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Design of a Five Phase Brushless DC Moto

Design of a Five-Phase Brushless DC Motor for a
Safety Critical Aerospace Application
Xiaoyan Huang, Member, IEEE, Andrew Goodman, Member, IEEE, Chris Gerada, Member, IEEE,
Youtong Fang, Member, IEEE, and Qinfen Lu, Member, IEEE
Abstract—This paper describes a five-phase brushless dc
(BLDC) motor designed for an electrohydrostatic actuation system
(EHA) suited to the thin and optimized wings. The foundation
of the design is a motor with fault tolerance and high reliability,
compact structure, and low weight. The motor power rating is
12 kW at 12 000 rpm, and a “wet” form of construction is used
where hydraulic oil is present in the motor to reduce the number
of oil seals of the EHA for enhanced reliability and lifetime. The
losses and thermal behavior are evaluated for an optimized design.
Fault tolerance for BLDC motors is discussed. A five-phase motor
has been manufactured, and test results are presented to validate
the design.
Index Terms—Brushless dc (BLDC) motor, electrohydrostatic
actuation system (EHA), fault tolerant.
ITH THE development of new composite materials
thinner wings can be used for either improved aerodynamic efficiency or improved structural efficiency. This will
however severely limit the installation space for flight control
surface actuation systems. The limitation of space will make
the current actuator assemblies infeasible. Since the more electric aircraft concept emerged, electrical actuation system has
been intensively investigated due to advantages in terms of reduced weight, compact structure, easier maintenance, increased
safety, and enhanced reliability [1], [2]. Proposed electrical
solutions employ either an electrohydrostatic actuation system
(EHA) or an electrical mechanical actuation system (EMA).
The EMA uses mechanical gearing and a ball or roller
screw to amplify the motor torque and provide linear actuation.
This approach may lead to difficulty in realizing the safety
requirements of a primary flight control surface where multiple
Manuscript received November 3, 2010; revised April 9, 2011 and June 30,
2011; accepted September 23, 2011. Date of publication December 5, 2011;
date of current version April 13, 2012. This work was supported in part by
the U.K. Government under the DTI CARAD programme and by the National
Natural Science Foundation of China under Grant 51007078.
X. Huang was with the Power Electronics Machines and Control Group,
Department of Electrical and Electronic Engineering, University of Nottingham, Nottingham NG7 2RD, U.K. She is now with the College of Electrical
Engineering, Zhejiang University, Hangzhou 310027, China.
A. Goodman and C. Gerada are with the Power Electronics Machines
and Control Group, Department of Electrical and Electronic Engineering,
University of Nottingham, Nottingham NG7 2RD, U.K.
Y. Fang and Q. Lu are with the College of Electrical Engineering, Zhejiang
University, Hangzhou 310027, China (e-mail: luqinfen@zju.edu.cn).
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/TIE.2011.2172170
actuators are used to ensure continued operation in the event of
failure of one unit. When an EMA fails due to the mechanical
gearing, the result can be a jammed actuator that may also
prevent operation of the backup units. The EHA is a local,
electrically driven hydraulic actuation system, which retains a
similar function to the conventional hydraulic system. Where
multiple EHA actuators are used in parallel on a flight control
surface, failure of one unit can be accommodated using a simple
hydraulic by-pass valve. This leaves the hydraulic ram of the
failed unit free to move as the flight control surface moves
under the control of the remaining healthy units. Thus, an EHA
system is the basis of this study due to its fault tolerance, which
is the key issue in aerospace applications.
The aim of this paper is to design a fault-tolerant electrical
motor for EHA which could be fitted in a thinner and optimized
wing for better aerodynamics.
In the past decades, switch reluctance (SR) motor, permanent
magnet synchronous motor (PMSM), and brushless dc (BLDC)
motor were investigated for aerospace applications [3], [4]. The
SR motor is highlighted for safety critical applications due to
its inherent high reliability. A 25-kW, four-phase, 8/6 SR has
been designed and built by the University of Glasgow [5]. It was
considered to offer the best compromise between performance,
fault tolerance, and drive complexity for use in a large EMA
system. A 22-kW, four-phase, 8/6 high power density SR drive
for the MEA is reported in [6] which also discusses drives with
different number of phases. The higher number of the phases,
the more fault-tolerant the drive system is. However, the powerto-weight ratio reduces as the number of phases increases.
