Innovations in Axial Flux Permanent Magnet Motor Thermal Management for High Power Density Applications

Abstract

For aerospace applications, power density is a major driving force in the design of electrified powertrains. At the forefront is the challenging design of electric motors with high efficiencies, torque, and power capabilities. Due to its high performance, the Axial Flux Permanent Magnet (AFPM) Motor is expected to be one of the leading technologies to meet the demands of these industries. Finding the balance between the cooling system’s effectiveness and subsequent parasitic losses is key to utilizing these performance benefits. Single stator double rotor topologies achieve the best torque density and lower stator losses, however are more challenging to cool as the stator is in the center of the motor. Single stator single rotor and double stator machines are less challenging to cool but typically have lower power density. Rotor air cooling is discussed including the effectiveness of blades, meshes, and vents which can be optimized to prevent demagnetization. Stator cooling is critical as many machines maximize current density producing a large amount of heat. The chosen strategy depends on the machine topology and can be accomplished by several strategies, including jackets, fins, channels, immersion cooling, hollow coils, and heat pipes.

Full text

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Innovations in Axial Flux Permanent Magnet
Motor Thermal Management for High Power
Density Applications
Colleen Jenkins, Student Member, IEEE, Samantha Jones-Jackson, Student Member, IEEE,
Islam Zaher, Student Member, IEEE, Giorgio Pietrini, Member, IEEE,
Romina Rodriguez, Member, IEEE, James Cotton, Senior Member, IEEE, and Ali Emadi, Fellow, IEEE
Abstract—For aerospace applications, power density is a
major driving force in the design of electrified powertrains.
At the forefront is the challenging design of electric motors
with high efficiencies, torque, and power capabilities. Due
to its high performance, the Axial Flux Permanent Magnet
(AFPM) Motor is expected to be one of the leading tech-
nologies to meet the demands of these industries. Finding
the balance between the cooling system’s effectiveness
and subsequent parasitic losses is key to utilizing these
performance benefits.
Single stator double rotor topologies achieve the best
torque density and lower stator losses, however are more
challenging to cool as the stator is in the center of the
motor. Single stator single rotor and double stator machines
are less challenging to cool but typically have lower
power density. Rotor air cooling is discussed including
the effectiveness of blades, meshes, and vents which can
be optimized to prevent demagnetization. Stator cooling
is critical as many machines maximize current density
C. Jenkins, colleenmjenkins@gmail.com, S. Jones-Jackson,
jonesjas@mcmaster.ca, I. Zaher, zaheri@mcmaster.ca G. Pietrini,
pietring@mcmaster.ca, R. Rodriguez, rominarodriguez88@gmail.com,
J. Cotton, cottonjs@mcmaster.ca, A. Emadi, emadi@mcmaster.ca,
McMaster Automotive Research Center, McMaster University,
Hamilton, Ontario, Canada,
Copyright (c) 2023 IEEE. Personal use of this material is permitted.
However, permission to use this material for any other purposes
must be obtained from the IEEE by sending a request to pubs-
permissions@ieee.org.
producing a large amount of heat. The chosen strategy
depends on the machine topology and can be accomplished
by several strategies, including jackets, fins, channels,
immersion cooling, hollow coils, and heat pipes.
Index Terms—Axial flux permanent magnet synchronous
motor, high power density, motor cooling, thermal manage-
ment.
I. INTRODUCTION
The transportation sector is undergoing intense elec-
trification to decrease greenhouse gas emissions (GHG).
There have already been significant advancements in
passenger vehicles. Electric car sales increased 41% in
2020, and the global electric car stock reached the 10
million units milestone [1]. However, improvements for
heavy- and light-duty trucks have lagged in comparison.
For example, global heavy-duty electric truck stock in
2020 was just 31,000 [1]. Recently, some of the key
players in the heavy-duty truck manufacturing industry
(e.g., Freightliner, Meritor, Kenworth, and Tesla) have
started to propose electric powertrains based on motors
with power ratings ranging from 200 kW up to 450 kW
[2]–[4]. Heavy-duty trucks consume high amounts of en-
ergy, require batteries for long range driving, and require
high torque for hauling heavy loads. Electric motors
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must have high power density, efficiency, and torque
capabilities [5]. Typically, an electric motor’s torque
target for long-haul heavy-duty trucks’ is 1000 Nm
[6]–[8]. These ambitious specifications would not be
possible without a reliable cooling system to control the
temperature of the electrical machine.
The More Electric Aircraft and All Electric Aircraft
(MEA and AEA) are other transportation fields with
challenging powertrain requirements. Aerospace appli-
cations inherently need high power to weight ratios,
thus occasionally pushing the continuous operating mode
close to the overload limit of the machine. For example,
the Evolito d1500/1x3 and d1500/2x3 have respectively
135 kW and 270 kW of continuous power, but just
140 kW and 280 kW (+3.7% compared to the continuous
mode) of peak power [9]. The automotive overload time
is typically short (20 – 30 sec), and the difference
between the nominal and the overload power ratings is
significant compared to aerospace applications. Conse-
quently, the cooling system must reach thermal steady
state at load levels much higher than the automotive
sector.
Among the most widespread technological trends in
eVTOL (electric Vertical Take-Off and Landing) ve-
hicle design is direct-drive propulsion, which removes
the bulky transmission components between motor and
propeller, thus enabling highly distributed propulsion
systems [10]. As no gearbox connects the electric motor
to the propeller, the aerospace industry’s torque require-
ments are increasingly demanding.
The automotive and aerospace industry’s primary con-
cerns are reliability and safety. The wide range of op-
erating scenarios due to the altitude creates challenging
design constraints for aerospace cooling systems. A ro-
bust thermal management system is crucial to increasing
motor performance and lifetime. Increased temperatures
can lead to permanent magnet demagnetization, material
failure, and bearing lubrication deterioration, increasing
losses in the machine [11]. Emerging high-power, high-
torque devices require innovative thermal management
solutions that can perform in numerous operating condi-
tions and under tight packaging constraints with minimal
performance impact.
One of the most promising electric motor technolo-
gies for the two transport electrification applications
mentioned above is the Axial Flux Permanent Magnet
(AFPM) Motor. An AFPM motor can achieve high
torque and high power densities [19]–[21]. AFPM mo-
tors market is expected to reach $353 million by 2026 at
a compound annual growth rate of 10.6% between 2021
and 2026 [22]. E-mobility, MEA, AEA, and eVTOL
are among the main drivers of this growth. Table I
reports some of the highest performance AFPM motors
currently available. It can be observed that most AFPM
motor manufacturers already have machines in their
catalog compatible with the requirements for buses and
heavy-duty trucks. Besides, some AFSPM companies
specifically focus on aerospace electrification, such as
Evolito, and others have special classes of motors for
this industrial sector (e.g., AFE Eagle series).
This paper will first discuss axial flux machine topolo-
gies in Section II and the primary sources of losses
in the machine in Section III. Then it will investigate
the thermal management of two main sections of a
permanent magnet axial flux motor: the magnets in
Section IV and coils in Section V. Finally, materials and
coolants related to the thermal designs will be discussed
in Section VI. Future trends and innovations will be
discussed in section VII and concluded in section VIII.
