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Design and Performance of a High-Speed Permanent Magnet Generator with
Amorphous Alloy Magnetic Core for Aerospace Applications
Article in IEEE Transactions on Industrial Electronics · March 2019
DOI: 10.1109/TIE.2019.2905806
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Viacheslav Vavilov
Università di Pisa
Ufa State Aviation Technical University
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Denis Gusakov
Ufa State Aviation Technical University
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IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS
Design and Performance of a High-Speed
Permanent Magnet Generator with
Amorphous Alloy Magnetic Core for
Aerospace Applications
Flyur Ismagilov, Luca Papini, Member IEEE, Viacheslav Vavilov, Member IEEE,
and Denis Gusakov, Member IEEE
Abstract—This paper deals with the design,
prototyping, and testing of a high-speed permanent
magnet generator with a stator magnetic core made of
amorphous magnetic material for aerospace applications.
The target operative speed range is 30,000-60,000 rpm
with an output power range between 50 kW and 300 kW.
Two types of winding are examined and compared: toothcoil and distributed. Special attention is dedicated to the
rotor dynamics aspects of the solid disk magnet rotor
assembly. The multidisciplinary design process is
described. Electromagnetic calculations are provided for
different operative conditions where the magnetic field
and loss distribution are determined. Based on the design
results, a full-size prototype rated at 120 kW and 60,000
rpm rotational speed featuring a stator core made of
amorphous magnetic material has been developed and
tested. The test results proved the high efficiency
achievable with the described technical solutions. It has
been proved that the use of amorphous magnetic
materials in high-speed electrical machines enables to
minimize losses in the stator magnetic core by 5-7 times in
comparison with traditional materials.
Index
Terms—Amorphous
Permanent magnet machines.
magnetic
materials;
I. INTRODUCTION
H
IGH-speed synchronous electrical machines with
permanent magnets (HSEMPMs) operating at a rotational
speed of 30,000-60,000 rpm and a power of 50-300 kW are of
high interest for the aerospace industry. The interest in that
speed-power range is mainly due to the capabilities of direct
connection to the aircraft power plant. This allows to eliminate
the gearbox and thereby to increase the reliability of the entire
Manuscript received April 17, 2018; revised January 23, 2019;
accepted February 23, 2019. This work was supported by the Russian
Science Foundation, project № 17-79-20027.
F. R. Ismagilov, V. E. Vavilov and D. V. Gusakov are with the
Department of Electromechanics, Ufa State Aviation Technical
University, Ufa, Russia (corresponding author to provide phone:
+79174014545; e-mail: [email protected]).
L. Papini is with the University of Nottingham, Nottingham, United
Kingdom (e-mail: [email protected]).
power supply system. The high rotational speed of the
HSEMPM rotor makes it possible to maximize the electrical
power generated while minimizing the weight of the electric
machine (EM), which is one of the main criteria determining
the EM applicability in the aerospace industry.
HSEMPM has a great potential for use in both traditional
aircrafts and in aircrafts with a hybrid power plant. The
concept of More Electric Aircraft emphasizes the
implementation of high performance EMs [1]-[3]. Therefore,
special requirements on the efficiency are imposed on
HSEMPM in addition to the minimum weight, specifications
on the volume, and reliability capabilities. Improving the
energy efficiency of HSEMPM directly translate in the aircraft
fuel consumption reduction and increase in the efficiency of
the whole power supply system. Indirectly, the energy
efficiency requirement leads to minimize the amount of
refrigerant required for heat extraction from the EMs. The
latter is an extremely important factor determined by a limited
amount of cold resources on board of an aircraft, for example
oil. All these advantages allow to increase the environmental
friendliness of aircrafts as well as to expand their
functionality.
The study of loss reduction techniques in HSEMPM [4]-[7]
have become an important topic for academia and industry as
a consequence of the above considerations. General
approaches to loss analysis in HSEMPM, studies of losses in
permanent magnets of the rotor, aerodynamic losses, as well
as losses in the HSEMPM winding are given in [4]-[6].
Studies of losses in HSEMPM with a tooth-coil winding are
given in [8], [9]. The advantage of tooth-coil winding for
electrical machine, in addition to the one listed in [8] and [9],
is the minimum winding end-turn length, which makes them
sometimes an effective solution for the aerospace industry
[10]. However, a significant disadvantage is the increase of
eddy currents losses in permanent magnets caused by the air
gap spatial harmonics. Therefore, it is necessary to search for
a balance between the overall dimensions and efficiency of
HSEMPM when choosing the winding arrangement.