Higher power density can be achieved with PMSM motors.
Fault tolerance can be improved in these machines in a similar
degree as in the SR motor drives [7]. A six-phase, eight-pole
PMSM drive has been reported in [7] which can produce 29%
more torque than its SR counterpart with high fault tolerance. A
PMSM draws sinusoidal currents when supplied by sinusoidal
voltages. This is an advantage of the PMSM over the SR motor,
when a minimum current harmonic content is required [8].
PM motors with different structures had been designed and
evaluated for fault tolerance and high reliability [9], [10].
The BLDC motor and PMSM share certain advantages
and disadvantages. The BLDC motor has the advantage of a
simpler control than the PMSM. No complex microprocessor
is required, which to some degree could reduce the risk of
failure. Multiphase BLDC motors with high power density
and reliability for an EHA are reported [11]–[13]. The BLDC
motor has the potential to provide the highest power density
of any kind of motor because of the trapezoidal current supply
0278-0046/$26.00 © 2011 IEEE
and back electromotive force (EMF). The disadvantage of the
BLDC motor is the high input current harmonic content due to
the nonsinusoidal back EMF and supply current.
In safety critical applications, the fault tolerance and reliability of the whole electrical drive including the motor and power
converter are equally of great importance. In most industrial
applications, the type of the motor is selected first and followed
by the power electronics and control strategy to reach various
requirements. In this design, a different procedure is followed.
A power converter with high reliability and simple control
strategy is first determined. Then, the motor which suits the
power converter and the control strategy with a specific power
rating will be designed.
From the power converter point of view, a matrix converter
is highlighted for high reliability due to the absence of bulky
capacitors which cannot operate reliably in the extremities of
aerospace environmental conditions [14], [15]. The disadvantages are the higher risk of short circuits and more complicated commutation techniques which will lead to the use of a
dedicated powerful microprocessor. In aerospace applications,
the reliability of such a microprocessor can be a significant
drawback. Therefore, the single-sided matrix converter (SSMC)
which keeps the inherent advantages of a matrix converter but
avoids the complex commutation problems is selected as the
motor drive in this application; details are given in [16]. The
SSMC can only provide unidirectional dc current. The BLDC
and SR motors are all suitable candidates for use with the
SSMC. Finally, the BLDC motor was selected over the SR
motor due to its superior power density. In this paper, a BLDC
motor will be designed to work with the SSMC and furthermore
to meet the requirements of aerospace applications.
seen as limiting the diameter of the motor and limiting the
space for the power electronics that is usually mounted
close to the actuator. A compact structure is essential.
“Wet” operation
To minimize the number of sliding surface oil seals in
the actuator for higher reliability and longer life time, the
motor is immersed internally in Skydrol to avoid shaft
Viscous losses
Skydrol produces a hydraulic power loss as it circulates
through the relatively short air gap, and this loss must be
minimized by controlling the rotor dimensions. The loss
is greatest at high speed.
Maximum temperature limit
The maximum operating temperature for Skydrol is
120 ◦ C.
Fast response
For most of the time, the actuator is at the required position, and the motor is running at a low speed (2000 rpm)
with minimum load so that the oil seals are continuously
lubricated so keeping wear to a minimum and extending
operational life. When the new position reference signal
arrives, a higher motor power at high speed (12 000 rpm)
will be produced and then transmitted by the fluid flow
to the actuator to change the position until it reaches the
required position. The motor should have a low inertia to
respond rapidly to changes in load accelerating from 2000
to 12 000 rpm in about 0.5 s. The high-speed operation
lasts only for as long as the high load is required before
the motor returns to the low-speed condition.
Accommodation to be driven from a SSMC
The SSMC only provides a positive current half cycle
in each stator coil. If the same design of motor were
fed from a conventional converter, then it is obvious that
for a given torque level, the peak current level would
be half that of the SSMC fed machine. Correspondingly,
the power loss would also be half that of the SSMC
machine. The design of the motor is therefore required
to accommodate the special form of SSMC power electronics which provides greater inherent commutation and
control reliability for the converter and thus the complete
drive system.