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TABLE I: Summary of the Main AFPM Motors Available on the Market [9], [12]–[18]
Manufacturer Model Peak Power
(kW)
Peak torque
(Nm)
Max Speed
(RPM) Mass (kg) Power Density
(kW/kg)
Cooling
System
AFE motors Cheetah Small 32.5 41.8 10000 3.5 9.3 Liquid
AFE motors Cheetah Medium 130 185 8000 16 8.1 Liquid
AFE motors Cheetah Large 280 520 6500 28 10 Liquid
AFE motors Cheetah Extra Large ∗430 1020 5000 47 9.1 Liquid
AFE motors Eagle Small †21 41.8 5000 3.5 6 Liquid
AFE motors Eagle Medium †77.5 185 4000 16 4.8 Liquid
AFE motors Eagle Large †180 500 3500 28 6.4 Liquid
AFE motors Eagle Extra Large †290 930 3000 47 6.2 Liquid
AFE motors Sailfish Small 32.5 41.8 10000 3.5 9.3 Liquid
AFE motors Sailfish Medium 130 185 8000 16 8.1 Liquid
AFE motors Sailfish Large 280 520 6500 28 10 Liquid
AFE motors Sailfish Extra Large 430 1020 5000 47 9.1 Liquid
AVID EVO AF125 100 220 12000 22 4.5 WEG
AVID EVO AF130 140 350 8000 30.5 4.6 WEG
AVID EVO AF140 220 600 5000 42.5 5.2 WEG
AVID EVO AF230 280 700 8000 57.5 4.9 WEG
AVID EVO AF240 ∗440 1200 5000 82 5.4 WEG
AVID EVO AF340 ∗660 1800 5000 122 5.4 WEG
Emrax 188 60 90 6000 6.8 8.8 Air/WEG
Emrax 208 75 140 6000 9.1 8.2 Air/WEG
Emrax 228 100 230 5500 12 8.3 Air/WEG
Emrax 268 230 500 4500 19.9 11.6 Air/WEG
Emrax 348 ∗300 1000 4000 39 7.7 Air/WEG
Evolito d500 †125 500 2500 20 6.25 Oil
Evolito d1500/1x3 †140 1350 2500 38.5 3.6 Oil
Evolito d1500/2x3 †280 1350 4000 38.5 7.2 Oil
MAGELEC M19Px 145 131 15000 19.5 7.4 WEG
MAGELEC M21Px 150 206 12000 23 6.5 WEG
MAGELEC M21Sx 130 225 12500 23.5 5.5 WEG
MAGELEC M21Rx 164 290 12500 23.7 6.9 WEG
MAGELEC M24Px 201 368 9000 32.5 6.2 WEG
MAGELEC M27Px 218 580 7350 43.5 5.0 WEG
MAGELEC M34Px ∗182 1114 4650 69.5 2.6 WEG
MagnaX AXF185 100 100 12000 8 12.5 Oil
MagnaX AXF225 170 250 10000 14 12.1 Oil
MagnaX AXF275 300 500 8000 26.5 11.3 Oil
MagnaX AXF350 ∗480 1000 4000 42 11.4 Oil
Phi Power Phi271 155 240 12000 20 7.8 WEG
Phi Power Phi301 160 320 9000 29 5.5 WEG
Phi Power Ph381 125 410 6000 38 3.3 WEG
Phi Power Ph382 ∗250 820 6000 68 3.7 WEG
YASA Ltd. 750R ∗200 790 3250 37 5.4 Oil
YASA Ltd. P400R 160 370 8000 28.2 5.7 Oil
* Good candidates for heavy-duty trucks, buses, etc.
†Motors specifically designed for aircraft applications.
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4
II. AX IA L FLU X MACH IN E TOPOLOGIES
The axial flux motor architecture can be configured in
multiple topologies [19], [23]–[28], as shown in Fig. 1.
Based on the number and the position of rotors and
stators, three main categories can be defined:
•The single-sided configuration consists of just one
rotor and one disk-shaped stator. It is also known
as SSSR (Single Stator Single Rotor).
•In the double-sided AFIR (Axial Flux Internal Ro-
tor) configuration, one rotor is sandwiched between
two stators, so there is a Double Stator Single Rotor
(DSSR).
•In the double-sided AFIS (Axial Flux Internal Sta-
tor) configuration, one stator is placed between two
rotors (Single Stator Double Rotor, or SSDR).
Thanks to the double active air gap, double-sided
AFPM motors exhibit a superior torque capability com-
pared to the single-sided configurations [19]–[21], [24].
The main difference between DSSR and SSDR AFPM
motors is that with the first, the rotor yoke can be
removed, while the stator yoke can be removed in the
second. In either case, the mass of the machine decreases
while the power density is enhanced.
A. Single Stator Single Rotor
This machine topology is conceptually simple and
similar to conventional radial flux machines, as the
magnetic circuit requires a yoke in both the stator and
the rotor to be closed [26]. However, the disk shape
makes the manufacturability of these yokes much more
challenging compared to radial flux motors. Lamina-
tions must be arranged in a spiral or cylindrical way.
Therefore, normal stamping or laser-cutting cannot be
used to make the slots directly in the stator’s lamination
pack. Another option is EDM, but the manufacturing
cost increases significantly. The stator can also be split
into teeth (which can be laser-cut and glued) and yoke
(which could be made by employing roll slitting fol-
lowed by mandrel wrapping of the steel sheet strips), but
the assembly complexity rises. An exciting alternative
for the stator and rotor’s iron cores is SMC (Soft
Magnetic Composites). The SMC material is designed
to make magnetic cores with unusual geometries, and
its specific eddy loss is relatively low [29]. Sadly, its
magnetic performance is inferior to high-end SiFe and
CoFe laminations, particularly regarding saturation level
and relative magnetic permeability. SMC’s mechanical
properties are poor: it is a highly brittle material, and
its tensile ultimate/yield stress is typically one order of
magnitude lower than steel due to its compressed iron-
powder nature.
Another disadvantage of the single-sided configuration
is the high axial force acting between the stator and
the rotor [24], [25], [27], thus stressing the bearings,
reducing their life, and increasing the friction loss. The
main advantage of the SSSR topology is the cooling sys-
tem. The stator yoke provides a large contact surface for
the heat sink, and the laminations’ thermal conductivity
is high along the axial direction. However, due to the
manufacturability issues described above and the lower
performance, most of the AFPM motors on the market
are not single-sided.
B. Double Stator Single Rotor
The AFIR configuration typically requires less Per-
manent Magnet (PM) material because it only has one
rotor. However, the PM mass savings is offset by more
copper since the end windings are longer than the
SSDR topology, as they must cross the motor axially
to connect the coils in the adjacent stator [19]. Longer
windings result in higher phase resistance and increased
copper loss. Moreover, the machine construction issues
described for the single-sided stator yoke also apply
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Conductors
Rotor
backiron
Magnets
Stator core
(a)
Rotor backiron
Magnets
Stator core
Conductors
(b)
Rotor backiron
Magnets
Stator core
Conductors
(c)
Fig. 1: Three main categories of axial-flux permanent magnet machines topologies [19], [23]–[28], a) Single Stator Single Rotor
b) Double Stator Single Rotor c) Single Stator Double Rotor.
to the DSSR stators. Finally, air-cooling the rotor is
challenging since the two sides are magnetically active
and face a stator that can transfer its heat through a thin
air gap.
Because of these disadvantages, the DSSR configura-
tion is not the most widespread in the market, although
some manufacturers such as MAGELEC Ltd. and Phi-
Power AG have selected this machine topology [15],
[17].