Analysis of works [4]-[13] shows that the total losses of
HSEMPM with a power of 120 kW and a rotational speed of
30,000-60,000 rpm are about 2-8 % when using a distributed
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winding while increase up to 6-9 % with a tooth-coil winding.
The rough distribution of losses in the active elements for
HSEMPM with a rotational speed of 30,000-60,000 rpm and a
power of 50-300 kW with a tooth-coil winding and distributed
winding is given in the pie chart of Fig. 1. As can be seen
comparing the two charts, the main losses in the range of the
indicated powers and rotational speeds are developed in the
stator magnetic core and stator windings. The iron losses in
the stator of electrical machines rated at 120 kW - 60,000 rpm
with a 0.1-0.18 mm sheet thickness of the silicon steels stator
core, reach 700- 750 W at a frequency of 1000 Hz and flux
density of 1.2-1.3 T [6].
The study here presented aims to demonstrate an effective
technique to minimize the losses in the stator magnetic core.
The initial goal that has been set is to reduce the losses in the
stator magnetic core at least 4 times using amorphous
magnetic materials (AMMs). It is technologically difficult to
make a magnetic core from AMM tape with thickness of 2530 μm. In addition, the literature lack in publications that
provides an effective technological solution for the realization
of a stator structure out of an AMM tape. In this paper,
practical results and measurement data are obtained based on
the theoretical analysis presented in [14], where the in-house
technology for manufacturing a stator with AMM has been
developed and described.
The design, manufacturing and testing of a 120-kW
60,000-rpm HSEMPM for aerospace applications is here
presented featuring a stator magnetic core made of AMM.
EMs) with AMM magnetic core for a similar power-speed
node are not described in the literature, showing the novelty of
the achieved results. The maximum power of known EMs
with a magnetic core made of AMM does not exceed 20 kW
[27]. After a description of the technology around the AMMs
of Section II, the multidisciplinary design process and the
refined design of the prototyped HSEMPM are described in
Section III. Based on the calculations, a full-size prototype of
HSEMPM has been manufactured and the test results are
given in Section IV. According to the experimental results it
was found that the losses in the stator magnetic core are no
more than 100 W at a frequency of 1000 Hz and flux density
of 1.3 T.
II. AMORPHOUS MAGNETIC MATERIAL: TECHNOLOGY
REVIEW AND CHALLENGES
AMM consists of steel rolled to relatively thin sheets.
Different manufacturing process are used, among which the
annealing, and then rapidly spray quenched in liquid nitrogen
Fig. 1. Distribution of losses in HSEMPM with a toothcoil winding (right) and with distributed winding (left)
being the most common. The result is a steel characterized by
a disordered crystals (amorphous) state. The steel features
reduced hysteresis losses and eddy current losses, the latter
mainly due to an increase in resistivity. However, the
manufacturing of core for EMs or transformers is challenging
due to the limited width of the strips that can be nowadays
produced. The main problem that prevented the wide use of
the AMM in HSEMPM is the lack of an economically and
technologically efficient solution for manufacturing the stator
core. Traditional stator manufacturing techniques such as
stamping, laser or electro-erosion cutting are ineffective to
produce the AMM stator cores due to several technical
peculiarities that define the entire process of the design
technology for EMs with AMM:
– The AMM tape thickness is 25-30 µm and the AMM
layers in the stator core are not isolated each other. The
stacking factor does not exceed 0.8 due to various
technological defects and impurities in the AMM.
– The AMM produced by industry has a low saturation
magnetic flux density which varies from 1.3 to 1.55 T. This
value is 60 % lower than that of various cobalt alloys. The
alloying of AMM with Cobalt makes it possible to increase
the magnetic flux density up to 1.8 T. However, this also
increases the specific losses due to the increase in electrical
conductivity. Therefore, it is necessary to limit magnetic flux
density in the air gap when designing the EM with AMM.
– The Vickers hardness of AMM exceeds 1000 HV. This
does not allow to effectively use stamping when creating the
magnetic cores from AMM.
– Magnetic cores made of AMM have a high
magnetostriction coefficient of 26·10-6. Because of this,
magnetic cores made of AMM produce more noise during
operation than conventional magnetic cores.
– Mechanical and temperature effects can significantly
change the magnetic properties of AMM. During laser cutting,
local overheating of the AMM tape might occurs, leading to
local short circuit between AMM sheets in the stator core and,
as a result, to an increase in the eddy-current losses.
In literature can be found examples of AMM used for
manufacturing cores for electro-magneto-mechanical devices.
Dr. Wenming Tong et al. [15] demonstrated several EMs with
the AMM stator core produced by using the electro-erosion
machining. The manufacturing technology of the AMM
magnetic cores for the axial flux EM is presented in [16], [17].