The design objectives of the motor need to be identified before starting the motor design. The conventional design criteria
for the complete drive are listed in Table I which includes the
power supply, electrical performances, and mechanical performances.
The key problems of the motor design are:
1) Fault tolerance and high reliability
Fault tolerance and high reliability are of great importance in aerospace applications. A multiphase motor and
drive will be designed toward this end.
2) High power density
High power (torque)-to-weight (volume) ratio is a key
criterion to save space for aerospace applications. The
effect of the reduction of installation space on motor is
In this paper, a typical motor design procedure is followed.
First, the main dimensions are determined by the power rating,
the empirical current loading, and magnetic loading in this
power range. Then, the stator lamination dimensions can be
calculated based on the specific number of the phases and poles,
followed by the design of the rotor structure. Finally, electrical
power losses and hydraulic losses are then evaluated. The
temperature rise of the motor is investigated using Motorcad.
A. Main Dimensions
The initial design of the BLDC includes the determination
of the main dimensions including stator inner diameter and
Fig. 1. Comparison of output torques for motors with different number of
phases under healthy and faulty conditions.
effective length of the rotor. It should be noted that for the
specific flux density, electric and magnetic loading, and speed,
the output power is proportional to rotor volume. The choice
of stack length and rotor diameter was influenced by the low
inertia, the hydraulic loss, and the critical shaft speed. A small
rotor diameter is required to ensure the low inertia and then the
fast response. Hydraulic loss increases rapidly with diameter of
the rotor and directly with its length. The hydraulic loss can
easily dominate all other losses in the machine at 12 000 rpm.
Therefore, the air gap length has a minimum value. Thus,
magnet thickness is mostly defined and consequently is the
diameter of the solid rotor shaft and rotor core. This in turn
restricts the minimum rotor diameter both through increasing
inter pole magnetic leakage flux and through the shaft critical
speed. As the rotor diameter reduces, the core length and the
distance between the bearing centers increase with a consequent
reduction in the shaft critical speed.
B. Number of Phases and Poles
Since fault tolerance and high reliability are the foundations of the design, the number of phases should therefore be
carefully selected. Generally, motors with a high number of
phases have the potential advantage of higher fault tolerance
and consequently reliability [17]–[19].
Finite-element method models of different phase BLDC
motors were built to investigate their respective performances.
In this section, the inner diameter, stack length, the rotor
structure, and the rotating speed of multiphase BLDC motors
are kept constant. Only the stator structures are affected when
considering different numbers of phases. The comparison is
based on the assumption that the same quantity of copper is
used. The number of turns varies with the number of phases to
keep the amount of copper constant.
Fig. 1 shows the output torques for motors with different
number of phases and unidirectional supply currents under
healthy and faulty conditions. It can be seen that at least two
phases should be conducting at any time for improved reliability of the drive. This allows a phase malfunction while still
maintaining some drive functionality as there will always be at
least one phase conducting current. The requirement for two
conducting phases in an N-phase motor would imply that each
phase current would conduct for 2/N of a cycle. Considering
a four-phase motor, the current pulse conduction angle needs
to be 180◦ , leading to large torque ripple. For five- and sixphase machines, the conduction angles are 144◦ and 120◦ ,
respectively, which are confirmed by the curve shape of back
EMFs. The torque ripple can therefore be reduced significantly.
The disadvantage of the five and six phase is of course the extra
power devices.
Compared to the six-phase drive, the torque produced by the
five-phase drive is 2.98%, 4.19% less respectively at healthy
and faulty condition. However, 20% extra power electronic
components are required.
Since the motor is fed from SSMC, the BLDC performance
under the fault condition with one IGBT of the SSMC shortcircuit need to be investigated. The conduction period of current
in the faulty phase is longer than the healthy phase which results
in a large transient torque at the time. The torque can be double
the rated torque if the faulty phase current cannot be switched
off on time. The transient torque for all phase number drives
rises by approximately 50% under this faulty condition. From
the motor design point of view, the one IGBT short-circuit fault
condition has similar impact on the BLDC motors with different
number of phases.