C. Single Stator Dual Rotor
As mentioned above, the SSDR configuration needs
less copper but more PM than DSSR [19]. However,
thermal management of the magnets is easier; since the
outside of the two rotors plays no magnetic role, it can be
used for rotor air-cooling by adding effectively designed
fins, as explained in the following sections. The two
external rotors are exposed; therefore, two covers are
required and must be designed to allow adequate airflow.
The SSDR configuration reduces the overall losses
by, in some cases, utilizing shorter end-windings and
removing the stator yoke [24], [25]. As the AFPM motor
is a synchronous machine, the fundamental component
of the magnetic field spins synchronously with the rotor,
thus representing a constant magnetic flux in the rotor’s
reference frame. Therefore, the rotor’s yoke loss is much
smaller than the stator’s (although not negligible, as will
be discussed later).
Due to several advantages, the SSDR topology is
probably the most studied, so there are multiple con-
figurations belonging to this category. The torus stator
represents a well-known example, which can be either
slotless or slotted [30]. In the slotless torus, the windings
are wound around an iron ring [24]. This way, however,
the windings are highly exposed to the rotor’s magnetic
field; hence the AC copper loss reaches very high levels.
Further, the air gap must be expanded to accommodate
the coils’ thickness, lowering the electromagnetic perfor-
mance. The problem is addressed by the slotted version
of the torus SSDR [24], which consists of two ”types”
based on the PM magnetic polarization: the NN-type,
where the poles of the two rotors are aligned, and the
NS-type, where the rotors are shifted by 180 electrical
deg [31]. A problematic peculiarity of every torus with
the stator yoke is the extra mass and manufacturing
issues (strip wound electrical steel or SMC are the main
options). However, the NS-type technically does not
utilize the yoke for the equivalent magnetic circuit [32].
The Yokeless and Segmented Armature (YASA) topol-
ogy is the AFIS configuration with the most market
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6
penetration. Structurally, the main difference between a
YASA machine and a torus is that a YASA is segmented,
meaning that the stator is formed by distinct bar-shaped
teeth [24], [25]. There is no stator yoke, and, in the case
of Fractional Slot Concentrated Winding (FSCW), each
of the teeth can be wrapped in a single coil and eventu-
ally welded together in the desired winding scheme. This
specific NS-type slotted stator is easily built compared to
the other AFPM machines because it uses electric steel
laminations cylindrically arranged. Furthermore, YASA
motors outperform the other AFPM motor configurations
in terms of torque and power densities, weight, and iron
loss [25].
The possibilities of the axial flux topology have not
been completely explored yet, and new interesting ma-
chine architectures are proposed in literature [33] every
day. Most of them, however, are small power prototypes
and their true industrial potential is still to be proven. As
the main focus of this paper is the thermal management
and not the electromagnetic design of the AFPMs, we
will consider only the main categories that have found
industrial application especially in the electric transporta-
tion.
III. MAIN LOSS SOURCES
Even if this type of machine shows good maximum
efficiency, the high power density achievable by these
motors makes the rotor and stator cooling systems
challenging. Due to the small volume and the machine
construction, AFPM machine thermal limits often restrict
overall performance. A trade-off between structural and
thermal properties typically drives the material selection
to sub-optimal cooling solutions.
This section examines the three dominant electromag-
netic loss sources of the AFPM motors, with special
regard to the YASA topology: permanent magnets’ eddy
loss, winding’s copper loss, and iron core loss [21]. As
previously explained, the stator iron core of the YASA
topology is substantially reduced compared to equivalent
radial flux machines, so the iron core loss is usually less
concerning.
A. The permanent magnets
The AFPMs are synchronous machines, so the funda-
mental term of the rotating magnetic field spins at the
same speed as the rotor. This property minimizes the
rotor losses compared to asynchronous motors, as the
magnetic flux density Binto the rotor is almost time-
constant.
However, the magnetic field at the air-gap is not a
perfectly sinusoidal spectrum and contains spatial har-
monics (such as those related to the stator slotting effect
[34]) and sometimes sub-harmonics. These spectral com-
ponents are not synchronous to the rotor spinning and
therefore excite non-negligible losses in the rotor. YASA
motors often adopt FSCW arrangements, characterized
by sub-harmonics with significant amplitude [35].
Both hysteresis and eddy losses are a function of the
amplitude and frequency of B. While the frequency of
sub-harmonics is lower than the fundamental, in AFPM
motors, it should not be neglected. Thanks to their disk
shape, YASA machines are suitable for rotor configura-
tions with a high number of poles, thus maximizing the
torque and power performance [36]. For instance, all the
Emrax AFPM motors examined here have 20 poles [14].
On the other hand, the high number of poles increases
all the electric frequencies of the motor, sub-harmonics
included.
Consequently, the rotor losses, with special regard to
the PM eddy losses, reach critical levels, and careful
attention must be paid to the rotor cooling. Furthermore,
in AFPM motors, the permanent magnets are completely
exposed to every air gap flux line. The relatively low
resistivity of the sintered NdFeB magnets results in
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TABLE II: Comparison of the AFPM Topologies
Topology Torque & power
density
Permanent magnet
mass
Manufacturing
complexity
Electromagnetic
losses Cooling complexity
SSSR - - + + -
DSSR + + ++ ++ +
SSDR (YASA) ++ ++ - - ++
high eddy currents unless proper countermeasures (e.g.,
segmentation) are taken [37]–[40]. However, segmen-
tation entails increased manufacturing complexity, thus
raising the production cost. In AFPMs, the rotors are
always attached to some structural machine part which
provides torque transmission to the shaft and mechanical
support in both axial and radial direction to the magnets.
Therefore, the magnets, together with the rotors, need to
fit into these rotor carriers. The higher the number of
edges and faces, the tighter the mechanical tolerances.
Segmentation increases not only the machining but also
the assembly complexity. The segments of each magnetic
pole should be packed as tight as possible, but this way
the magnetic forces are maximum.
B. Copper loss
The uneven current distribution in the electric ma-
chines’ windings is mainly due to the time-varying
magnetic flux leakage in the slots [8], [41]. The skin
and the proximity effects caused by the slot flux leakage
reduce the active cross-section of the conductors in the
slots, thus increasing the AC resistance and copper loss
[42].
Due to the large effective air gap, the slot flux leakage
in the SPMs can easily account for 50% or more of the
machine’s overall inductance [43]–[48]. Consequently,
the AFPM motors are typically affected by substantial
AC copper loss. Furthermore, the electric frequency of
the AFPM motors can reach significant values since the
number of poles is preferably high [14], [19]–[21], [36].
Many top-performing electric machines utilize copper
bar windings to maximize the slot’s fill factor. High
current densities are also used to improve torque and
power density but compromise the windings’ copper loss
DC and AC copper losses. Litz wire is another possible
solution to the extra AC copper loss [49], [50], although
due to the high cost, it is not used normally in large scale
production. Litz wire can also decrease the fill factor of
the stator slots.
The copper loss represents the dominant part of the
electromagnetic losses in AFPM motors [21]. Therefore,
highly effective cooling systems are necessary for the
windings to prevent failures of the electric insulation
layer, stator bar lamination adhesive, and bearing lubri-
cation breakdown due to high temperatures [51].