Axial flux EM with wound AMM stator core is investigated in
[18]; radial flux EM and with a wound AMM magnetic core
are considered in [19]. The manufacturing technology of the
stator teeth from separate rectangular AMM sheets is
presented in [20]. Various technologies of using AMM in the
radial flux EM with a power of up to 15 kW are presented in
[21]. Studies of the slotless EM structures with the AMM
magnetic core power of up to 100 W and a rotational speed of
up to 1,000,000 rpm are presented in [22]-[24]. The axial-gap
EMs with the AMM magnetic core are presented in [25].
In certain cases, individual geometric figures can be made
by stamping the AMM tape into the desired shape before
being assembled together in a single structure. According to
this technology, Radam Motors Company created radial flux
EMs. This technology has a special advantage for EMs
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featuring tooth-coil winding structure as the coils can be premanufactured and subsequently mounted around each
individual tooth, maximizing the slot fill factor. On the other
hand, the complexity in the manufacturing results the main
disadvantage.
One of the promising areas of application for the AMM is
the design of the modular stator cores from the AMM
individual elements. The AMM magnetic core elements of the
HSEMPM modular structure are made from assembly of
simple geometric shapes as shown in Fig. 2. The cores are
produced from the AMM tape where the subsequent stacking
and gluing of sector elements leads to a single magnetic core.
This technology, which allows to achieve stacking factor of
0.8, has a few drawbacks. The presence of additional air gaps
and the difficulty of providing reliable fastenings of the stator
core elements are the main problems faced. The authors
carried out detailed studies of this technology. As a result, one
of its significant disadvantages was revealed. For radial flux
EM made according to the above described technology
featuring an axial stator core length of more than 10 mm, the
eddy-current losses in the stator core become quite significant
[26]. This leads to the opposite effect with respect to the
desired one: the stator core losses were not decreased but
increased.
Therefore, we propose a structure where the single stator
core building blocks are AMM magnetic cores with a
maximum axial sector length of 5 mm. The axial stacking
factor of magnetic cores that can be achieved is 0.7-0.75. This
approach is effective in the creation of both slotted and
slotless EMs with AMM stator core. Moreover, in the
proposed manufacturing technology for a stator magnetic core
made of AMM, the individual elements of the magnetic core
are independently wound, followed by their gluing and
impregnation as described in detail in [14].
III. HSEMPM DESIGN
The design process of EM with AMM stator core requires
the selection of the most suitable material. Among the range of
high performance AMMs, the amorphous alloys 5BDSR and
1SR grade, produced by the “Asha Metallurgical Plant”, and
electrical steel grade AMAG321, produced by the OJSC
“Mstator” are considered. In Table I, the main characteristics
of AMMs are presented and compared with high-performance
electrical steel commonly used for high performance EMs.
The magnetization curves of the AMMs can be found in [28].
Based on the criteria of the lowest specific losses, amorphous
a
b
Fig. 2. AMM individual element (a) and assembly of
simple geometric shapes in technological mandrel (b)
TABLE I
MATERIALS COMPARISON
Properties
Flux density, T
Loss density, W/Kg
(1.3 T, 400 Hz)
Mass density, g/cm3
5BDSR
1SR
AMA
G
321
VACO
FLUX50
HIPERCO
40
1.3
1.5
1.8
2,35
2,4
1.1
3
6
28
17
7.6
7.6
7.3
8,12
8,2
alloy 5BDSR has been selected for manufacturing the stator
core.
Being the iron losses density dependent on the operative
flux density and frequency, preliminary considerations has
been done to minimize the losses. The high mechanical
rotational speed imposes high fundamental electrical
frequency in the iron cores which is minimized by electing a
2-pole structure of the rotor. A preliminary design of the
machine had been performed considering a general sizing
technique [29]. The geometrical dimensions of the stator
magnetic core of the HSEMPM were chosen in a way that the
flux density in the stator teeth and back iron is identical.
Whereas the specific losses in the stator teeth and back iron
will be the same, the total losses in the stator core will depend
on the weight of the teeth and the stator back iron. In the
HSEMPM, the stator back iron is manufactured separately
from the teeth. The weight of the back iron of the stator
magnetic core at the end of the initial design process results
30% less with respect to the one of the stator teeth part. The
small difference in weight allows to consider the losses
distribution in the stator core to be practically uniform.
Furthermore, considering a three-phase terminal system, it is
important to choose the minimum number of slots to reduce
the economic impact of stator manufacturing on the
production of HSEMPM with amorphous alloy magnetic core.
