highspeed permanent magnet (PM) motor is used to drive the centrifugal compressor instead of theconventional positive displacement compressor. Since the electrically operated supercharger allows
use of the centrifugal compressor and makes a mechanical-linkage-free system possible, many
advantages are obtained, e.g., more efficient operation, higher rotation speed, higher boost pressure,faster response of the inlet air compression, and smaller mechanical dimensions than those of theconventional displacement compressor based system. Furthermore, such a mechanical-linkage-freedesign gives freedom of mechanical placement around the engine, and makes it possible to reduceoverall mechanical losses as well as to eliminate the complicated linkage mechanism.
This paper discusses an optimum design of the ultra high-speed PM motor of which specificapplication is the electrically operated supercharger of the automotive engine. The motor is fed by athree-phase inverter with a 12-V battery as a DC power supply for an automotive application; thus the
motor design must meet a high-current and high-frequency operation requirement without sacrificing
the motor efficiency and power density. The investigated motor has a rated output power of 1.5 kW
and the maximum rotation speed of 150,000 r/min, respectively. In order to achieve this goal, various
technical issues must be solved, e.g., drastic reduction of the synchronous impedance, minimization of
the iron and the copper losses, further improvement of the motor efficiency and the power density,
mechanical stabilization in high-speed operation range, and so forth. In addition, these electricaldesign requirements must be satisfied all together with a compact and robust mechanical design. On
the way of optimization process in the machine design, a finite element method (FEM) based
electromagnetic field analysis is conducted to make fine-tuning of the detailed motor shape and to
maximize the efficiency and the power density at the same time. Consequently, the efficiency has
been improved over 97 % (excluding a mechanical loss) and the power density has been raised toapproximately 13 W/cm3 in the prototype motor design.

Required Specifications of Ultra High-Speed PM Motor
Assuming a 1,500-cc class automotive engine, the ultra high-speed PM motor is required to have therated power of 1.5 kW at the maximum rotation speed of 150,000 r/min to achieve the electricallyoperated supercharging system. When boosting the inlet air compression, extremely fast response isrequired because the compressor must be accelerated from several ten thousand r/min to the maximum
rotation speed in approximately 0.5 s, which surpasses a response time of the conventional
supercharger. In order to meet this requirement, the motor must have a short-duration overload
capacity, which is a double of the rated output power for 1 s. Table I shows target specifications of the
motor determined by the above requirement.

Basic Design Concept of Ultra High-Speed PM Motor
In order to achieve the highest efficiency and the highest power density among various sorts of electricmotors, a two-pole three-phase surface permanent-magnet synchronous motor (SPMSM) is focused on

Table I: Specifications of developed ultra high-speed PM motor
Assumed engine

1500 cc class

Rated output power

1.5 kW

Maximum rotation speed

150,000 r/min

Rated torque0.0955 Nm

Overload capacity and duration

3 kW (200 % load) for 1 s

as the best choice for the ultra high-speed motor drive because of the simple rotor structure and nomagnetizing current required, which implies higher efficiency than other motors.
The stator has a six-tooth six-slot structure and concentrated windings, which is remarkably effective
to reduce the copper loss and the leakage inductance thoroughly as well as the synchronous inductance.
Each phase has a pair of single or double-turned windings in parallel, and the windings are notordinary wires but copper bars of which shape is like an alphabetical letter “b.” It is necessary to
reduce as much stator iron loss as possible even at 150,000 r/min operation, so high-performance 6.5% silicon electromagnetic steel plates of which thickness is only 0.1 mm are employed to compose alaminated stator iron core.
On the other hand, the rotor consists of a strong Nd-Fe-B permanent magnet and a molybdenum alloy
shaft. Use of the Nd-Fe-B permanent magnet allows not only motor efficiency improvement but also
drastic reduction of the rotor size and inertia. In addition, such a strong Nd-Fe-B permanent magnet
that has BHmax = 310-kJ/m3 makes it possible to widen the air gap, which is essential to reduce thesynchronous inductance and to obtain a sinusoidal electromotive force (e.m.f.) regardless of theconcentrated stator winding structure.
Table II is a summary of the basic conceptual design parameters of the ultra high-speed PM motor to
be investigated in this paper.