C. Iron core loss
The iron core losses in the electric machines are due
to the BH curve hysteresis and the eddy currents. Both
these two effects are function of the electric frequency
of the motor, and as discussed in the sections above,
the fundamental frequency of the AFPMs is not low
typically, because of the high polarity [14], [19]–[21],
[36]. The most common approach to limit the iron loss
without sacrificing the iron core magnetic performance is
by utilizing higher grade magnetic laminations, thinner
and with a higher amount of silicon or just switching to
cobalt steel CoFe and SMC. The AFPM designer can use
higher grade material as well, but choosing the YASA
topology can also represent a game changer.
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By getting rid of the stator yoke, not only the mass
of the machine is reduced, but also all the iron losses
associated with that part are canceled. This concept
applies even more to the DSSR, since there are two stator
yokes.
The YASA topology enables also the utilization of
grain-oriented electrical steel (GOES) [52]–[54], which
was originally developed for the transformers, and shows
much lower iron specific losses compared to the non-
oriented lamination steel, especially in overload condi-
tions. The GOES magnetic properties are anisotropic,
therefore it cannot be used normally in rotating electric
machines, as the magnetic flux direction in the teeth is
different from that in the yoke. On the contrary, the flux
in the YASA stator bars (or teeth) is mainly along the
axial direction, thus making possible the application of
this material to the YASA machines [52]–[54].
The only danger from the iron loss standpoint for
the YASA motors is related to manufacturing. Some-
times machining is used to achieve particularly tight
mechanical tolerances or to make the wedge shape of the
stator bars. This non-ideal practice can compromise the
insulation between the laminations of the stack in areas
where the magnetic flux density is high, thus increasing
the eddy loss. However, most of the manufacturers just
relies on laser-cutting or stamping of the laminations
without post-machining for this type of motors.
IV. PERMANENT MAGNET THERMAL
CONSIDERATIONS
One critical machine failure mode is the demagnetiza-
tion of the permanent magnets resulting in torque loss.
The primary cause of demagnetization is overheating.
Permanent magnets undergo reversible demagnetization
causing a reduction of their magnetic flux density as their
temperatures rise. For Neodymium magnets (NdFeB),
commonly used in electric machines for the automo-
tive industry, magnetic flux and magnetic torque reduce
linearly as their temperature increases due to reversible
demagnetization, as shown in Fig. 2. The reduction
becomes more significant past 120 ◦C and reaches an
irreversible demagnetization limit, and complete ma-
chine power loss is typically around 150 ◦C [55]. As an
alternative to NdFeB, Samarium Cobalt (SmCo) magnets
can be selected. Thanks to their higher Curie’s tempera-
ture, SmCo magnets exhibit better magnetic performance
at high temperature. However, SmCo typically has a
remanence flux density Brm and a maximum energy
product BHmax 25%-30% lower than NdFeB [56], so
the SmCo performance at regular temperatures is worse
compared to NdFeB, whereas SmCo price is usually
higher than NdFeB. Due to this low benefit/cost ratio,
SmCo is still a niche market material, while NdFeB
represents the most prevalent permanent magnet option
available in industry.
Fig. 2: Effect of magnets temperature on per unit residual flux
density and per unit magnetic torque, from [55].
Overheating the permanent magnets in AFPM ma-
chines becomes a severe risk in high power density
applications. Magnet eddy-current losses increase sig-
nificantly due to higher excitation flux generation in
the stator for higher magnetic torque within the often-
limited cooling capacity of the rotor. Other thermal loss
sources that contributes to the magnets overheating are
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9
the bearing mechanical losses, the rotor iron losses, and
the windage heating losses. However, these losses can
be significantly reduced by using low friction bearings,
non electromagnetic rotor components. The impact of
the windage heating losses on the magnets temperature
in AFPM machines is usually minimal due to their low
speed operation (<10,000 RPM). A Representation of
the losses and the cooling mechanisms affecting the
permanent magnets in an AFPM machine is shown in
Fig.3
Fig. 3: Main losses and the cooling air paths for fully vented
AFPM rotor.
A. Rotor Air Cooling
The rotor cooling capacity depends heavily on the
cooling method, machine geometry, and operating speed.
One advantage of the disk-shaped rotor of an AFPM
machine is the inherently higher effective cooling surface
area and relative air speed due to the larger machine
diameter compared to its radial counterparts. Such an
advantage makes air-cooling a favored cooling method,
given the simplicity and lower parasitic power consump-
tion of air cooling systems compared to other types such
as two-phase and liquid cooling systems [57]. Despite
the large body of literature body available on the topic
of AFPM cooling. There is generally less emphasis on
the topic of the rotor cooling compared to the stator
cooling due to the lower losses in the former, especially
in the machines of relatively low power density. The
AFPM rotor cooling is usually discussed in the light
of the more general topic of cooling of rotating disks
due to the close resemblance to the rotor carriers. Rotor
air cooling can be challenging due to the relatively low
thermal properties (thermal conductivity, k, and specific
heat capacity, C) of air as a coolant. Moreover, the
cooling rate depends on the machine’s operating speed
and ambient conditions, causing further cooling system
limitations. The cooling capacity can be improved by
optimizing the rotor geometry to utilize the high relative
air speed and turbulent flow patterns over the rotor and
maximize the heat transfer rate over its surfaces.
1) Air Flow Over Rotating Disks: The rotor of an
AFPM motor is commonly designed as a flat disk with
surface mounted or interior -buried- permanent magnets.
The rotor disk is either enclosed between the stator and
a cover in SSSR or SSDR machines, or sandwiched
between two stators in a DSSR machine. Both topologies
have a narrow axial clearance between the rotor surface
and the stationary part of the machine, usually referred to
as the air gap. The two opposing rotating and stationary
disk surfaces in the air gap dominate the cooling system’s
effective heat transfer surface area, especially for an air-
cooled rotor. Therefore, studying the flow over rotating
disks is of great interest in understanding the associated
flow patterns over the rotor surfaces and designing a
suitable cooling system.
The flow patterns at the air gap is critical to the
rotor heat transfer, with a direct influence on the mag-
nets cooling. The air film temperature and the average
convective heat transfer over the magnets surfaces are
greatly affected by the type of the air gap flow and the
net throughflow. The flow over the disk surfaces of the
rotor can be classified into two main types according
to the solutions originally introduced by Batchelor [58]
and Stewartson [59] as shown in 4. In a Batchelor flow
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type, two boundary layers are formed on each disk
surface, separated by a non-viscous rotating fluid core.
The boundary layers at the stator and rotor can be either
turbulent or laminar depending on the rotor rotational
speed and the resulting Reynolds number. Contrarily,
the Stewartson flow type has no core rotation, and the
tangential flow velocity across the air gap varies from
zero at the stator boundary to a nonzero value near
the rotor surface. Despite the contrariety of the two
solutions, it was concluded through many studies that the
Batchelor flow is mostly associated with enclosed rotor-
stator assemblies [60] [61]. In contrast, the Stewartson
flow type can be found in open periphery throughflow
systems. Nevertheless, Batchelor flow type can also exist
in throughflow systems at a low Reynolds flow.
Fig. 4: The representation of the streamlines and the mean ra-
dial velocity profiles associated with Batchelor and Stewartson
flow types in a typical disk rotor-stator system.