Therefore, a HSEMPM with 6 slots was selected. The winding
topology selection is later described in the section however
being limited to the structures possible with 6 stator slots. The
HSEMPM reference structure is shown in Fig. 3. The details
of the rotor and winding design process are presented in the
following subsections.
A. Rotor Design.
The rotor structure of the HSEMPM was made from 6
cylindrical magnets (SmCo brand, Br=1.07 T, Hc=756 kA/m)
with a diameter of 60 mm and an axial length of 25 mm. The
rotor magnets are axially segmented to reduce the eddy
currents losses in permanent magnets caused by spatial and
temporal air gap harmonics. Mechanical calculations highlight
Fig. 3. The HSEMPM design. Here: 1 – Stator axial
stacks made of AMM; 2 – PMs; 3 – rotor sleeve; 4 –
wedge; 5 – cooling channels; 6 – housing; 7 – windings;
8 – bearings; 9 – sealing.
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the need of containment sleeve to guarantee the integrity of
the rotor at high speed due to the poor mechanical properties
of the permanent magnet material. The rotor sleeve thickness
is calculated considering the model presented in [29] with a
safety factor ksafe=2.5. Inconel 718 is selected as the material
for the containment sleeve whose thickness has been
computed to be 4 mm. Having secured the integrity of the
rotor, the choice of the bearing supports was carried out
accounting for the rotor dynamics aspects, computed by means
of Ansys software. The dynamics of the rotating structure has
been predicted considering the dimensions of the turbine
wheel, the compressor wheel, and the rotor structure of the
generator being mounted on a common shaft.
The gas turbine has a diameter of 131 mm with a length of
65 mm. It is made of high-alloy structural, corrosion-resistant
and heat-resistant steel 07Х16Н6 (analogue of 301S21) with
the strength of 1180 MPa. The compressor is made of BT6C
(analogue of IMI318ELI) with the strength of 980 MPa. The
diameter of the compressor impeller is 122 mm with a length
of 57 mm. Fig. 4 shows the three-dimensional models of a gas
turbine and a compressor.
The vibrations which the turbine transmits to a single shaft
structure have been considered at the design stage. The rotor
dynamics largely depends not only on the type of bearing
supports, but also by the length of the rotor, and hence by the
winding end-turn length. Both distributed and tooth-coil
windings were considered at this stage to compute the critical
rotational speeds of the rotor. The comparison of the
performances between the uses of the different winding
configuration is discussed further focusing on their impact on
the losses distribution, weight and efficiency. At this stage, the
parameters of interest are the rotor length and the total axial
length of the EM including the end-winding as strongly impact
on the bearing location and therefore on the rotor-dynamic
performance. According to the calculations, the rotor length
was computed to 274 mm for tooth-coil winding, while
resulting in 314 mm if distributed winding were considered.
The winding end-turn length of the HSEMPM is 35 mm and
15 mm for distributed and tooth-coil winding topology,
respectively.
Having defined the axial length for both the rotor and the
overall structure, different bearing technology and their
arrangements have been investigated. Two variants of rotor
dynamics calculation were considered featuring different
bearing arrangements: a two-bearing scheme (bearings are
located only on the HSEMPM rotor) and three-bearing scheme
(bearings are located between the turbine, the compressor and
the HSEMPM rotor). The axial cross-section of the rotor
assembly is shown in Fig. 5.a. It was found that the use of
mechanical bearings is ensuring the operation of HSEMPM at
subcritical speeds in the entire operative speed range. Gas
bearings were also considered at the design stage but they do
not guarantee safe subcritical rotor operation for the
HSEMPM equipped with the tooth-coil winding due to their
lower stiffness. The use of magnetic bearings could be
effective, however featuring higher cost and control
complexity. Mechanical rolling bearings are the selected
topology as they allows to achieve safe subcritical rotational
speeds of the rotor regardless of the winding type. Ball
bearings with oil lubrication were finally selected for the
prototype. The use of oil lubricant reduces both efficiency and
reliability of HSEMPM but is necessary to guarantee safe
rotor-dynamic operative conditions. Fig. 5.b and Fig. 5.c
shows the results of rotor dynamics calculations for a threebearing and two-bearing schemes, respectively. From the
calculations, the three-bearing scheme turned up more
effective as guaranteeing wider subcritical operative range and
results in a stiffer structure. The general view of HSEMPM is
shown in Fig. 6.
a
b
c
Fig. 5. The results of rotor dynamics calculations. a –
rotor dimension; b – bending vibrations of the rotor at
the first critical speed for three-bearing arrangement; c
– bending vibrations of the rotor at the first critical
speed for two-bearing arrangement.
