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Cylindrical Rotor Design for Acoustic Noise
and Windage Loss Reduction in Switched
Reluctance Motor for HEV Applications
Kyohei Kiyota, Student Member, IEEE, Takeo Kakishima, Student Member, IEEE,
Akira Chiba, Fellow, IEEE, and M. Azizur Rahman, Life Fellow, IEEE
Abstract—A switched reluctance motor (SRM) is one of the
candidates of rare-earth-free motors for hybrid electric vehicles
(HEVs). An SRM has been developed with same interior permanent magnet synchronous motor (IPMSM) dimensions having
competitive maximum torque, operating area, and maximum efficiency. However, this SRM has a windage loss of 1.3 kW due to the
salient poles of the SRM driven at the maximum rotational speed.
In addition, considerable acoustic noise is caused by the salient
poles. In this paper, a simple design of a cylindrical outer shape
rotor is proposed, and a comparison with the conventional SRM
rotor is carried out. It is found that the efficiency is improved due
to the windage loss reduction. It is also found that the acoustic
noise is significantly reduced in the proposed rotor design.
Index Terms—Acoustic noise, efficiency, hybrid electric vehicle
(HEV), switched reluctance motor (SRM), windage loss.
I. I NTRODUCTION
T
HERE are considerable demands for developing electronic motors with high torque density, high efficiency,
and low cost for hybrid electric vehicles (HEVs), plug-in HEVs,
and electric vehicles (EVs) [1]. An interior permanent magnet
(PM) synchronous motor (IPMSM) with rare-earth NdFeB PMs
is one of the most popular electric motors [2]. The cost and
the supply are problems of rare-earth PMs. In recent years,
there are strong demands of a rare-earth-free motor or lessrare-earth motor for HEVs and EVs. There are several types of
rare-earth-free motors or less-rare-earth motors. A synchronous
motor with ferrite-type PMs is one of the investigated motors
[3], [4]. Those motors have competitive efficiency and power
Manuscript received February 20, 2015; revised May 18, 2015; accepted
July 6, 2015. Date of publication August 11, 2015; date of current version
January 18, 2016. Paper 2015-EMC-0123.R1, presented at the 2014 IEEE
Energy Conversion Congress and Exposition, Pittsburgh, PA, USA, September
20–24, and approved for publication in the IEEE T RANSACTIONS ON I NDUS TRY A PPLICATIONS by the Electric Machines Committee of the IEEE Industry
Applications Society.
K. Kiyota, T. Kakishima, and A. Chiba are with the Department of Electrical
and Electronic Engineering, Tokyo Institute of Technology, Tokyo 152-8552,
Japan (e-mail: kiyota.k@belm.ee.titech.ac.jp).
M. A. Rahman is with the Department of Electrical and Computer Engineering, Memorial University of Newfoundland, St. John’s, NL A1B 3X5, Canada
(e-mail: arahman@mun.ca).
Color versions of one or more of the figures in this paper are available online
at http://ieeexplore.ieee.org.
Digital Object Identifier 10.1109/TIA.2015.2466558
density, although torque density is inferior to that of rare-earth
PM motors. Mechanical strength is also a critical point of those
motors. Field-winding-type motors and induction motors have
been also one of the prospective rare-earth-free motors [5], [6],
but improvement of efficiency and torque density are necessary.
A switched reluctance motor (SRM) is one of the rare-earthfree motors. SRMs do not need PMs; only low-loss silicon steel
and stator concentrated windings are needed. Thus, SRMs have
several advantages, such as low cost and possible operation
in high-temperature environment for an internal-combustion
engine. Rotor robustness is also one of the advantages because
of the simple rotor structure.
However, SRMs have four problems: 1) low power and
torque densities, 2) low efficiency, 3) high noise and vibration,
and 4) need of special inverter. Improvements of efficiency have
been the first challenging problem of this project. There are
several papers, which study power density or efficiency improvements [7]–[12]. The torque and power densities of SRMs
are in the range of 7–30 Nm/L and 0.5–6 kW/L, respectively. In
the previous papers, the torque and power densities have been
improved as 35 Nm/L and 18.4 kW/L [13]–[15], respectively.
In addition, the efficiency has been enhanced to be competitive
to those of IPMSMs [13]–[16]. Hence, the decrease of acoustic
noise with competitive torque and power densities, as well as
efficiency, is the second challenging problem. An SRM has
salient poles in the rotor. It results in high windage loss and
acoustic noise caused by salient poles at the high-speed operation. Thus, acoustic noise is one of the problems to be solved.