The studies in literature showed that the air gap flow
type is mostly a function of the rotor geometry and
rotational speed. The air gap flow patterns and heat
transfer in the reviwed studies is monitored and governed
using the following parameters: Rotational Reynolds
Number, Reθ, and non-dimensional mass flow rate, Cω,
the Nusselt number, Nu, the air gap ratio (G=s=R),
the flow entrainment coefficient, K(the ratio between
the flow tangential velocity at the rotating core and the
disk surface at the corresponding location),and the local
flow rate coefficient, Cqr, and Rossby number, Ro. The
discussed variables can be expressed as follows:
Reθ=ωR2
ν(1)
Cω=˙m
µR (2)
G=s/R (3)
Nu =hR
k(4)
Cqr=QtRe1/5
θ,r
2πr3ω(5)
Ro =Qt
2πR2eω (6)
where ωis the rotor rotational speed in rad/s, Ris the
outer rotor radius, νis the kinematic viscosity, ˙mis
the mass flow rate, sis the air gap thickness, his the
convective heat transfer coefficient, and kis the thermal
conductivity of air, Qtis the volumetric flow rate, Reθ,r
is the local rotational Reynolds number, and ris the
radius.
Daily and Nece [62] studied the flow in an enclosed
rotor-stator assembly and classified the flow in the air
gap into four different flows. They studied the different
flow types by varying the rotational Reynolds number
and the air gap ratio. Experiments were done using
different air gap ratios at rotational Reynolds numbers
ranging from 1E+3 to 1E+5. The authors proposed four
possible flow regimes and related their existence to
different combinations of the air gap thickness and the
flow Reynolds number. The mentioned flow regimes can
be classified into turbulent and laminar flows of merged
or non-merged boundary layers. According to Fig. 5, the
following can be observed:
1) Laminar flow with merged boundary layers exists
at low Reynolds numbers and small gap thick-
nesses.
2) Laminar flow with non-merged boundary layers
exists at larger air gap thicknesses. While for
a given air gap thickness, the laminar boundary
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layers can become unmerged at higher Reynolds
numbers. For small gap thicknesses, the boundary
layers will stay merged even at high Reynolds
numbers before the transition into turbulent bound-
ary layers.
3) Merged turbulent boundary layers exist at high
Reynolds numbers and small gap thicknesses. At
larger gap thicknesses, turbulent boundary layers
may never become merged.
4) Non-merged turbulent boundary layers can exist at
any gap thickness if the Reynolds number is high
enough.
Fig. 5: Four air gap flow regimes as defined by Daily and Nece
[62] based on Gand Reθ.
It was also observed that a minimal radial outflow
exists midplane across the air gap and that the average
tangential velocity through the gap decreases as the
gap thickness increases at a constant Reynolds number.
Poncet et al. [63] investigated the turbulent flow in a
rotor-stator system with and without a throughflow. The
flow was studied experimentally and numerically at three
different Reynolds number values and two different gap
aspect ratios. The study concluded that the flow remains
of Batchelor type for closed systems and when a weak
throughflow is allowed. The study also showed that
increasing Reynolds number at a given throughflow rate
changes from Stewartson to Batchelor type. The results
showed that both flow types can co-exist at different disk
radial locations and that the Stewartson flow tends to
dominate the flow at the smaller radii.
The transition from Batchelor to Stewartson flow
primarily depended on the throughflow rate. Two cor-
relations that relate the flow entrainment coefficient, K,
and the local flow rate coefficient, Cqr, were derived
from the collected experimental data shown in Fig. 6.
Note that high K values denote the Batchelor flow type.
In a further study, Poncet et al. [64] studied the flow
transition from the Batchelor type to Stewartson type
under the influence of superimposed throughflow. The
results showed that the direction of the throughflow
determines the flow type. Superimposed centripetal flow
maintained Batchelor flow type while it had faster core
rotation than in case of closed system Batchelor flow.
The study also characterized the transition from Batch-
elor to Stewartson flow type using the Rossby number
considering the radial position in the gap.
Fig. 6: Flow type dependency on local flow coefficient, Cqr
[63].
2) Heat Transfer in Rotating Disks: The different
flow types discussed earlier are determinants of heat
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transfer rate over the rotor surfaces. Many studies in
the literature estimated and compared the heat transfer
associated with the different flow patterns over rotating
disks under different conditions. Cobb and Saunders [65]
and Dorfman [66] studied the heat transfer over a heated
free disk- without a stator wall- and derived correlations
to estimate the Nusselt number for laminar and turbulent
flows. The correlations provided by Cobb and Saunders,
and Dorfman are only applicable for a heated free
disk. Accordingly, these correlations are insufficient to
represent the heat transfer in an AFPM machine rotor,
given the small clearances around its surfaces, especially
at the air gap. However, the flow between the rotor and
the shroud -or cover- can be of the free disk type if the air
gap ratio is large enough. Owen and Rogers [67] defined
a limit for gap ratio over which the stator influence on
the air gap flow can be neglected, and the rotor can be
assumed as a free disk. This limit is correlated with
the rotational Reynolds number and can be calculated
through the following equation:
Glim = 1.05Re0.2
θ(7)
For any smaller gap ratio, the heat transfer at the air gap
has to be analyzed based on the type of flow exhibited
at the gap. In the case of fully enclosed shrouded
rotor-stator systems, Owen and Rogers [67] analytically
derived four mean Nusselt number correlations corre-
sponding to each of the four flow regimes discussed by
Daily and Nece [62].
Their results showed that for low air gap ratios (G <
0.01) where the flow corresponds to either regime I
(laminar) or III (turbulent) with a merged viscous bound-
ary layer, the mean Nusselt number reduced as the gap
ratio increased. At larger air gap ratios (0.01 < G <
0.06), where the flow exhibits regime II or IV, the mean
Nusselt number is an increasing function of the air gap
ratio. However, a rotating core between the two boundary
layers reduces shear stress causing the mean Nusselt
number to drop significantly compared to the smaller
air gap ratio nearing that of the free disk.
Soo [68] analyzed laminar flow in a fully enclosed
rotor-stator system with a superimposed inward / out-
ward flow through an opening at the stator center. The
study concluded that enforcing a net outward flow in
the air gap is more beneficial for the rotor heat transfer
compared to enforcing an inward flow. Poncet et al. [64]
experimental study showed similar results and observed
that the superimposed outward flow transitioned the flow
from Batchelor type to Stewartson type. This transition
results in better heat transfer over the rotor surface
due to increased shear and lower average air gap bulk
temperature.
Kapinos [69] and Owen [70] investigated the flow in
an unshrouded rotor-stator system with a superimposed
flow. The different flow types discussed earlier are de-
terminants of heat transfer rate over the rotor surfaces.
While this increase is dependent on the imposed mass
flow rate, air gap ratio, and rotational speed. The study
also concluded that the convective heat transfer is more
dependent on the mass flow rate and less dependent on
the rotational Reynolds number values at low Reynolds
numbers for narrow air gaps (G < 0.06).
Boutarfa and Harmand [71] conducted a similar study.