Fig. 4. 3D model of gas turbine (a) and compressor (b)
a
b
Fig. 6. The general view of HSEMPM (a) and its
dimensions (b)
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B. Winding Selection
Considering that the different winding structures do not
impact on the rotor-dynamic aspects, according to the previous
considerations, their selection is mainly driven by the losses
minimization. The stator winding type strongly impact on the
HSEMPM losses distribution and, consequently, on the
efficiency. At this stage of the design process, it is expediently
decided to consider the use of different types of winding based
on the electromagnetic characteristics of HSEMPM. A toothcoil and a distributed winding are considered when analyzing
the windings type. Table II show the calculation results
obtained considering the above-mentioned winding
arrangements.
The winding end-turn length is an important factor as it
impacts on the total length of the machine, therefore affecting
the rotor-dynamics performance as discussed in the rotor
design section. However, from the electromagnetic point of
view, the tooth coil winding topology raises several problems,
the main one being a significant amount of eddy current losses
induced in the rotor sleeve, which reaches 230 W. The
electrical conductivity of the rotor sleeve (Inconel 718) is 1.25
μΩm at 20 oC. The losses in the sleeve can be reduced to 2022 W with the use of carbon fiber sleeve which, however, does
not provide the necessary mechanical strength to the rotor.
Furthermore, there are significant eddy currents losses in
permanent magnets that occur with the use of tooth-coil
winding. The segmentation of the magnets in the axial
direction, even if this complicates the rotor manufacturing
process, allow a significant reduction of the induced eddy
TABLE II
PARAMETERS OF HSEMPM
Parameters
Distributed
winding
Power, kW
120
Number of rotor poles
2
Number of stator slots
6
Rotational speed, rpm
60,000
RMS phase idling voltage, V
125
RMS phase voltage at rated load (cosφ=1), V 121
RMS phase voltage at rated load (cosφ=0.8)
115.71
RMS phase rated current, (cosφ=1), A
342
RMS phase rated current, (cosφ=0.8), A
422
Current density, A/mm2
10
Short-circuit current, A
1600
Number of turns in phase
4
Magnetic flux density in the stator back, T
1.25
Magnetic flux density in the stator teeth, T
1.25
Magnetic flux density in the air gap, T
0.6
Active stator length, mm
150
Inner stator diameter, mm
72
Outer stator diameter, mm
160
Rotor outer diameter, mm
60
Tooth width, mm
20
Tooth height, mm
19
End-winding length, mm
35
AMM sheet thickness
25 microns
Saturation flux density of the AMM, T
1.35
Permanent magnet type, Br(T)/Hc(kA/m)
SmCo,
1.07/756
Thickness of the sleeve, mm
4
Material of the sleeve
Inconel 718
Weight of active parts, kg
21
Stator length including end-winding, mm
220
Refrigerant
Air
Inlet refrigerant temperature, oC
20
Tooth-coil
winding
120
2
6
60,000
127
117
111
342
429
14.3
1320
8
1.25
1.25
0.6
150
72
160
60
20
19
15
25 microns
1.35
SmCo,
1.07/756
4
Inconel 718
18
180
Air
20
current in the rotor permanent magnet and therefore a decrease
of the rotor losses. Moreover, the winding coefficient for
tooth-coil winding is 0.5 for the 6-slot stator and 2-pole rotor,
while being unitary for a 1 slot per pole per phase distributed
winding arrangement. This reduces the efficiency and, as can
be seen from the calculations, leads to an increase of the
current density in the winding by a factor of 1.4.
The electromagnetic calculations of the final structure
selected of the HSEMPM were performed considering four
operating modes: idling, three-phase short circuit, under load
with cos(φ)=1 and cos(φ)=0.8.
Fig. 7.a and Fig. 7.b shows the results of the computer
simulation for the machine featuring distributed winding
structure where the magnetic field distribution and losses in
the permanent magnets and sleeve of the rotor are presented,
respectively.
The distribution of eddy currents losses in the rotor and the
sleeve of tooth-coil and distributed windings are shown for
comparison in Fig. 7.b and Fig. 7.d, respectively. The
difference in both distribution and amplitude of the losses is
due to the higher space harmonics of the air gap field imposed
by a tooth-coil winding with a fractional number of slots with
respect to the distributed winding layout. Consequently, the
losses in rotor permanent magnets and sleeve results 5-8 times
higher in HSEMPM with the tooth-coil winding with respect
to the one in HSEMPM with distributed winding. From the
simulation results it was found that the eddy currents losses in
a
b
c
d
Fig. 7. Results of computer modeling of 120 kW
HSEMPM at rated load and speed: a – flux density
distribution for distributed winding structure; b – rotor
eddy current losses for distributed winding structure; c –
flux density distribution for tooth-coil winding structure;
d – rotor eddy current losses for tooth-coil winding
structure.