In this paper, a cylindrical-shape rotor is designed for the reduction of the windage loss and acoustic noise. First, the windage loss and the acoustic noise of the salient poles of the 60-kW
SRM are measured. It is shown that the windage loss caused by
the salient poles is as high as 1.3 kW at the maximum rotational speed of 13 900 r/min. The windage loss resulted in the
decrease of efficiency in high speed and for low power region.
Thus, the reduction of the windage loss has been investigated.
There are a few papers to reduce the windage loss by the rotor
rib [17], [18]. Those papers evaluate the total losses at the noload condition, while the measurement of the efficiency of the
load test and the acoustic noise are not evaluated. Not only the
low mechanical loss but also the efficiency improvements at
the load condition, as well as the low acoustic noise, are important improvements in a cylindrical outer shape rotor. Moreover,
the detailed analysis of the mechanical strength is not described
0093-9994 © 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission.
See http://www.ieee.org/publications_standards/publications/rights/index.html for more information.
KIYOTA et al.: CYLINDRICAL ROTOR DESIGN FOR ACOUSTIC NOISE AND WINDAGE LOSS REDUCTION IN SRM
155
Fig. 1. Power flow from the input to the output.
Fig. 3. Fabricated cylindrical dummy rotor.
rotational speed is 13 900 r/min. Thus, the maximum frequency
of current and the rotor peripheral speed are 2.78 kHz and
132 m/s, respectively. The iron stack length of the stator is
87.25 mm. Deep-groove ball bearings are installed in the SRM.
The outer diameter of the shrouds and the bottom diameter of
the rotor slot are 130 mm and 144 mm, respectively. There are
no shrouds at both ends of the rotor to avoid the additional iron
loss at the shrouds. Therefore, axial air gas flow may cause
additional windage loss and acoustic noise at both ends of the
rotor. The expression of mechanical loss Wmp is given as
Fig. 2. Conventional SRM rotor and stator.
Wmp = Wwp + Wwc + Wb
in those papers. In this paper, the rotor outer diameter is around
6 times large with respect to those in [17] and [18], whereas
the rib width is as small as 0.2 mm to enhance efficiency. This
thin rib is possible because the iron plate thickness is as thin as
0.1 mm. However, the thin rib causes deformation at high rotational speed; thus, an offset structure has been proposed, and the
mechanical strength analysis has been carried out. It is found
that the rib has enough mechanical strength at the maximum
rotational speed. Moreover, the analysis of the load efficiency
improvement is carried out. It is found that the efficiency is
improved by 3.9% at 30-kW output and at the maximum rotational speed. It is also found that the acoustic noise is reduced
by 14.2 dB. Finally, the prototype machine is fabricated and
experimentally tested.
II. E STIMATION OF THE W INDAGE L OSS AND ACOUSTIC
N OISE C AUSED BY THE S ALIENT P OLE OF THE SRM
In an SRM, there are copper loss, iron loss, bearing loss,
windage loss, etc. Fig. 1 shows the power flow from electrical
input to mechanical output. There is additional windage loss
caused by salient poles in SRMs, which does not exist in
IPMSMs and induction machines. The mathematical calculation of the windage loss caused by the salient poles of the
SRM is previously presented in [19] by some of the authors.
To avoid overlap, only test results of the conventional rotor and
the cylindrical-shape rotor are presented here.
Fig. 2 shows the pictures of a conventional SRM reported
in [15]. The outer diameter of the stator iron core and the
rotor is 264 mm and 182 mm, respectively. The maximum
(1)
where Wwp is a windage loss of salient poles, Wwc is a windage
loss of a cylindrical rotor, and Wb is a bearing loss. Note that
the windage loss of an SRM is supposed to be a sum of the one
caused by cylindrical rotor and the additional loss caused by
rotor salient poles.
Fig. 3 shows the cylindrical-shape dummy rotor with the
identical outer diameter and bearings of the test SRM rotor.
The mechanical loss Wmc of the cylindrical dummy rotor is
given as
Wmc = Wwc + Wb .
(2)
From (1) and (2), Wwp is given as
Wwp = Wmp − Wmc .