As a result, the five-phase BLDC motor was finally chosen
as a compromise between fault tolerance, weight, and volume.
Another important parameter that needs to be taken into
consideration is the number of poles. The higher the number
of poles, the smaller is the stator outer diameter for a given
rotor diameter. Furthermore, from the motor point of view, the
copper loss is reduced when increasing the number of poles,
because the end windings become shorter. However, increasing
the number of poles will lead to a higher basic switching
frequency and hence an increase in iron losses. This increased
loss is however mitigated due to the reduced stator back iron
depth and thus reduces to some extent the iron losses. The risk
of demagnetization for the PMs also reduces with the increasing
number of poles. From the power converter point of view, a
higher switching frequency is required for a higher number
of poles. In this project, the switching frequency capability
of the IGBTs on the power board is limited which can be
easily reached if a large number of poles applied due to the
low inductance of the BLDC. A pole number of four was a
reasonable compromise.
C. Stator Lamination
According to the steel datasheet, the steel started to saturate
severely around 1.8 T. As a result, the flux densities in stator
teeth and air gap are 1.8 and 0.75 T, respectively. For specific
flux density required in air gap and stator teeth, the width of
stator teeth was determined by stator inner diameter and the
number of slots.
The slot area was determined by the current density, number
of turns, and slot fill factor. In aerospace application, more
space in the slots needs to be left for the insulation material for
higher reliability. Consequently, a conservative slot fill factor
was used.
The width of stator back iron is one of the key factors
determining the iron losses. The deeper the stator back iron is,
the heavier the total stator laminations are, and in turn, the more
Fig. 4. Peak cogging torque with various magnet arcs.
Fig. 2.
Flux density in the air gap with different magnet thickness.
Fig. 3.
Shapes of the PMs.
Fig. 5. Flux density in the motor no-load (Ia = 0 A).
iron losses. Contradictorily, the wider the stator back iron is,
the smaller the flux density is in the back iron. The width of the
back iron was limited to control the iron losses in the machine.
D. Permanent Magnet Design
The PM thickness is a significant parameter since it affects
the flux density in the air gap, the demagnetization withstand
capability, and the cost. A special feature of this BLDC motor
is the “wet” operation. Therefore, a large air gap will be used to
minimize viscous losses. The thickness of magnet is expected to
be greater than traditional designs. The simulations are carried
out with PMs of various thickness based on the same stator
lamination dimensions and the rotor structure as shown in
Fig. 2. It can be seen that flux density in the air gap (Bg )
increases as the thickness of the PMs increases. However, it
reaches to the maximum 0.75 T at 8 mm before the core starts
to saturate and interpole leakage becomes substantial. This
confirms 8 mm as the maximum sensible value.
A surface-mounted PM was used in this design due to higher
power density, which was slightly modified to an octagon to
prevent sliding between the rotor surface and the PMs as shown
in Fig. 3.
The arc of PMs is also optimized to minimize the torque
ripple. The peak cogging torque with different magnet arcs is
shown in Fig. 4. It can be seen that the peak cogging torque
is relatively small, and therefore the magnets are not skewed
in this application. The 75◦ magnet pole pitch produced the
least cogging torque. However, the flat top area of the back
EMF waveform reduces as the arc reduces. Large torque pulsation will occur during commutation when the arc of the PMs
decreases, which will also reduce the average output torque.
Fig. 6. Flux density in the motor full-load (Ia = 40 A).
Consequently, an arc of 80◦ is a good compromise between the
cogging torque ripple and the commutation torque ripple.
A. Motor Operation
FEM model of the BLDC motor is built. The flux densities
in the motor under no-load and full-load conditions are shown
in Figs. 5 and 6, respectively, which verify the motor design.
The PMs can operate reliably when the current is less than
120 A without demagnetization according to the Fig. 7. If the
motor winding has a turn-turn short circuit, the pulse of current
may be greater than this level. Therefore, demagnetization
might still be possible in this condition; however, the likelihood
of it happening did not compromise the required system reliability level. However, the PMs have the potential risk of demagnetization which may result in the motor failure. Investigation
of the resistance to demagnetization with armature reaction is
thus of great importance to ensure high reliability.