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permanent magnets caused by the spatial harmonics increase
by 10% at cos(φ)=0.8. In Fig. 7.a and Fig. 7.c it is possible to
notice that the maximum value of the magnetic flux density in
the teeth of HSEMPM does not exceed 1.25 T. Furthermore,
the flux density in the stator teeth and the stator back is evenly
distributed.
In general, the results of computer modeling prove the
operability of the chosen HSEMPM design and the possibility
of its practical implementation. Based on all the above
arguments and in combination with the rotor-dynamics results,
the distributed winding topology is chosen for the prototype.
More in detail, a 1 slot per pole per phase distributed winding
with a rated current density of 10 A/mm2 is selected and
manufactured from a stranded wire with the diameter of 1.6
mm. The wire diameter was chosen based on the skin depth
penetration of copper at the maximum operative speed, aiming
to minimize the skin effect losses. The insulation of the wire
was chosen to be polyamide featuring a temperature index of
220 °C.
The use of a magnetic core made of AMM, a novel
technology in the field of power engineering, does not allow to
determine the losses with enough accuracy due to the limited
experience and shortage of reliable experimental data.
Therefore, the thermal calculations of the HSEMPM are
carried out after the experimental studies since the exact loss
value are required as input of the thermal model.
presented in Fig. 8.b. The manufactured winding end-turn
length achieved by using a distributed winding was physically
measured to be 37 mm. It can be clearly seen from Fig. 8.b
that the stator magnetic core is recruited from several sectors
in the axial direction. This is done to reduce the eddy current
losses, similarly to what is done with laminations in
conventional magnetic cores, however leading to a more
complex assembly procedure since it was performed manually
for this experimental prototype. The automation and
improvement of the assembly stage for production of stator
magnetic cores is a future target.
A. Experimental Test Results
Fig. 9 shows the rotor, the assembled prototype and the
experimental layout of the HSEMPM assembly. It should be
noted that the total weight of the HSEMPM is 28 kg at a
power of 120 kW (accounting for the weight of the housing
and the bearing shields), therefore resulting in a power density
for the experimental prototype of ~4.3 kW/kg. However, this
indicator could be improved during batch production where
the auxiliary components and structural elements (e.g.
housing-bearing structure) can be optimized for mass
minimization [13]. Furthermore, with experience in
manufacturing AMM stator cores, the performance can be
further improved avoiding alignment deviations, profile
irregularities, material treatment, minimization of un-wanted
IV. EXPERIMENTAL STUDY OF THE HSEMPM
A full scale prototype of the HSEMPM is manufactured to
validate the modeling and simulation results, as well as to
determine the losses the overall performances. The key
parameters of the prototype are presented in Table II
(distributed winding) and schematics representation of the
structure is reported in Fig. 6.
The teeth of the stator structure are the result of the
assembly of the 6 U-shaped cores in which the stator consists
of. The AMM strip has been packed together. AMM is very
flexible in a non-annealing form and can be shaped by a
special machine. The stator magnetic core made of AMM,
created according to the in-house technology, is presented in
Fig. 8.a. The axial length of each U-core is ~5 mm. The total
axial length is achieved by means stacking together the
obtained structure. A full ring of full axial length is also
produced out of 0.08 mm FeSi laminations and used as stator
back iron, embracing the U-cores structure manufactured. The
coils have been placed in the dedicated stator slots that can be
seen in Fig. 8.a and the stator structure after impregnation is
a
b
Fig. 8. a - Stator magnetic core teeth elements made of
AMM without winding; b - Impregnated wound stator
magnetic core (left).
a
b
c
Fig. 9. a - HSEMPM rotor; b – Assembled HSEMPM; c
- Test bench and experimental layout. Here: 1-load; 2test stand; 3- HSEMPM; 4-oscilloscope.
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air gaps.
The HSEMPM installed on the test bench is shown in
Fig. 9.c. Initially, the phase voltage of the generator was
recorded during no load operation. Fig. 10 shows the trend of
voltage at idle (full speed of 60,000 rpm) in comparison with
the results obtained by means of computer simulation. It can
be noticed that the discrepancy between experimental and
computer simulation data does not exceed 8-10 V and
therefore contained within 10%. These discrepancies might
have been caused by additional air gaps that occurred during
the assembly of the stator magnetic core. Nevertheless, the
stator core stacking factor is estimated around 0.75 due to the
oscillations of the amorphous material caused by inclusions
and irregularities in the surfaces. Targeting to meet the
requirements of the target specification, one more wire turn
must be laid in the HSEMPM coils.