(3)
The mechanical losses Wmp and Wmc are the product of the
measured torque and the rotational angular speed when the load
electrical machine drives the shaft with no excitation in SRM
stator, with an SRM rotor and the cylindrical dummy rotor,
respectively. Note that the weights of the dummy rotor shaft and
the SRM rotor shaft are 20.9 kg and 16.6 kg, respectively. Thus,
the measured Wmc includes additional bearing loss caused
by the increased weight. It is also noted that the bearings
are changed with the rotor shaft for the measurement of the
cylindrical dummy rotor. Hence, the measured mechanical loss
includes the error of bearing condition. In the measurement, the
bearing loss is corrected by the difference of the estimated noload torque at zero speed.
Fig. 4 shows the mechanical loss of the conventional rotor
and the cylindrical dummy rotor. The mechanical loss is an
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IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 52, NO. 1, JANUARY/FEBRUARY 2016
Fig. 4. Mechanical loss of the cylindrical rotor and the salient-pole rotor.
Fig. 6. No-load acoustic noise of the cylindrical rotor and the salient-pole
rotor at 13 900 r/min.
Fig. 5. No-load acoustic noise of the cylindrical rotor and the salient-pole
rotor (overall).
average of 60-s measurement with the sampling time of 0.1 s.
In the correction, the increased torque (0.19 Nm) of the bearings
is removed from the test result of the cylindrical dummy rotor.
The corrected mechanical loss of the cylindrical rotor is 0.7 kW
at 13 900 r/min. The difference of the mechanical loss at the
maximum rotational speed is 1.3 kW, as shown in Fig. 4. Hence,
the experimental windage loss caused by the salient poles of
the SRM is 1.3 kW at 13 900 r/min. Note that the calculated
windage loss caused by the salient poles was estimated as
1.3 kW at the maximum rotational speed in [19]. Thus, the
windage loss can be well calculated by the simple equation.
Fig. 5 shows the overall no-load A-weighted acoustic noise
of the test system, the conventional rotor, and the cylindrical
dummy rotor. The acoustic noise difference between the test
system and the cylindrical dummy rotor is almost constant,
whereas that between the test system and the conventional rotor
is increased. The difference of the acoustic noise level between
the conventional rotor and the dummy cylindrical rotor is
14.2 dB at the rotational speed of 13 900 r/min.
Fig. 6 shows the frequency component of the no-load
acoustic noise of the two rotors at 13 900 r/min. The 36th harmonic acoustic noise of 8340 Hz, i.e., the least common multiple of the rotor poles 12 and the stator poles 18 is the highest
in the conventional rotor. This means that the acoustic windage
noise is mainly generated when the rotor poles pass the stator
poles. The 24th harmonic acoustic noise at 5560 Hz that is multiple of the number of the rotor poles is also apparent. Thus, the
salient poles of the conventional rotor generate the significant
acoustic noise.
Fig. 7. Conventional rotor (left) and the proposed cylindrical rotor (right).
III. D ESIGN AND A NALYSIS R ESULT OF THE
C YLINDRICAL O UTER S HAPE ROTOR
A. Design of the Cylindrical-Shape Rotor
Fig. 7 shows the conventional rotor and the proposed cylindrical outer shape rotor. The proposed cylindrical rotor has
12 poles as the conventional rotor. The outer diameter, the radius of rotor slot bottom, and the pole structure of the proposed
cylindrical rotor are the same as those of the conventional rotor.
Note that there are thin ribs between the salient poles in the
proposed cylindrical rotor for the reduction of the windage
loss and the acoustic noise level. The thickness of the rotor
ribs are one of the most important design improvements of the
cylindrical-shape rotor. When the magnetic material ribs are
installed between the salient poles, the unaligned inductance
is increased; then, the output torque and the efficiency are decreased. Thus, the ribs must be as thin as possible. On the other
hand, the rotor peripheral speed at 13 900 r/min is 132 m/s;
thus, the ribs must have enough mechanical strength to tolerate
the centrifugal force. In addition, the ribs must withstand the
excitation force generated by the stator coils. Note that these
rotors employ 6.5% high silicon steel characterized by low
iron loss to improve the efficiency. The yield point of the iron
material is 550 MPa; thus, the target maximum Mises stress is
set to 275 MPa because the safety factor is set to as 2.0, as for
the first step.
Fig. 8 shows the enlarged view of the interval of the salient
poles of the proposed cylindrical rotor. The width of the ribs
is set to 0.2 mm that is twice the width of one sheet of the
KIYOTA et al.: CYLINDRICAL ROTOR DESIGN FOR ACOUSTIC NOISE AND WINDAGE LOSS REDUCTION IN SRM
157
Fig. 8. Enlarged view of the interval of the proposed cylindrical rotor.