Fig. 7. Flux density in the motor when Ia = 120 A.
Fig. 9.
Stator iron loss versus speed.
tions (12 000 rpm), respectively. However, the loss would be
increased if the winding temperature rises above that which is
B. Iron Losses
Fig. 8. Flux density in the motor at 150 ◦ C.
B. Motor Operation at High Temperatures
Samarium cobalt is selected due to its low temperature
coefficient. FEM simulations under no-load condition were
performed to predict the performance of the motor at 150 ◦ C,
as shown in Fig. 8. The peak value of flux density reduces from
1.57 T at 20 ◦ C to 1.49 T at 150 ◦ C. The average back EMF
reduces from 28.72 to 28.0 V. The torque is proportional to
the back EMF at the same speed and armature current, and
therefore it reduces by 2.57%, which is within the tolerance
range for the design and is easily accommodated with a small
increase in phase current.
The thermal design is critical for this application to get a
compact structure. The losses need to be thoroughly evaluated
for the thermal analysis [20].
The main losses in a BLDC motor consist of the copper, iron,
mechanical losses, and stray load loss. The copper loss can be
limited by reducing the current or the resistance of the winding.
The resistance can be reduced by using conductors with large
diameter and operating under cool temperatures. Clearly, large
slot area is required as well as large machine size. A higher Bg
is an advantage so that the armature current is reduced to deliver
the same torque level. The increased Bg will lead to extra iron
losses. The compromise between size and losses (copper and
iron losses) needs to be made for best motor performance.
A. Copper Loss
For the ideal square wave current supply, the copper loss
can be calculated simply by I 2 · R. The approximate resistance
is estimated to be 0.121 Ω. The copper loss will be 6.05 and
387.2 W at minimum load (2000 rpm) and full-load condi-
The iron losses in the BLDC include the iron losses in
stator, rotor, shaft, retaining can, and the magnets. It is difficult
to calculate the iron losses accurately. Several improved iron
losses prediction methods have been presented based on the
analytic or FEM method in the literature which involves complex magnetic field calculations [21]–[25]. The other method is
to calculate the iron losses based on empirical formula using
the empirical loss data provided by the manufacturer. This was
applied in this design owing to its simplicity. The stator and
rotor peak flux densities used were derived from the FEM
calculations. The stator iron losses at different rotating speeds
are shown in Fig. 9.
It should be pointed out that the method used above is for
sinusoidal current excitation. The BLDC produces nonsinusoidal current waveforms, and the stator winding is fed from
the switched AC source from the SSMC, which will bring high
frequency harmonics into the stator flux. As a result, there is
expected to be modest correlation between the estimated value
using the classic method and the measured iron losses.
C. Retaining Can Eddy Current Losses
The retaining can has generally low electric resistance, and
the loss could be significant. Strength and high resistance
are the important criteria in selecting the retaining can material.
The losses in the can are derived from FEM software as shown
in Figs. 10 and 11.
During most of the operation time, the losses are small
because the flux variation in the can is slight. The greatest
losses occur during the armature current commutation, where
the rapid change of the current in armature winding leads to a
rapid flux variation in the retaining can. Under this condition,
the can losses cannot be ignored. The average losses in one
electrical cycle are 1.5 W and 81.7 W, respectively, under
2000 rpm and 12 000 rpm full-load conditions.
The losses in the PMs are similar to those in the can.
However, the relative longer distance from the stator tooth and
Fig. 10. Retaining can eddy current loss at 2000 rpm.
Fig. 12. Temperature rises of the components.
Fig. 11. Retaining can eddy current loss at 12 000 rpm.
the shielding due to the eddy currents generated in the can make
these losses smaller.The PMs are made into segments to reduce
these losses. The average losses in one electrical cycle are 0.6 W
and 22.8 W, respectively, under 2000 rpm and 12 000 rpm fullload conditions.