However, the rotor self-pumped the throughflow air with-
out superimposing additional airflow from an external
source. The results showed that the throughflow allowed
through the stator opening increased the Nusselt number
values on the rotor surface regardless of the Reθand
Gvalues compared to closed systems. The study was
carried out for a range of rotational Reynolds numbers
5.87E+4 < Reθ<1.4E+6, at G= 0.01, 0.02, 0.06,
and 0.17, as shown in Fig. 7. The comparison of flow
in conjunction with the heat transfer at different gap
ratios showed that for smaller air gap ratios (G≤0.01),
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the Nusselt number over the rotor across the air gap
is nearly constant. This agrees with Owen and Rogers,
with a higher mean value than larger gap ratios at the
same Reθ. The author related the reduced mean Nusselt
number in a higher gap ratio to the fact that they result
in lower air flow rate and Batchelor flow pattern across
the air gap. They explained that for G= 0.01, the air
gap flow is dominated by the viscous forces resulting
in a Couette flow type filling the entire gap, similar to
regimes I and III discussed by Daily and Nece. The
result is a higher heat transfer rate at the inner radii,
with relatively small boundary layer thickness. As the
flow progresses to the gap periphery, the boundary layer
thickness increases, and viscous forces become more
significant, deteriorating the heat transfer. Moreover, the
air temperature increases as it flows radially, reaching
temperatures close to the rotor surface, which further
deteriorates the heat transfer. Accordingly, the local heat
transfer coefficient decreases proportionally with the
radius. The results also showed that the heat transfer
is enhanced significantly when the flow reaches the
transition region.
Fig. 7: Mean Nusselt number as a function of Reθ, and G[71].
Rasekh et al. [72] provided a comprehensive numeri-
cal study for the rotor heat transfer in AFPM machines.
The study was carried out for an unshrouded rotor-stator
assembly with axial openings in the rotor for throughflow
ventilation. The rotor had surface mounted magnet poles
with annular air channels in between each pair, acting
as a centrifugal fan and enhancing the system airflow,
hence the heat transfer. The study varied multiple pa-
rameters, and correlations for the mean Nusselt number
were derived accordingly. The correlation constants are
optimized so that their results are not dependent on the
system dimensions and ambient temperature. The results
showed that for values of G < 0.01, the heat transfer
over the magnets on the rotor is directly proportional
to the air gap ratio, in contrast to the results discussed
by [67] and [71] for air gap ratios larger than G=
0.01. The difference in results can be related to the
much smaller air gap ratios tested in Rasekh’s study.
However, two other major differences that were not
considered might have influenced the results—first, the
key difference between the studied rotor geometries.
Second, the existence of annular air channels affects the
flow field at the air gap as they become the primary
air flow passage in the system, according to the results
presented by Howey et al. [73] and Airoldi et al. [74].
Howey et al. [75] reviewed previous air gap convec-
tion studies in axial and radial flux electrical machines.
The authors concluded that for small-sized machines to
reach high power densities, introducing a superimposed
flow at the air gap and increasing the roughness of
the internal flow surfaces is recommended to increase
the air gap heat transfer. This will maintain better flow
turbulence and higher temperature gradients over the
effective heat transfer surface area, hence higher cool-
ing rates. However, the study by Nece and Daily [76]
showed that surface roughness does not affect on the
frictional resistance on the rotating disk when the flow
is laminar. While for turbulent flows, increasing the disk
surface roughness significantly increased the windage
torque -the fluid frictional resistance torque- reaching
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14
nearly double the value of the smooth disk at high
rotational speeds, given the direct relationship between
the frictional moment and the average heat transfer over
rotating disks [77]. It can be concluded that increasing
the surface roughness of a disk operating at laminar flow
speeds will not yield better heat transfer. In contrast, a
significant enhancement in heat transfer can be yielded
with turbulent flows but at the cost of increased frictional
resistance. Therefore, this approach should be carefully
studied concerning the machine operation speeds.
The discussed studies emphasize maintaining a small
gap ratio (G≤0.01) and high rotational Reynolds
number for better cooling of the magnets through the
air gap in an AFPM machine. However, relying solely
on the air gap to cool the magnets might not be enough.
Hence, indirect cooling of the magnets through the rotor
is usually necessary, especially for machines of higher
power densities. Several studies have been performed to
test the effect of different geometrical alterations on the
improvement of rotor cooling.
B. Approaches for Rotor Air Cooling Enhancement
The low cooling rate of rotor air systems can be
maximized under the same machine operating conditions
by optimizing the rotor geometry to enhance turbulence,
air mass flow rate, and therefore overall heat transfer.
Many researchers have discussed this subject, and many
approaches were proposed in the literature to improve
rotor air cooling by enhancing its geometry, especially
for SSDR machines. The following sections discuss
the main approaches to improving rotor cooling in the
literature.
1) Ventilated Rotors: Rasekh et al. [78] investigated
the air flow in discoidal rotor-stator systems at a range
of rotational Reynolds numbers 2.5 E+4 ≤Reθ≤
2.5 E+5, and air gap ratios G=0.0067, 0.0133, and
0.02667. The study results showed that vents/holes in
the rotor body enable a net radial outflow through the
air gap. Hence, increasing the air gap heat transfer effect.
However, the study did not quantify or measure the
claimed enhancement compared to a non-vented rotor.
Airoldi et al. [74] also argued the benefit of having air
admitted through the rotating rotor boss (hub) on the
air gap heat transfer. The study focused on the heat
transfer over only the stator side of the air gap, and the
enhancement due to ventilation was also not quantified.
Chong et al. [79] discussed the effect of rotor axial and
radial holes in an air cooled AFPM machine. The study
showed a significant increase in the throughflow mass
flow rate by introducing the holes, which enhanced the
stator cooling result. Similar to Airoldi et al. study, the
effect of the holes on rotor cooling was not assessed in
this study.
Fig. 8: Representation of the different rotor carrier geometries
variations based on cooling features discussed in literature.
2) Surface-Mounted Protruding Magnets: The effect
of surface mounted protruding magnets on air cooling
in AFPM machines was investigated by Airoldi et al.
[74]. The thickness of the surface mounted magnets,
and the rotational speed was varied in a validated CFD
(Computational fluid dynamics) model. The study found
that increasing the height of the channels in between
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15
the magnets increased the pumping capacity of the rotor
and the resulting windage losses. This effect resulted
in enhanced heat transfer on the stator side of the air
gap. Howey et al. [73] conducted a similar experimental
study with two rotor configurations, one with a flat rotor
surface, as shown in Fig. 8 ( a,e) another with surface
protruding magnets, as shown in Fig. 8 ( c,g). The study
concluded that the average heat transfer over the stator
side was increased by 20%-30% at similar operating
conditions when surface protruding magnets were used.
Rasekh et al. [72] studied rotor protrusions in AFPM
numerically. Unlike the geometries presented in Airoldi
et al. and Howey et al., the machine was throughflow
ventilated through annular openings (channels) in the
rotor. The geometry of the magnets, gap ratio, and
rotational speed was varied in the study. The effect of
each parameter on the heat transfer in the stator and the
rotor was studied. The results showed that the overall
heat transfer in the machine could improve substantially,
especially on the magnets, by having both the rotor pro-
trusions and the rotor openings simultaneously featured
in the design. It can be argued that rotor protrusions are
only effective in ventilated machines; hence, the author
emphasized on the equal importance of both features in
the design for enhanced cooling.
3) Rotor Embedded Radial Fan: In the efforts to
maximize the rotor cooling in AFPM machines, Van-
sompel [80] suggested the addition of fan blades to
the back side of the rotor to enhance the rotor heat
transfer and structural rigidity. A well known concept
in the automotive industry for brake rotors cooling but
can only be applied to AFIS topologies. Fawzal et al.