Fig. 11. Von Mises stress of points A and B of the rotor at 13 900 r/min.
Fig. 9. Enlarged view of the distribution of von Mises stress and the displacement of the rotor rib at 13 900 r/min.
base of the rotor rib. This value is about half of the yield point
of the iron material. The maximum deformation of the rib, on
the other hand, is 107 μm at the center of the rotor rib, which
is 21% of the gap length between the rotor poles and the stator
poles. Thus, the 0.1-mm offset is proposed in the rotor rib. Thus,
the air gap length between the rotor ribs and stator poles is only
7 μm decreased at the maximum rotational speed.
Fig. 11 shows the variation of von Mises stress of the
points A and B, when the A-phase coil is excited at the current
of 300 A, i.e., 1.25 times high current with respect to the
maximum peak current. The maximum Mises stress values of
points A and B are 60.6 MPa and 32.3 MPa, at the mechanical
rotor positions of 24◦ and 15◦ , respectively. These values are
about one-fifth and one-tenth of the yield point of the iron
material, respectively; thus, the rotor rib has enough robustness
to the excitation force generated by the stator coils.
B. Comparison of the Efficiency Between the Conventional
Rotor and the Proposed Cylindrical-Shape Rotor
Fig. 10. Von Mises stress and the displacement of the rotor at 13 900 r/min.
steel material thickness of 0.1 mm. Note that there is a 0.1-mm
offset from the rotor outer radius. This offset value has been
carefully designed based on the further analysis described as
follows. Fig. 9 shows the displacement and von Mises stress
of the rotor outer radius, and the deformation and the contour
of Mises stress distribution of the rotor rib at the rotational
speed of 13 900 r/min, respectively. The deformation in Fig. 9 is
expanded to 100 times the amount of actual displacement.
Fig. 10 also shows these results in the 2-D graph. The horizontal
axis in Fig. 10 is the mechanical angular position corresponding
to the rotational coordinate in Fig. 9. The centrifugal force of
100 MPa is stressed at the rib. The maximum Mises stress is
213 MPa at the mechanical angular position of 6◦ that is the
In order to confirm the efficiency improvement due to the
provision of the ribs between the rotor poles, a 2-D finiteelement method analysis is carried out. The stator and winding
parameter of the proposed rotor is identical to that of the
conventional rotor presented in [13]; thus, the difference of the
motor performance only depends on the difference of the rotor
structure. Table I shows a comparison of the proposed rotor and
the conventional rotor at three drive points: 30-Nm output at
2768 r/min and 30-kW output at 7500 r/min and 13 900 r/min.
The upper part compares the excitation conditions. Note that
the current hysteresis control is applied at the 4200-r/min and
lower speed region, while the one-pulse control is applied at
the 5400-r/min and higher speed region; thus, the referenced
peak current is shown at only 2768 r/min. The dc voltage of the
inverter is 650 V, at each drive point, in both the proposed and
conventional rotors. In the low speed region, the turn-on and
turn-off angles of the proposed rotor are the same as those of
the conventional rotor, whereas the referenced peak current is
increased by 3 A to compensate torque reduction. In the high
rotation speed region such as 7500 r/min and 13 900 r/min,
in contrast, the turn-on angle of the proposed rotor is advanced
from that of the conventional rotor. Note that the turn-off
angle of the proposed rotor is also advanced from that of the
conventional rotor to avoid the current continuous control.
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TABLE I
C OMPARISON OF THE C ONVENTIONAL ROTOR AND THE P ROPOSED ROTOR AT T HREE D RIVE P OINTS
In the lower part in Table I, the losses and efficiencies are
compared. Note that the efficiency ηe is calculated as
ηe =
Po
Po + WCu + WFe
(4)
where Po is the calculated output power, WCu is the copper
loss, and WFe is the sum of the iron losses. To see the efficiency
influence of the mechanical loss Wm , the total efficiency ηm is
calculated as follows:
ηm =
Po − Wm
.
Po + WCu + WFe
(5)
The mechanical losses of the conventional rotor and the proposed rotor are given by the test results of the conventional rotor
and cylindrical dummy rotor in Fig. 4. The RMS current of the
proposed rotor is almost the same as that of the conventional
rotor, although the turn-on angle is advanced. In the proposed
rotor, the unaligned inductance is higher than that of the conventional rotor in the low current region because of the rotor rib
iron; thus, the current rise is slow. The iron loss of the proposed
rotor is rather high in the rotor increase of the current peak.