Optimization of Permeance Coefficient and Stator Windings
Since a permeance coefficient determines an operating point on a B-H curve of the permanent magnet,
operation characteristics of the PM motor dominantly depend on the permeance coefficient. The
permeance coefficient is basically proportional to the e.m.f. unless other physical dimensions of theinvestigated motor are changed. Therefore, the permeance coefficient has a strong influence on the
motor efficiency because there is a trade-off relationship between the copper loss and the iron loss ofthe motor, depending on the e.m.f. Assuming that the investigated PM motor has a uniform
permeance distribution along the air gap, the permeance coefficient pu of the motor is expressed by thefollowing equation:ag
Dm + l gllm
pu = m
=
(1)
,
a K l
D −l
K l
m

C g

m

m

C g

where ℓm is a permanent-magnet thickness, am is an averaged cross section area of the permanent
magnet, ag is an averaged cross section area of the air gap between the rotor and the stator, ℓg is aradial air gap length, Dm is an outer diameter of the permanent magnet, and KC is a Carter’s coefficient.
Since KC normally takes a value of approximately 1.2 to 1.5, ag can be regarded as almost same as
amKC; thus, the following approximated expression is obtained:lpu ≈ m(2)
lg .
This equation shows that the permeance coefficient is determined by the ratio between the permanentmagnet thickness and the radial air gap length as illustrated in Fig. 2.

Table II: Conceptual design parameters of developed ultra high-speed PM motor
Motor typeSurface Permanent-Magnet Synchronous Motor (SPMSM)

Number of phases

3 phaseNumber of poles

2 poleStator winding configuration

Concentrated

Winding configuration

1- or 2-turn windings in parallel per phase
10JNEX900(0.1-mm thick, 6.5-% silicon, µs = 23,000, Bmax = 1.8 T)
Nd-Fe-B
N-39SH(Br = 1.28 T, bHc = 955 kA/m, BHmax = 310 kJ/m3)
Angular ceramic-ball bearings
with grease lubrication

Electromagnetic steel plates
Permanent magnet
Bearings

The synchronous impedance of the investigated PM motor must ultimately be reduced because themotor is operated under high-current and high-frequency operation conditions. Otherwise, it is
impossible to satisfy the target specifications, delivering the rated power of 1.5 kW at the maximum
rotation speed of 150,000 r/min with a 12-V DC power supply. This ultimately low synchronous
impedance can be accomplished by having a wide air gap owing to such a strong Nd-Fe-B permanent
magnet as BHmax = 310 kJ/m3, and by reducing the number of winding turns drastically. Figure 3 (a)indicates a stator winding configuration, where two models of the winding structure are illustrated in
the same figure, i.e., a one-turn and a two-turn structures. The copper-bar windings are cut out of bulk
copper plates so that their shape is like an alphabetical letter “b” as illustrated in Fig. 3 (b). The
leakage inductance can effectively be reduced by inserting these windings closely to teeth of the statoriron core.
In general, a back e.m.f. of the PM motor is expressed as follows:e = 2 E = 2 pωk w N φ g ,(3)where p is the number of pole pairs, ω is an operation speed in rad/s, kw is a winding coefficient, N is
the number of winding turns and φg is an air gap magnetic flux. As the above equation indicates, theback e.m.f. can be enlarged by increasing Ν or φg. Therefore, the two configurations of the stator
windings shown in Fig. 3 (a) are investigated from the viewpoint of the efficiency and power density
maximization and the voltage utilization of the inverter output. Suppose that each motor line current
is controlled to be in phase with the back e.m.f. of the corresponding phase, the total phase voltage
including a voltage drop of the inverter can be expressed by the following equation:
V = RFET I + Ra I + jωLa I + E ,(4)where RFETI is a voltage drop of the MOSFETs used in the inverter (approximately 2 mΩ/phase), RaI is
a voltage drop of the stator winding resistance, ωLaI is a voltage drop of the stator winding inductance,and E is the back e.m.f. The voltage drops of the motor with one-turn stator windings are listed in theleft column of Table III. As can be seen in the table, the voltage drop due to the synchronous
inductance is hardly observed as well as the leakage inductance drop. In addition, the back e.m.f. is
Permanent Magnet

Shaft

Coil (Winding)
Stator CorelglmFig. 2: Cross section diagram to investigate permeance coefficient.
Winding
Stator teeth

2 turn1 turn(a) Stator core and windings.
(b) Stator winding.
Fig. 3: Stator winding configuration.