D. Rotor and Shaft Losses
Owing to the low rate of flux cutting, the iron losses in the
rotor and shaft have a similar waveform to those of the retaining
can and the PMs, but the magnitudes are much smaller. The
average losses in one electrical cycle are 0.029 W and 3.497 W,
respectively, at 2000 rpm and 12 000 rpm, which are less than
1% of the total losses.
E. Viscous Loss
The Skydrol hydraulic fluid in contact with the rotor surface
generates a viscous drag loss, which can be calculated using the
P = 0.5 · π · ρ · ω 3 · Rr4 · l · Cm
where ρ is the density of the fluid, Rr is the rotor radius, ω
is the angular velocity, Cm is the friction coefficient. Cm can
be expressed as (2), when the fluid is turbulent, with a high
Reynolds number [26]
Cm = 0.065(lg /Rr )0.3 · Re−0.2
ρ · · R r · lg
Re =
here, lg is the air gap length and μ is the fluid viscosity.
Consequently, the viscous loss is 6 W and 817 W at 2000 rpm
and 12 000 rpm, respectively. The losses are summarized in the
Table II.
The power-to-weight/volume ratio is mainly restrained by
the operating environment temperature and the cooling method
once the power rate is determined. The thermal performance
of the motor should be carefully examined to maximize the
In this design, the weight/volume is compromised by the
requirement of fault tolerance and hydraulic losses. A Motorcad
thermal model was built based on the motor dimensions and the
losses mentioned above. The duty cycle of the motor varies with
the actuation cycle, which cannot be predicted precisely. In this
design, the maximum operation period of 60 s is expected at
high speed under full-load conditions. The simulation considered the worst case scenario. That means 180 s per cycle; the
motor started from 2000 rpm at 0 ∼ 20 s, then accelerated to
12 000 rpm in 0.5 s and maintained speed for 60 s, decelerated
to 2000 rpm and maintained speed for another 100 s. The
ambient temperature was set to 70 ◦ C at the beginning.
The temperature rise waveforms of the components including
stator winding and housing are shown in Fig. 12. The maximum
temperature rise occurred at the rotor surface because of the
most significant viscous loss. All the components were operated
within the temperature limit, which verifies the thermal design
of the motor.
It can be seen that the maximum temperature designed for is
around 100 ◦ C which is less than the temperature limits of the
components used. The margin was left by the design for three
Fig. 13. BLDC motor prototype. (a) Stator. (b) Shaft, rotor with magnets. (c)
Housing and end plates. (d) Assembled BLDC motor.
reasons. The first reason is the actual losses may be greater
than the predicted value due to the difficulty in predicting
the iron losses accurately. Furthermore, it can be noted that
the temperature of the Skydrol was set to 70 ◦ C during the
simulations; however, in reality, a higher temperature than the
ambient temperature is expected as this will carry the heat generated in the pump. The third reason is that the heat dissipation
is greatly dependent on the leakage flow rate of the Skydrol
from pump to the motor. The flow rate could not be precisely
predicted at this stage of the design as it relies partly on the
high pressure leakage past the pump pistons. Consequently,
the maximum temperature of 100 ◦ C was expected as a worst
case scenario. Maximum pump loss allowed in this design is
1197 W.
Fig. 14.
Five-phase back EMF waveforms at 2000 rpm.
Skydrol oil, the radial fins are designed to accommodate the
hydraulic system losses as well as the motor losses as shown in
Fig. 13(d).
A. Results to Verify the Electromagnetic Design
The stator stack is built from 0.35 mm thick laminations
as shown in Fig. 13(a). The ideal lamination material should
have low magnetic reluctance but high electrical resistance.
The single layer winding is employed for higher reliability and
smaller magnetic coupling between phases. Particular attention
should be given to the insulation material, which must be
Skydrol proof.
Owing to the large effective air gap length in the design, a
solid rotor could be employed for better mechanical integrity.
The material used for the solid rotor is EN24 steel. The assembled rotor, shaft, the magnet segments, and the retaining can
are shown in Fig. 13(b). The 1-mm retaining can is designed to
withstand the centrifugal force of the magnet segments at high
speeds. The material used for the can is nonmagnetic stainless
All the remaining components including the housing, two
end plates, and the resolver housing are shown in Fig. 13(c).