The next stage of the test campaign was the measurement of
the HSEMPM efficiency. For this purpose, the mechanical
power consumed by the generator at idle was estimated. In this
case, all losses in the generator (magnetic, mechanical and
additional) are estimated except for electrical losses in the
winding which, under the load condition tests, have been
computed from the resistance and the phase current.
The readings were taken from a torque sensor mounted on
the shaft of the measuring stand. To ensure the accuracy of the
measurements, the readings were taken at three different
times, respectively after 10, 40 and 60 minutes of operation.
From the test result it was found that the losses in the stator
magnetic core do not exceed 10 W at 400 Hz. At a frequency
of 1000 Hz and flux density of 1.25 T, the specific losses
measured experimentally were 8 W/kg. The total losses in the
iron of the stator magnetic core are 90 W at nominal rotational
speed. These values are 5-7 times less than those for
conventional electrical silicon steel 2421 [30]. According to
the results of experimental studies, the goal of minimizing
losses in the stator iron was achieved.
The next stage of the tests was to measure the HSEMPM
under load condition, supplying current up to 297 A. First, the
DC-resistance of the HSEMPM phase has been measured to
be 0.0023 Ω (+10% with respect calculation); the inductance
was measured to be 10.5 mH (+10% with respect calculation)
at a frequency of 1000 Hz.
To measure the output voltage of the generator, the lead
ends were connected to the load and to the measuring
Fig. 11.
Comparison of simulation and in the
experimental test results for the dependency of the load
voltage with respect to the load current.
instruments of the test stand: ammeter, voltmeter, and
tachometer. The generator is smoothly accelerated. The
rotation speed of the generator shaft was monitored by the
tachometer. Fig. 11 shows the dependency of the load voltage
with respect the load current. The simulation and the test
results are compared and a difference of no more than 8-10%
is shown. The temperatures of the hot-spots of the generator
were measure by thermocouples tightly pressed against the
surface of the winding end-turn and the stator iron and
monitored to remain within the limits imposed by the
materials used.
Based on the test results, the losses in the elements of the
HSEMPM were determined and are reported in Table III. The
methodology for measuring losses was as follows. Initially,
the HSEMPM was started with a dummy rotor (a rotor without
permanent magnets but whose dimensions correspond
completely to the rotor with permanent magnets) to full
rotational speed. The torque sensor measured the mechanical
power on the shaft of the EM, which in fact were a
combination of mechanical and aerodynamic losses.
Afterwards a rotor with permanent magnets was installed. In
addition, the EM was started at no load; the torque sensor
measured the mechanical power on the shaft of the EM, which
consist in mechanical, aerodynamic and stator magnetic core
losses. The difference between the measurements at no load
with the dummy rotor and the rotor with permanent magnets
are the losses in the stator magnetic core. The results look very
promising since the losses are much smaller than the one
declared in [4]-[10]. The measured losses distribution enables
to perform thermal calculations of the HSEMPM using the
Ansys IcePack software package. Moreover, it has been
possible to compare the thermal FEA results with the thermal
management measurements as described in the following
subsection.
B. Thermal Management
The HSEMPM thermal management has been designed
using air-cooling technology. Air is blown through the end
TABLE III
LOSSES IN HSEMPM
Fig. 10. Measurements of the voltage at idle (60,000
rpm).
Aerodynamic and mechanical losses, W
Losses in the stator winding, W
Losses in the stator magnetic core, W
Eddy currents losses in permanent magnets and rotor
sleeve, W
Total losses, W
Efficiency, %
1253
608.6
90
75
2026
98.3
IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS
winding of the drive end of the machine to flow in the main air
gap and exit passing through the non-drive end winding region
with a flow rate of 0.038 m3/s. As a result, it was found that
the temperature of the winding does not exceed 120 оС, the
temperature of permanent magnets is not more than 55 оС, and
the temperature of the stator magnetic core is not more than 70
о
С. The results of thermal calculations are shown in Fig. 12,
where the slice is taken in the middle of the machine stack
length. The heat is removed from the windings at the endwindings regions. The end winding structures are not shown in
the air flow speed plot of Fig. 12 to highlight the speed
distribution in the end regions.
A results comparison between the thermal measurements
and the computer simulation is given in Table IV. In the
experimental studies, only the temperatures of the stator
magnetic core and the stator winding were measured.