unnecessarily too low to utilize the DC power supply voltage. Therefore, since both of the leakage
inductance and the synchronous inductance are sufficiently small, it is possible to operate the motorwith a 12-V DC power source even if the number of turn is double as indicated in the upper part of Fig.
3 (a), resulting in quadruple synchronous inductance and twice in the back e.m.f. In addition,changing the permeance coefficient pu from 1.67 to 0.882 to improve the motor efficiency, the voltage
drops of the motor with the two-turn stator windings are as listed in the right column of Table III. As
described previously, the winding resistance and the inductance become 4 times of those in the oneturn motor, but the increase of these voltage drops can be restricted within 2.5 times to 2.9 times
because the operating current is effectively reduced by approximately 30 % by optimizing the backe.m.f. Assuming that the maximum phase voltage applied to the motor is 4.9 Vrms, which can be fedby the inverter with the 12-V DC power supply, it is possible to utilize 90.0 % of the 12-V power
supply voltage at 1.5-kW operation and to utilize 97.5 % even in a 200-% over load condition without
a voltage saturation.

Optimization of Detailed Dimensions for Power Density Maximization
In order to maximize the power density of the motor without sacrificing the efficiency, the detailed
stator iron core shape, i.e., a yoke width, a outer diameter and a tooth width, is investigated throughFEM based electromagnetic analyses as shown in Fig. 4. The stator inner diameter is fixed at 28 mm,
which is derived from the calculation result of the permeance coefficient described above and thestack length of the stator iron core is limited to 30 mm to prevent harmful mechanical vibrations of theS

N

Teeth width

Outer diameter ( 70 mm )Shaft

Inner diameter( 28 mm)

Stator core
Winding( 2 turn )Permanent magnet

Back yoke widthFig. 4: Cross section diagram of 2-turn PM motor.

Table III: Voltage drops per phase at rated operation and other design parameters
Number of winding turns

1

2
2 mΩ

Resistance of inverter MOSFET RFET
Stator winding resistance Ra

0.072 mΩ0.200 mΩStator winding inductance La

0.070 µH0.294 µH

Voltage drop of inverter MOSFET RFETI

0.353 V

0.243 V

Voltage drop of the stator winding resistance RaI

0.0127 V

0.0243 V

Voltage drop of the stator winding inductance ωLaI

0.194 V

0.562 V

Back e.m.f. E

2.84 V

4.11 V

Total voltage drop V

3.21 V

4.41 V

Stator iron core stack length L

30 mm

Permanent-magnet thickness ℓm

5 mm

3.75 mm

Radial air gap length ℓg

3 mm

4.25 mm

Permeance coefficient pu

1.67

0.882

rotor shaft over the whole operation speed range. Figure 5 shows loss analysis results of the motorwith respect to the specified dimension of the stator iron core.
Figure 5 (a) shows electrical losses of the motor when only the stator yoke width is changed whileother parameters, i.e., the stator outer diameter and the tooth width, are kept constant. As the stator
yoke gets wider, the iron loss gradually decreases because the magnetic flux density is reduced in theiron core, but the copper loss becomes dominant among the losses. Since the copper loss is affectedby the slot cross section area for the windings, the wider stator yoke reduces the slot cross section area,
resulting in higher current density and higher resistance of the windings. On the other hand, when thestator yoke is too narrow, the motor is unable to deliver the torque due to the magnetic saturation in
the stator yoke. Therefore, it can be found that 6 mm is the best value for the stator yoke width.
Keeping the stator yoke width at 6 mm, a loss analysis result is shown in Fig. 5 (b) as the stator outer
diameter is changed. The reduction of the stator outer diameter makes the total iron core loss lessbecause the magnetic flux path gets shorter together with whole volume of the motor. However, thepercentage of the copper loss becomes more remarkable as the stator outer diameter gets smaller. The
reason of this copper loss enlargement is excessive reduction of the slot cross section area for thewindings. Therefore, the stator outer diameter of 70 mm takes the minimum loss, which is the mosteffective to optimize the overall motor dimensions from the viewpoint of the efficiency and the power
density.
70
60

Loss (W)50
40
30

Iron loss (yoke)

20

Iron loss (teeth)

Copper loss
(windings)