[81], [82] studied the implementation of this concept.
Three fan geometries (backward-curved, radial/straight,
and tear drop pillar blade) from other applications were
investigated. Each geometry was modeled separately
as an embedded fan section in an AFPM rotor. The
performance of each fan design was judged based on
the cooling capacity, pumping capacity, pressure drop,
and windage losses. The study was done numerically
using a validated CFD model of the machine. Although
the study showed a significant enhancement in the rotor
cooling performance with the addition of the rotor fan
blades, this enhancement was not quantified. The results
showed that the backward inclined had the best thermal
performance among other geometries since it provides
high cooling capacity at low additional windage losses.
On the other hand, the radial blade offered slightly
better cooling capacity but at the high cost of added
windage losses. The authors introduced an index, the
Rotor Cooling Performance Index (RCPI), to compare
designs based on their convective cooling capacity and
windage losses. Despite helping rank the designs based
on two of the most crucial rotor design parameters,
the index is calculated at equal temperatures of the
rotor across all designs, hence, varying thermal load.
Accordingly, the index comparison might not be directly
applicable when comparing designs at constant thermal
load.
In a further study, Fawzal et al. [11] studied the
ventilation of the bladed rotor air cooling system and its
effect on the machine’s thermal performance. Three rotor
shroud designs were studied with different variations of
the location and shape of inlet and outlet ports. The
first design had a radial inlet and radial outlet. The
second design had a tangential inlet and tangential outlet.
Finally, the third design had an axial inlet and tangential
outlet. The three designs were modeled and simulated
using CFD, where all simulations prescribed an equal
mass flow rate at the inlet. The axial inlet-tangential
outlet design resulted in the lowest pressure losses and
significantly higher average heat transfer coefficient over
the rotor at a comparable increase in windage loss values
compared to the other designs resulting in an overall
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16
enhancement of the rotor thermal performance.
4) Rotor Liquid Cooling: Other motorized compo-
nents in an electric vehicle can be integrated in to
the AFPM machine rotor to improve its cooling per-
formance and power density accordingly. Nishanth et
al [83] studied the integration of a hydraulic radial
piston pump with the rotor of a single stator AFPM.
The design is optimized to maximize the pump leakage
flow to improve the machine cooling capacity. The study
showed a that a power density of a 7.7 kW/kg at
12500 RPM rated speed using conventional materials.
Chuan et al. [84] proposed utilizing hollow shaft liquid
cooling in conjunction with forced in AFPM rotor. the
study showed a significant improvement in the magnets
temperature through utilizing the hybrid cooling method
over the air cooling method. The study compared the
different cooling configurations using Lumped Parameter
Thermal Network (LPTN) model. However, the hybrid
cooling methods has to be further validated through
detailed numerical and experimental studies.
Based on the discussed cooling features presented
in the literature, different combinations of the rotor
cooling features can be used together to achieve higher
air mass flow and cooling rate. This can be achieved
at the same operating condition without significantly
changing the machine design or dimensions. Zaher et
al. [85] numerically studied and compared the thermal
performance of six different rotor geometries, shown in
Fig. 8. considering different combinations of rotor cool-
ing features (rotor vents, protruding magnets, and rotor
fan blades). The different designs were ranked based
on their cooling efficiency considering the maximum
magnets temperature and the associated windage losses.
The study concluded that having rotor vents that allow
throughflow air is crucial for designing an efficient rotor
air cooling system. Moreover, employing the rotor vents
and the protruding magnets in the same rotor resulted
in a significant reduction in the magnets temperature at
a marginal cost of added windage losses. Adding fan
blades to the rotor resulted in significantly higher mass
flow rate, turbulence, and heat transfer over the rotor
surfaces. However, the added windage loss (or the power
consumed by the fan) might not be justifiable unless
a sufficient cooling rate could not be met using only
direct cooling methods. This emphasizes the importance
of maximizing the direct cooling of the magnets through
either the air gap, or air channels in between surface
mounted protruding magnets for efficient cooling of the
rotor. Despite the drawbacks of using embedded fan
blades for cooling the rotor, such systems can be opti-
mized to maximize its cooling effect while maintaining
the fan power moderately low. This can be done by
optimizing the fan blade geometry using turbomachinery
concepts as discussed in [86].
V. STATOR WINDING COOLING STRATEGIES
Stator cooling methods range widely from natural
convection to two-phase liquid cooling, including forced
air, outer liquid cooling jackets, flow channels, direct
cooling, and heat pipes. The design of these cooling
methods can significantly impact the machine’s overall
thermal, electrical, and mechanical performance. Since
increased temperatures can reduce the machine efficiency
and cause adverse effects on motor stresses, the thermal
management system must be selected such that it can
effectively remove excess heat. Each cooling method
comes with its benefits and drawbacks. Generally, natural
and forced air-cooling for high power density motors is
insufficient for stators in axial-flux motors. The follow-
ing sections will focus on single-phase and two-phase
liquid cooling. Liquid cooling in literature is summarized
in Table III, which displays the type of cooling method,
the machine configuration, the coolant used, the machine
power output, and the machine losses. When the machine
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17
losses are not readily available, the stator losses for the
machine are substituted and marked with an asterisk.
A. Jackets
Stator cooling jackets are liquid cooling solutions that
indirectly cool the copper coils since they are not in
direct contact with the conductors. Generally, the coolant
is in contact with the stator encapsulation or motor
housing. A cooling jacket is either the sole stator cooling
method or the manifold connection between different
cooling channels. An outer jacket that contributes to the
sole cooling function of the stator is one of the easiest
and most common liquid cooling methods. While a wide
range of design parameters can be altered depending on
the design constraints, there are two common groups, the
outer diameter (OD) cooling jackets, and the end cover
jackets. Typically water, ethylene glycol (EG), and water
ethylene glycol (WEG) are used in these jackets (further
discussed in Section VI-A).
Most commonly, stator cooling jackets for axial flux
motors are located at the outer diameter (OD) of the
motor, shown in Fig. 9, or at the end covers, as shown
in Fig. 10. The size and cooling surface area of the
jacket have a large effect on the heat transfer and, in
turn, the stator temperature. Most commonly, the OD
cooling jackets are used in the single-stator and double-
rotor configurations, while end cover jackets are used in
double-stator and single-rotor configurations. The goal
is to ensure the most heat can be removed from the
stator, so the cooling location changes with the changing
configuration.
Veg and Laksar [89] studied the impact of these two
cooling systems with their respective configurations. In
this study, the end cover cooling jacket was only required
to cool a single stator (i.e., two end cover cooling
jackets for the motor). It was found that the double-
stator end cover cooling system had almost 30% lower
total temperature than the single-stator OD jacket. This
lower temperature is due to each system required to
cool half of the losses the single-stator jacket dissipates.
Additionally, the end cover jacket allowed for greater
temperature uniformity. However, this double-stator mo-
tor configuration requires more space than the single-
stator.
Furthermore, the greater surface area of the double-
stator configuration allows for greater potential power
output. This increased power is due to the greater
cooling performance, which allows for greater power
at the same temperature limitations. Although double-
stator configurations are generally larger than single-
stator configurations, the increased power output means
similar power densities.
1) Outer Diameter Jackets: Outer diameter cooling
jackets are limited in design due to the relatively low
space availability. Commonly, these jackets are con-
structed from aluminum, but their flow path design range
from spiral jackets [91], [111], axial flow paths [91],
[111], parallel flow channels [87], [98], or a single
channel [88]–[90], as shown in Fig 9.