Moreover, there is an additional iron loss of the rotor rib in the
proposed rotor; therefore, the total iron losses of the proposed
rotor are 10 W, 104 W, and 10 W higher than that of the
conventional rotor at 2768 r/min, 7500 r/min, and 13 900 r/min,
respectively. Thus, the efficiency ηe of the proposed rotor is
0.1%–0.3% low, with respect to that of the conventional rotor,
because of the increase of the unaligned inductance. Referring
to Fig. 4, on the other hand, the estimated total mechanical loss
of the proposed rotor is 24 W, 308 W, and 1258 W low, with respect to that of the conventional rotor. Consequently, it is found
that the total efficiency ηm of the proposed rotor is 3.9% high,
with respect to that of the conventional rotor, at 13 900 r/min,
i.e., the maximum rotational speed. At 2768 r/min, the total
efficiency ηm of the proposed rotor is almost the same as that
of the conventional rotor. Note that Table I also shows the
peak-to-peak torque ripple of the conventional rotor and the
Fig. 12. Efficiency variation of the conventional rotor and the proposed rotor
at 2768 r/min and 7500 r/min.
proposed rotor. The torque ripple of the two rotors are depending on the rotational speed and output torque, but the values
are generally as high as 140% to 210%, as expected. Thus,
load systems generally have gears and torsional shaft to absorb
torque ripple.
Fig. 12 shows the total efficiency ηm of the proposed rotor
and the conventional rotor at 2768 r/min and 7500 r/min. The
maximum RMS current is set to 141 A, in both SRMs. The
efficiencies of the proposed rotor and the conventional rotor
are drawn by the solid and broken curves, respectively. The
total efficiency ηm of the proposed rotor is superior to that of
the conventional rotor, at all output ranges, at 7500 r/min. The
maximum total efficiency of the proposed rotor is 0.6% high,
with respect to that of the conventional rotor, at 7500 r/min.
The efficiency improvement is high in the lower power region
because the reduction of the mechanical loss is significant in
this region. The output power of the maximum efficiency point
of the proposed rotor is lower than that of the conventional rotor, at both 2768 r/min and 7500 r/min, because of the reduction
of the mechanical loss. At 2768 r/min, the total efficiency ηm of
the proposed rotor is superior to that of the conventional rotor
at 25 kW and low output power region. The reduction of the
windage loss is still effective to improve the total efficiency
at 2768 r/min.
KIYOTA et al.: CYLINDRICAL ROTOR DESIGN FOR ACOUSTIC NOISE AND WINDAGE LOSS REDUCTION IN SRM
159
The fabricated rotor is installed in another SRM, which has
an identical stator and motor case design to the conventional
rotor. The winding resistance was 99.9 mΩ that was 2.2% high,
with respect to that of the conventional rotor. In these test
machines, forced cooling is not implemented, i.e., only natural
cooling is used.
B. Test Condition
Fig. 14 shows a test system configuration of the proposed rotor. The load machine, i.e., a motor/generator is driven by a voltage source inverter. The maximum speed of the load machine
is limited to 7500 r/min. The tested SRMs are driven by the
SRM inverter, of which the maximum peak is limited as 150 A;
thus, the maximum torque test is not carried out. The dc voltage
of the SRM inverter is fixed to 650 V by the dc–dc converter. A
torque and speed transducer (Onosokki DD-206S) is connected.
A three-channel digital power meter (Hioki 3192) is inserted to
measure the electrical input power. The input power measurement error of the digital power meter is 0.2% of the product of
the range of the RMS current and RMS voltage with additional
0.2% of reading value. Note that another three-channel digital
power meter (Hioki 3193) is inserted during the test of the conventional rotor. The full scale error of Hioki 3193 is identical to
Hioki 3192, whereas the reading error is 0.1% lower than that
of Hioki 3192; thus, the efficiency accuracy of the proposed
rotor is 0.1% low, with respect to that of the conventional rotor.
A precision sound level meter is located on the rotational axis
with a distance of 10 cm from the motor rear cover.
Fig. 13. Fabricated proposed rotor. (a) Rotor core and shaft. (b) Enlarged view
of the rotor core.