100

Eddy-current loss (PM)2

4
6Back yoke width (mm)810(a) Loss analysis result with respect to back yoke width.
1000
900

Loss (W)50
40
30

Iron loss (yoke)

20
Copper loss
10 (windings)
0
50

55

Iron loss (teeth)
Eddy-current loss (PM)

60
65
70
75
80
Stator core outer diameter (mm)8590(b) Loss analysis result with respect to stator outer diameter.
60
50
Loss (W)40
30 Iron loss (teeth)

Iron loss (yoke)

20 Eddy-current loss (PM)10
Copper loss (windings)

0

2

4

6
8
Teeth width (mm)1012(c) Loss analysis result with respect to teeth width.
Fig. 5: Loss analysis results with respect to detailed stator iron core shape.

In a similar way, the tooth width can be determined to be 10 mm to minimize the total loss. As thetooth width gets narrower, it is found that the eddy current loss on the rotor permanent magnet
becomes higher. This is because the wider slot opening due to the narrower tooth causes moredetrimental permeance variation along the air gap.
Consequently, the detailed stator iron core and rotor dimensions are summarized as listed in Table IV.
Since the total volume of the stator and the rotor including the air gap is only 115.5 cm3, the power
density at the rated output power reaches 13W/cm3, which is approximately 10 times of that ofcommon electric machines. The outer diameter of the rotor is as small as 19.5 mm; thus, thecircumference velocity of the rotor reaches 153.2 m/s, which is less than half of the sound speed. As
described later on, however, the rotor is mechanically reinforced by glass fiber threads with epoxy
resin to prevent the permanent magnet rotor from destruction due to large centrifugal force.
Figure 6 shows loss comparison at the rated operating condition between the one-turn and the two-turnmotors designed in this investigation. As can be seen in the figure, the two-turn motor achieves
drastic reduction of the electrical losses down to 36 W, compared with those of the one-turn motor, i.e.,
62 W. The copper loss of the two-turn motor is enlarged because of the small slot cross section area,
but each of the eddy current loss on the rotor permanent magnet and the stator iron core loss iseffectively reduced. It is inferred that the total efficiency of the designed two-turn motor excluding amechanical loss reaches 97.6 % although the motor is driven by such a low-voltage DC power supply
as a 12-V battery.

Prototype Ultra High-Speed PM Motor
As described in the basic design concept, the prototype has a special electrical and mechanicalstructure. Figure 7 shows photographs of the prototype. The laminated stator iron core consists ofapproximately 300 sheets of 6.5-% silicon electromagnetic steel plates, of which outer diameter is 70
mm, inner diameter is 28 mm, and axial stack length is 30 mm. The stator winding copper bars areinserted in the stator teeth with keeping electrical insulation from the stator iron core by polyimide
taping, and are connected to a neutral-point end ring all together. Each of the stator windings has across section area of 16 mm2, resulting in a current density of 7.6 A/mm2 at the rated operating
condition. Every clearance between each tooth and each stator winding is less than 0.3 mm, which
effectively improves the magnetic coupling with reducing the leakage flux.
On the other hand, the rotor is simply assembled with a ring-shaped Nd-Fe-B permanent magnet and amolybdenum alloy shaft, and is magnetized so that the flux distribution becomes sinusoidal. Afterassembling the rotor, 2.5-mm thick layer of the non-electrically-conductive glass fiber is formed withspecial epoxy resin on the permanent magnet surface against large centrifugal force. Mechanicalreinforcement with glass fiber can be seen in the photograph of the rotor exterior.

Table IV: Summary of designed dimensions
Stator outer diameter DO

70 mm

Stator stack length L

30 mm

Stator inner diameter DI

28 mm

Rotor PM outer diameter dO

19.5 mm

Stator yoke width YW

6 mm

Rotor PM thickness l m

3.75 mm

Stator tooth width TW

10 mm

Radial air gap length l g

4.25 mmFig. 6: Loss comparison between one-turn and two-turn motors at rated operating condition.

Figure 8 illustrates a three-dimensional computer graphic (a bird’s-eye view) of the prototype motor
assembly. All of the metal components are made with a high-precision NC machining tools. The
prototype is designed and created to have extremely high mechanical accuracy of µm-order, especiallyin the bearing brackets and the rotor shaft parts.