Table IV and Fig. 12 shows that the discrepancy between
the calculated and experimental data does not exceed 5-7% – a
good result that gives confidence on the modeling and the
prototyping of the machine. In addition, based on the results of
both thermal calculations and thermal measurements, the
prototyped HSEMPM is cooled efficiently with air, thanks to
the minimization of the rotor and the stator magnetic core
losses. Considering that analogues machines with similar
power are cooled by liquid (presented in [4]-[10]), the
HSEMPM here presented looks more promising also thanks to
the simplification of the cooling system. Furthermore, is worth
to highlight that the voltage drop of HSEMPM during testing
at full load has been recorded to be no more than 14-15 V.
Therefore, the HSEMPM features a very tough external
characteristic. For aircraft application, the stiffness of the
external characteristic is a very important indicator, since high
rigidity makes it possible to minimize the requirements for the
HSEMPM voltage stabilization system.
V. CONCLUSION
The paper presents calculation and measurement results for
a 120 kW HSEMPM featuring a magnetic core made of
AMM. It has been proved that the use of AMM in HSEMPM
makes it possible to minimize losses in the stator magnetic
core 5-7 times in comparison with traditional materials. This
result is confirmed experimentally and have a positive impact
on the overall machine efficiency and simplification of the
cooling system. Moreover, this paper describes the
Fig. 12. Distribution of air flow speed (left) and stator
temperatures (right)
TABLE IV
CALCULATED AND EXPERIMENTAL STATOR TEMPERATURES
Measured
Calculation
Discrepancy,
temperature, oC
results, oC
%
Stator winding
118
111.26
6.74
Stator magnetic core
70
75
5
multidisciplinary process of designing the HSEMPM with a
magnetic core made of AMM; the design and manufacturing
details of such electrical machines are given. The reduced core
losses enable the use of air-cooling techniques as proved both
in simulation and experimentally. This increases the
competitiveness of the proposed structure and its suitability
for aerospace applications where cold resources are precious
and limited. Furthermore, the winding type choice is discussed
and the selection of a distributed winding for the power range
and rotational speeds of the HSEMPM rotor is justified, in
contrast to the known works where it is recommended the use
of tooth-coil winding. The work pays special attention to rotor
dynamics evaluation in combination with the selection of the
winding type.
Based on the study, a full-size HSEMPM prototype rated at
120 kW with a maximum rotational speed of 60,000 rpm with
a stator core made of AMM was developed and tested. The
test results showed high efficiency and strong external
characteristic of the proposed technological solution.
The use of AMMs for stator cores in high power density
EMs is not extensively investigated and opens to new
challenges in material manufacturing, assembly and further
losses minimization exploiting dedicated stator topologies.
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View publication stats
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Flyur R. Ismagilov received the M.S. degree in
electrical engineering from Ufa Aviation Institute, Ufa,
Russia, in 1973 and defended his doctoral thesis in
Elements and devices of computer facilities and control
systems from Ufa Aviation Institute, Ufa, Russia, in
1998.
From 2000 until now, he is a head of department of
Electromechanics with the Ufa State Aviation
Technical University, Ufa, Russia. His research interest
includes the high-speed high-efficient machines for
aerospace application and other energy converters.
L. Papini is a research Fellow at the University of
Nottingham. He received his Bachelor degree (Hons.)
and Master degree (Hons.) in Electrical engineering
from the University of Pisa in 2009 and 2011,
respectively. He received a Ph.D. degree in 2018 from
the University of Nottingham where he has been a
research assistant since 2013. He was JSPS Fellow in
2018.
His main research interests are high speed, high power
density electric machines, machine control and
levitating system.
Viacheslav E. Vavilov (M’10) received the M.S.
degree in electrical engineering from Ufa State
Aviation Technical University, Ufa, Russia, in 2010
and the Ph.D. degree in electrical engineering from
Ufa State Aviation Technical University, Ufa,
Russia, in 2013.
From 2011 until now he is a Leading Researcher
with the Ufa State Aviation Technical University,
Ufa, Russia. His research interest includes the hybrid
magnetic bearings for electric machines and highspeed high-efficient machines for aerospace
application.
Denis V. Gusakov, (M’11) received the M.S.
degree in electrical engineering from Ufa State
Aviation Technical University, Ufa, Russia, in 2011
and the Ph.D. degree in electrical engineering from
Ufa State Aviation Technical University, Ufa,
Russia, in 2016.
From 2012 until now he is a Senior Researcher
with the Ufa State Aviation Technical University,
Ufa, Russia. His research interest includes the highefficient transformer-rectifier units for aerospace
application.
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