A special feature of the BLDC is the cold Skydrol oil fed from
the pump along the shaft through ten holes in the housing and
back to the reservoir. Owing to the method of cooling with the
The first step of the test was measuring the five-phase resistances and inductance which were recorded in Table III. The
resistances are 17.7% higher than the predicted value because
the resistance in the end winding is difficult to be predicted
accurately as well as the slot fill factor until the coils have
been assembled in the motor due to the practical limitations of
winding the machine. The end windings are actually longer than
anticipated as can be seen in Fig. 13(a). The average inductance
of the five-phase windings is 8.13% less than the predicted
value. The inductances varied slightly with the rotor position
because of the interaction between slot opening and the gap
between the PMs. Consequently, the inductances are slightly
different from each other at any rotor position. The measured
resistances and inductances indicated that the windings were
correctly constructed.
The open circuit test was done at 2000 rpm. The measured
five-phase back EMF waveforms at 2000 rpm were presented
in Fig. 14 which verified that the designed flux density in
the air gap was reached and the windings were correctly
constructed. The data were recorded using two synchronized
scopes. Each scope has a maximum of four channels. Therefore,
Fig. 16. Comparison of tested and predicted viscous loss.
Fig. 15. Tested iron losses, windage, and friction losses.
two synchronized scopes are required to record the correct back
EMF sequence. There should be two pulses in one mechanical
revolution due to the two pole pairs in the BLDC motor. It can
be seen from Fig. 14, there are two pulses in 30 ms which
conformed to the 2000-rpm operation. The average value of the
flat top area is 27.5 V. Compared to the simulation result of
28.7 V, it is 4.12% less. The reason is the unexpected large
voltage dip in the middle of the flat top area which is caused
by the gaps between the PMs and the slot opening. This voltage
dip also introduces undesirable torque ripple.
B. Motor Losses Determination
1) Iron Losses (Stator Iron, Retaining Can, Rotor and Shaft,
and the PM Losses): It is difficult to separate the stator iron loss
with the retaining can, rotor, shaft, and the PMs losses because
they are all generated by the same source. During the test, they
were all considered as one iron loss, which can be obtained
from the open circuit test results. In the BLDC motor, the flux
density is mainly determined by the PMs. Thus, once the motor
was constructed, the iron losses mainly varied with the speed
of the motor. Consequently, the iron losses cannot be separated
from the friction and windage losses as in an induction motor.
The mechanical power of the BLDC motor driven by the AC
motor at different speeds without the Skydrol immersion was
recorded as shown in Fig. 15. These included the iron losses,
friction, and windage losses. The losses, as expected, increase
more rapidly as the speed increases.
It can be seen from the figure that the tested loss is greater
than the predicted loss. This is mainly because of the difficulty in predicting the stator core losses accurately. The rotor
and shaft losses are very small compared to other losses as
simulated in the FEM software. However, the predicted losses
do not include the bearing losses, which is not insignificant.
Overall, therefore, the agreement is good between prediction
and measurement.
It should be noted that the rotor and shaft losses, the retaining
can, and the PM losses both as tested here and as predicted
are lower than for the fully loaded situation because there is
no commutation of the current during the test. However, from
the simulated results described above, large power loss pulses
during commutation would be expected at commutation, which
will increase the average losses above those measured.
2) Fluid Viscous Loss: The mechanical input power in the
open circuit test with Skydrol immersion includes the viscous
loss, iron losses, and friction and windage losses. As a consequence, the viscous loss could be found by subtracting noload mechanical losses for the nonimmersed condition from
the immersed condition. The results were compared to the
predicted viscous hydraulic loss in Fig. 16. The tested losses
are greater than the predicted losses because the viscosity of the
Skydrol is sensitive to the temperature. The viscosity at 20 ◦ C
was used in the predicted losses. The actual room temperature
during the test was 17.2 ◦ C which will increase the viscosity
of the fluid and lead to a larger viscous loss. Furthermore, the
measured power loss should be greater than the predicted loss
because the prediction only deals with fluid in the air gap and
ignores the ends of the rotor, the shaft, and the loss in the
Skydrol flooded bearings. Overall, the predicted hydraulic loss
can be seen to be in good agreement with predictions.