Experimental Setup and Test Results
Figure 9 shows a schematic diagram of an experimental setup to confirm basic operation
characteristics of the prototype motor. A pseudo current-source inverter was employed to drive themotor because of a high-fundamental-frequency over 2 kHz and extremely low-synchronous
inductance. The pseudo current-source inverter consists of a current-controlled buck-boost chopperand a six-step inverter. The former is operated with a DC-bus current feedback at a switchingfrequency of 48 kHz, resulting in significant reduction of the DC-bus reactor inductance. The latter
commutates the DC-bus current and generates 120-deg conduction patterns of the motor line currents.
Surge voltages during the current commutation are clamped by the DC-bus power source via
MOSFETs’ body diodes and a bypass diode in the chopper.
Table V represents comparison between the designed and the measured motor parameters. As listed
here, the measured motor inductance is slightly higher than the designed value because of the leakage
inductance and the line inductance to the motor.

(a) Stator iron core with two-turn windings.

(b) Nd-Fe-B PM rotor reinforced by glass fiber.

(c) Front and rear views of assembled motor.
Fig. 7: Photographs of prototype ultra high-speed PM motor.

Fig. 8: Three-dimensional computer graphic of motor assembly.

Figure 10 shows steady-state waveforms of the Hall-effect position sensor signal, the line current andthe terminal voltage of the motor operated at 150,000 r/min under no load condition. As can be seenin the figure, the 120-deg current pattern is properly generated in synchronism with the Hall-effectposition sensor signal. The motor terminal voltage is sinusoidal without a conspicuous harmonic
distortion, which implies the back e.m.f. is properly generated so as to be a sinusoidal waveform by
the rotor permanent magnet regardless of the concentrated winding structure of the stator.
Figure 11 shows an acceleration test result, which was conducted to examine the output torquecontrollability. Since measuring the mechanical output at such an ultra high-speed as 150,000 r/min israther difficult, the output torque of the prototype motor was estimated by a designed value of the rotor
inertia and an acceleration observed in the speed step response. As shown in the figure, it is inferredthat the maximum output torque delivered for the acceleration was approximately 0.08 Nm due to thecurrent limit of the inverter, which was 84 % of the rated value. Although the estimated output torquewaveform is choppy, it can be found from the waveform envelope that the speed is linearly regulated.

Conclusion
This paper discussed an optimum design to develop a 1.5-kW, 150,000-r/min ultra high-speed PM
motor fed by a 12-V DC power source, which is applicable to an automotive supercharger, from theviewpoint of efficiency and power density improvement. The prototype has various unique features in
its electrical and mechanical structure, which achieves a low-voltage, high-current and high-frequency
operation. Owing to the optimum design of the permeance coefficient and the detailed stator iron coreshape, the designed motor can achieve an ultimately high power density of 13 W/cm3 together with aremarkably high efficiency over 97 % (excluding a mechanical loss). The maximum-speed operation
under no-load condition and a speed step response with a 84-% output torque delivered wereexperimentally examined, and proper operation characteristics were confirmed through theexperimental tests.
Buck chopper

D
SC1

VDC
72 V

LC CT IL

Six-step
inverter

Su S v S w

SC2

C +

SPMSM

Sx Sy
48-kHz Carrier

ωm * +

-

PI

ωm

Ultra
high-speed
motor

+
-

Sz

Logic circuitPI

Hall-effectposition
sensor

Hu
Hv
Hw

F/V

Fig. 9: Schematic diagram of pseudo current-source inverter drive.

Table V: Measurement result of two-turn motor parameters
Motor parameters

Designed value

Measured value

E.m.f constant-5

2.74 x 10 V/r/min

2.67 x 10-5 V/r/min

Stator winding resistance Ra

0.200 mΩ0.151 mΩStator winding inductance La

0.294 µH

0.362 µH

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6
4
2
0
-2
15
10
5
0
-5
-10
-15
30
20
10
0
-10
-20
-30
0.0

0.4

0.8

1.2

1.6

2.0

Time (ms)

Estimated output torque (Nm)

Rotation speed (r/min)Fig. 10: Operating waveforms at 150,000 r/min under no load condition.
45000
40000
35000
30000
25000
20000
0.10
0.08
0.06
0.04
0.02
0.00
0.0

0.2

0.4

0.6

0.8

1.0

Time (s)Fig. 11: Speed step response and experimentally estimated output torque.