C. No-Load Test Results
IV. T EST R ESULT OF THE P ROPOSED
C YLINDRICAL -S HAPE ROTOR
A. Fabricated Cylindrical-Shape Rotor
Fig. 13 shows the pictures of a fabricated cylindrical-shape
rotor. The thin steel sheets are adhered and laminated around
30 mm. Then, the iron core is cut by wire. The wire-cut process
is not good in a point of view of efficiency. A punch-out process
is recommended; however, due to a limited budget, the wirecut process is used to see windage loss reduction. The outer
diameter of the test rotor is 181 mm. There are shrouds that
have the same material and the same outer diameter at both
ends of the rotor core to reduce the axial gas flow at the
interpolar region of the rotor core. Note that shrouds also can
reduce the windage loss caused by the salient poles of the
rotor [20]. When only shrouds are installed to the conventional
rotor, the additional windage loss is 0.5 times of the windage
loss of the cylindrical-shape rotor. The iron stack length of
the rotor core including the shrouds is extended to 91.03 mm,
to avoid the additional iron loss at the shroud. In Fig. 13(b),
the shrouds are not attended; thus, rotor ribs can be seen. The
fabricated rotor is tested by the same stator and motor case
design machine; thus, the other dimensions are the same as
that of the conventional rotor. The mechanical strength of the
rotor ribs is confirmed at the maximum rotational speed of
13 900 r/min.
Fig. 15 shows the measured no-load torque Tm of the fabricated rotor and the conventional rotor. The no-load torque
of the fabricated rotor is almost constant, whereas that of the
conventional rotor is increased at the high rotational speed
region. The increase of the no-load torque of the conventional
rotor is caused by the windage loss of the rotor poles; thus, the
windage loss is suppressed in the fabricated rotor. At the low
rotational speed of 2768 r/min, on the other hand, the torque
of the fabricated rotor is higher than that of the conventional
rotor. Note that the torque transducers have ±0.2% error with
respect to the full-scale torque of 200 Nm; thus, the torque measurement error caused by the torque detector is ±0.4 Nm. It is
also noted that grease is also used with automatic transmission
fluid for the lubricant; thus, the bearing loss of the fabricated
rotor may be increased by the difference of the lubricant.
Experiments were carried out due to the limited rotational speed
of 7500 r/min of the load system. At this stage, the efficiency
improvements were not confirmed experimentally because of
the iron process in the test machine fabrication. Iron surface
etching or the punch-out process will be needed in the future
project instead of the wire-cut process.
Fig. 16 shows the overall no-load A-weighted acoustic noise
of the test system and the proposed rotor. The acoustic noise difference between the test system and the proposed rotor is almost
constant as that of the cylindrical dummy rotor. The acoustic
noise difference between the test system and the proposed rotor
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Fig. 14. System configuration of the proposed rotor.
TABLE II
T EST R ESULT C OMPARISON OF THE C ONVENTIONAL ROTOR
AND THE P ROPOSED R OTOR AT T WO D RIVE P OINTS
Fig. 15. Mechanical loss of the fabricated proposed rotor and the conventional
rotor.
2768 r/min and 7500 r/min. The upper part compares the excitation conditions. The control strategy of the proposed rotor at
the machine test is the same as that at the analysis. The increase
of the referenced peak current is 5 A, i.e., the decrease of the
output torque per ampere of the test machine is higher than that
of the analysis. In 7500 r/min, the different of the turn-on angle
between the proposed rotor and the conventional rotor is also
expanded to 1◦ .
The lower part in Table II compares the losses and efficiencies of these SRMs. The total efficiency ηm is calculated as
Fig. 16. No-load acoustic noise of the fabricated proposed rotor and the
conventional rotor (overall).
is 11.9 dB at 7500 r/min. The maximum rotational speed of
the test system is limited to 7500 r/min. Fig. 16 also shows the
overall no-load A-weighted acoustic noise of the conventional
rotor previously shown in Fig. 6. The overall acoustic noise of
the fabricated rotor is 11.4 dB low, with respect to that of the
conventional rotor, at the rotational speed of 7050 r/min. Note
that the acoustic noise of the conventional rotor is measured
by a different load system in a different place. The test motor is
covered by the small acoustic chamber. Thus, the acoustic noise
is small in the low speed region.
D. Motor Efficiency of the Fabricated Rotor
Table II shows a comparison of the proposed rotor and
the conventional rotor at two drive points: 30-Nm output at
ηm =
Pom
Tω
=
Pi
Pi
(6)
where Pom and Pi are the measured output and input power,
as indicated in Fig. 1; T is the measured torque; and ω is
the shaft angular velocity. The iron and other losses WFe are
calculated as
WFe = Pi − T ω − WCu − Wm .