3) Thermal Design: It is clear, however, that loss prediction
is in agreement with experimental results though in all cases
the losses are underestimated. The tendency of the losses to
increase with the speed was as expected. Most significant is the
extra stator conductor loss on full load. However, a good safety
margin was left for the thermal design. The actual maximum
temperature rise predicted was 30 ◦ C with a 20 ◦ C margin left.
The new actual power loss estimates suggest that 7 ◦ C to 8 ◦ C of
the 20 ◦ C margin is consumed in the extra loss which still leaves
a safety margin for underestimation of temperatures, cooling,
and pump losses. It can be therefore seen that the BLDC motor
is capable of providing the designed performance.
A 12-kW, 12 000-rpm, five-phase, four-pole BLDC motor
was designed and manufactured for the EHA application. The
performance of the motor was determined by magnetic circuit
analysis and then FEM simulation. Although the motor performance in terms of torque per unit volume is sacrificed partly for
the aerospace reliability requirements and partly for the SSMC
limitations, the motor provides with the SSMC a highly faulttolerant drive system. Important design features based on the
FEM simulation results include the 80◦ arc PMs. These are a
compromise between considerations of cogging torque and the
commutation torque ripple.
The losses of the motor were investigated individually and
controlled. At high speed, the viscous loss is the most significant loss and degrades the motor efficiency. The thermal
design of the motor has been developed using Motorcad. Even
when the motor operates at the maximum ambient temperature,
the steady-state temperature will be within the limit set by the
It should be noted that the motor dimensions have been
optimized, but the size and volume of the housing particularly
the cooling fins turned out to be a large proportion of the
whole assembled motor. Further research into improved heat
dissipation and more accurate estimation of the hydraulic pump
losses considering the sizes of the heat dissipation components
would be useful.
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Xiaoyan Huang (M’09) received the B.E. degree, from Zhejiang University, Hangzhou, China,
in 2003, and the Ph.D. degree in electrical machines and drives from the University of Nottingham,
Nottingham, U.K., in 2008.
From 2008 to 2009, she was a Research Fellow
with the University of Nottingham. Currently, she is
a Lecturer with the College of Electrical Engineering, Zhejiang University, where she is working on
electrical machines and drives. Her research interests are permanent magnet machines and drives for
aerospace and traction applications, and generator system for urban networks.
Andrew Goodman (M’09) received the M.S. and
Ph.D. degrees in electrical and electronic engineering from the University of Nottingham, Nottingham,
U.K., in 2002 and 2007, respectively.
He is currently a Research Fellow in the Power
Electronics, Machines, and Control Group at the
University of Nottingham. His research interests include grid interface converters and motor drives.
Chris Gerada (M’05) received the Ph.D. degree
in numerical modelling of electrical machines from
the University of Nottingham, Nottingham, U.K.,
in 2005.
He subsequently worked as a Researcher at
Nottingham on high performance electrical drives
and on the design and modeling of electromagnetic
actuators for aerospace applications. He is currently
an Associate Professor in Electrical Machines within
the Power Electronics Machine and Control research
group at Nottingham. He is also the Project Manager
of the GE Aviation Strategic Partnership, and his research interests include
high-performance electric drives and machines.
Youtong Fang (M’11) received the B.S. degree and
Ph.D. degree in electrical engineering from Hebei
University of Technology, Hebei, China, in 1984 and
2001, respectively.
Currently, he is a professor with the College of Electrical Engineering, Zhejiang University,
Hangzhou, China. His research interests include
the application, control, and design of electrical
Qinfen Lu (M’10) received the B.E., M.E., and
Ph.D. degrees from Zhejiang University, Hangzhou,
China, in 1996, 1999, and 2005, respectively.
Since 1999, she has been with the College of
Electrical Engineering, Zhejiang University, where
she is currently an Associate Professor. Her research
interests include analysis and control of linear machines, permanent magnet machines, etc.