(7)
Thus, the iron and other losses include not only the iron loss
of the stator and rotor core but also additional wiring resistance
caused by the skin effect or other additional losses. The mechanical loss is measured after each efficiency measurement
because about 20% variation in the mechanical loss results
depending on the bearing condition. The copper loss of the
proposed rotor is 10 W and 60 W high, with respect to that
of the conventional rotor, at 2768 and 7500 r/min, respectively.
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161
ACKNOWLEDGMENT
The authors would like to thank M. Saito at Motion System Tech, for the test machine fabrication, and Myway Plus
Corporation, for fabricating a field-programmable gate array
controller and an inverter for switched reluctance drives.
R EFERENCES
Fig. 17. Acoustic noise of the conventional rotor and the proposed rotor at
7500 r/min, 30 Nm point.
The increase of the copper loss of the test machine is slightly
high, with respect to that of the analysis, because of the increased wiring resistance and the difference of the RMS current. Note that measurements of the proposed rotor are carried
out at a coil-end temperature of 75 ◦ C, which is the identical
temperature at the measurement of the conventional rotor. The
iron and other losses of the proposed rotor are 234 W and 497 W
higher than that of the conventional rotor at 2768 r/min and
7500 r/min, respectively; those values are significantly increased. Note that the conventional rotor is processed by the
punching; thus, the eddy current loss at the rotor surface is
considerably low. On the contrary, the fabricated proposed rotor
is processed by the wire cut after being adhered and laminated;
thus, the eddy current at the rotor surface is likely to be
occurred; therefore, the significant increase of the iron and other
loss is caused by not only the rotor ribs but also the eddy current
of the rotor surface. The mechanical loss of the proposed rotor
is rather high, with respect to that of the conventional rotor, because of the increase of the bearing loss shown in Fig. 15. Consequently, the total efficiency ηm of the proposed rotor is 3.9%
and 1.7% low, with respect to that of the conventional rotor
at 2768 r/min and 7500 r/min, respectively. The efficiency decrease at 7500 r/min is smaller than that at 2768 r/min because
of the windage loss decrease of the fabricated proposed rotor.
Fig. 17 shows the frequency component of the acoustic noise
of two SRMs at the identical excitation condition point of
7500 r/min and 30-Nm output shown in Table II. The 36th
harmonic acoustic noise of the proposed rotor is 8.4 dB low,
with respect to that of the conventional rotor.
V. C ONCLUSION
In this paper, the cylindrical-shape rotor in SRMs has been
proposed to reduce the windage loss and acoustic noise at high
rotational speed. The salient poles were connected by thin ribs,
so that the outer rotor surface is mostly cylindrical. The rib is
designed, considering the expansion at the high rotational speed
of 13 900 r/min. Mechanical and electromagnetic restriction
is considered and confirmed in the experiments. The acoustic
noise at 7050 r/min is reduced by 11.4 dB. Noise reduction
may be 14.2 dB at 13 900 r/min. Efficiency improvement is
estimated theoretically; however, the confirmation in the experiments was left.
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162
IEEE TRANSACTIONS ON INDUSTRY APPLICATIONS, VOL. 52, NO. 1, JANUARY/FEBRUARY 2016
Kyohei Kiyota (S’12) was born in Chiba, Japan, in
1987. He received the B.S., M.S., and Ph.D. degrees
from Tokyo Institute of Technology, Tokyo, Japan, in
2011, 2013, and 2015, respectively, all in electrical
and electronic engineering.
He is currently with the Department of Electrical and Electronic Engineering, Tokyo Institute of
Technology, where he is engaged in a rare-earthfree motor project. He has been a Research Fellow
of the Japan Society for the Promotion of Science
since 2013. He was a Visiting Research Scholar at
the Memorial University of Newfoundland in 2014.
Takeo Kakishima (S’13) was born in Kanagawa,
Japan, in 1989. He received the B.S. and M.S. degrees in electrical and electronic engineering from
Tokyo Institute of Technology, Tokyo, Japan, in 2013
and 2015, respectively.
He is currently with the Department of Electrical and Electronic Engineering, Tokyo Institute of
Technology, where he is engaged in a rare-earth-free
motor project.
Akira Chiba (S’82–M’88–SM’97–F’07) received
the B.S., M.S., and Ph.D. degrees from Tokyo Institute of Technology, Tokyo, Japan, in 1983, 1985, and
1988, respectively, all in electrical engineering.
In 1988, he joined the Department of Electrical
Engineering, Faculty of Science and Technology,
Tokyo University of Science, as a Research Associate. Since 2010, he has been a Professor with the
Graduate School of Science and Engineering, Tokyo
Institute of Technology. He has been studying
magnetically suspended bearingless ac motors, super
high-speed motor drives, and rare-earth-free motors for hybrid and pure electrical vehicles. He has so far published more than 954 papers, including the first
book on magnetic bearings and bearingless drives in 2005.
Dr. Chiba was a recipient of the Institute of Electrical Engineers of Japan
(IEEJ) Prize Paper Awards in 1998 and 2005. He was also a recipient of the
First Prize Paper Award from the Electric Machines Committee of the IEEE
Industry Applications Society (IAS) in 2011. He has served as Secretary, Vice
Chair, Vice Chair/Chair Elect, and Chair of the Motor Subcommittee of the
IEEE Power and Energy Society (PES) in 2007–2008, 2009–2010, 2011–2012,
and 2013–2014, respectively. He has been a member, Chair, and Past Chair
of the IEEE Nikola Tesla Field Award Committee in 2009–2011, 2012–2013,
and 2014, respectively. He served as Chair of the IEEE IAS Japan Chapter in
2010–2011. He also serves as an Editor and Associate Editor for the IEEE
T RANSACTIONS ON E NERGY C ONVERSION and the IEEE T RANSACTIONS
ON I NDUSTRY A PPLICATIONS , respectively.
M. Azizur Rahman (S’66–M’68–SM’73–F’88–
LF’07) was born in Santahar, Bangladesh, on January 9, 1941. He received the B.Sc.Eng. degree from
the Bangladesh University of Engineering and Technology (BUET), Dhaka, Bangladesh, the M.A.Sc.
degree from the University of Toronto, Toronto, ON,
Canada, and the Ph.D. degree from Carleton University, Ottawa, ON, Canada, in 1962, 1965, and 1968,
respectively, all in electrical engineering.
In 1962, he joined BUET as a Lecturer and was
promoted to Full Professor in 1975. In 1976, he
joined the Memorial University of Newfoundland, St. John’s, NL, Canada,
where he is a Professor and University Research Professor. He has 50 years
of teaching, including about ten years of full-time and concurrent industrial,
utility, and consulting experience at GE, Schenectady, NY, USA, GE Canada,
Peterborough, ON, Canada, Newfoundland Hydro, Dhaka Electric Supply,
Teshmont Consultants, Iron Ore Company of Canada, etc. He has been a
Visiting Professor and Research Fellow at Imperial College London, Technical University of Eindhoven, University of Manitoba, University of Toronto,
Nanyang Technological University, Tokyo Institute of Technology, University
of Hong Kong, Tokyo University of Science, and University of Malaya. He has
published more than 635 papers, in addition to eleven patents, two books, and
five book chapters. His current research interests are in machines, intelligent
control, power systems, digital protection, power electronics, and wireless
communications.
Dr. Rahman is a Registered Professional Engineer in the Province of
Newfoundland and Labrador, Canada. He is a member of the Institution of
Electrical Engineers, Japan; a Fellow of the Institution of Engineering and
Technology, U.K.; a Fellow of the Engineering Institute of Canada; a Life
Fellow of the Institution of Engineers, Bangladesh (IEB); and a Fellow of
the Canadian Academy of Engineering. He has been a recipient of numerous
awards, including the GE Centennial Invention Disclosure Award in 1978, the
IEEE Outstanding Students Counselor’s Award in 1980, the IEEE Notable
Service Award for contributions to IEEE and the engineering professions
in 1987, the IEEE Industry Application Society’s Outstanding Achievement
Award in 1992, the Association of Professional Engineers and Geoscientists
of Newfoundland Merit Award in 1994, the IEEE Canada Outstanding Engineering Educator’s Medal in 1996, the IEEE Third Millennium Medal in
2000, the IEEE Cyril Veinott Electromechanical Energy Conversion Award
in 2003, the IEEE William E. Newell Power Electronics Award in 2004, the
Khwarizmi International Award in 2005, the IEEE Dr.-Ing. Eugine Mittelmann
Achievement Award and the IEEE Richard H. Kaufmann Technical Field
Award in 2007, the A. D. Dunton Award of Distinction in 2008, the IEEE Power
and Energy Society (PES) Distinguished Service Award in 2008, and the IEB
Gold Medal in 2011.