Le et al. [91] compared two OD cooling jackets,
one spiral channel, to an axial parallel flow channel
jacket. They found that the spiral jacket impacted the
rotor thermal performance and reduced the maximum
rotor temperature by 3.4 K. However, they determined
that the overall motor temperature still needed to be
reduced, so they studied the effect of adding fins to the
inner diameter of the cooling system. The fin results are
presented in the following section (Section V-B).
Unlike the conventional OD cooling jacket design,
Zhang et al. [98] swapped the single-channel jacket for
five copper pipes in the stator. This water-cooling system
is potted in epoxy to reduce the thermal resistance
between the heat losses (generated in the stator seg-
ments and the windings) and the water-cooling system.
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18
TABLE III: Stator Cooling in Literature for Axial-Flux Permanent Magnet Machines
Author/Year Cooling Method Machine
Configuration Coolant Peak Power
(kW)
Losses
(kW)
Chuan et al. 2019 [84] OD jacket SSDR Water 1.5†0.177
Chang et al. 2021 [87] OD jacket SSDR Water 50 –
Zhang et al. 2014 [88] Single cavity/ OD jacket SSDR Water 65 4
Veg and Laksar 2019 [89] Single cavity/ OD jacket SSDR Liquid 25†1.67
Marcolini et al. 2019 [90] Single cavity/ OD jacket SSDR Water 25†3.2∗
Le et al. 2020 [91] OD spiral jacket SSDR Water 60†3.14
Lamp´
erth et al. 2015 [92] Single cavity/ end cover jacket DSSR Water 79 –
Veg and Laksar 2019 [89] Single cavity/ end cover jacket DSSR Liquid 25†1.67
Odv´
aˇ
rka et al. 2010 [93] Inner core channel with spoilers SSDR Water – 1∗
Rahman et al. 2006 [94] Single cavity/ OD jacket with fins SSDR Liquid 50 5.7
Qi et al. 2019 [95] Spiral end cover channels DSSR Water 56 3.2
Liu et al. 2019 [96] Cavity, spiral, tandem,
and Z-shaped end cover channels DSSR Water 50 –
Lai et al. 2021 [97] Z-shaped end cover jacket DSSR Water 100 3.03
Zhang et al. 2016 [98] Copper pipe OD channels SSDR Water 70 8
Chang et al. 2021 [87] OD jacket with winding channels SSDR WEG 65 –
Gerlando et al. 2020 [99] Full-winding channels SSDR WEG 68†6.8
Jones-Jackson et al. 2021 [100] Full inter-winding / partial inter-winding SSDR WEG – 4.4
Li et al. 2019 [101] Serpentine inter-winding copper pipe DSSR Water 130†8.44
Mohamed et al. 2021 [102] OD jacket with fins SSDR Water 4 0.016∗
Mohamed et al. 2022 [103] Axial OD jacket with fins SSDR Water – 2.08
Talebi et al. 2022 [52] Microchannels on end windings SSDR WEG 250 14.37
Camilleri et al. 2012 [104] Immersion cooling SSDR Liquid 100 6
Zhang et al. 2019 [105] Immersion cooling with epoxy baffle plate SSDR Oil 53.8 2.2
Geng et al. 2020 [106] Immersion cooling with baffles SSDR Oil 50 3
Wanjiku et al. 2021 [107] Immersion cooling with baffles SSDR EG / Oil 50 15
Camilleri et al. 2016 [108] Flooded stator with baffles SSDR Oil 100 6
Liu et al. 2021 [49] Immersion with flow directing fins SSDR Oil 150 4
Lindh et al. 2016 [109] Hollow copper coils SSDR Oil 100 3.4
Lindh et al. 2017 [50] Hollow copper coils DSSR Oil 100 5.2
Le et al. 2021 [110] Zigzag channels (series and
parallel) with heat pipes SSDR Water 60 2.8
* Stator losses only
†Nominal/continuous power
While this study focused on mechanical analysis, the in-
depth thermal performance analysis was out of scope.
However, a computation fluid dynamics (CFD) analysis
was done for the whole motor, primarily considering the
airflow over the rotor. It was found that the permanent
magnet temperature rises very quickly, and the air gap
has a relatively low speed. Therefore, the rotor cooling
was the limiting factor on speed and torque for this
machine [98].
Furthermore, Tong et al. [111] studied the effect of
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19
(a) (b) (c) (d)
Fig. 9: Examples of outer diameter liquid cooling jackets for
axial-flux permanent magnet machines, a) spiral cooling jacket
[91] b) axial channels [111] c) single radial channel [88]–[90],
and d) parallel channels [87], [98].
the gap thickness between the stator core and the water
jacket. A circumferential, or spiral, jacket design was
used for a low-speed permanent magnet machine. The
different gap thicknesses are considered based on differ-
ent manufacturing tolerances. They found that the gap
between the two materials linearly affects the maximum
temperature rise. When the clearance is larger than
0.15 mm, the water-cooling system is likely no longer
effective. The maximum temperature rise of the machine
was greater than 80 K for this condition [111]. Since
air has a much lower thermal conductivity than solid
materials, epoxy is used in empty spaces to increase the
thermal performance.
2) End Cover Jacket: End cover jackets, compared
to OD jackets, have more available space, and therefore,
the flow path can range widely, as shown in Fig. 10.
The most commonly found are the single fluid cavity
[92], [96], [112], [113], helical [96], and tandem [95],
[96], [114], [115]. In addition, Li et al. [96] studied
the thermal performance of these types and a Z-shaped
channel on a 50 kW motor. Also, their single-fluid cavity
was equipped with spoilers to help direct the fluid flow
and increase mixing. For each cooling system, the inlet
and outlet diameters and thermal conductivities were
equal, neglecting the rotor impact.
Liu et al. [96] found that the Z-shaped channel had the
greatest pressure drop, due to the sharp corners, with an
(a) (b) (c) (d)
Fig. 10: Examples of end cover cooling jackets for axial flux
permanent magnet machines, a) single channel or cavity [96],
[115] b) helical channels [96], c) tandem channels [116], d)
z-shaped channels [95], [96], [114], and a single fluid cavity
[89], [92], [112], [113].
inlet pressure of about 7.1 kPa. The cavity with spoilers
was found to have the lowest pressure since the flow area
was much larger. While this resulted in an inlet pressure
of only 0.4 kPa, the fluid velocity in the cavity was much
lower. The helical and tandem channels were found to
have comparatively similar inlet pressures at about 1.8
kPa and 2.8 kPa, respectively. In terms of the thermal
performance, however, all channels, besides the cavity
with spoilers, saw similar maximum fluid temperatures,
around 81 - 82 ◦C. Due to the low fluid velocity, the
maximum temperature in the cavity was about 91 ◦C
(using coolant at 25 ◦C and 30◦C for ambient).
Therefore, since the tandem channel saw the lowest
maximum temperature, at 81 ◦C, Liu et al. studied the
geometrical effects on thermal performance. They stud-
ied the impact of the number of channels, axial channel
length, radial channel length, and inlet velocity. When
the number of channels was studied, the cooling surface
area stayed the same, but the channel width decreased
with an increasing number of channels. It was found
that the pressure drop increased while the temperature
decreased with increasing channels. To balance these
two factors, they concluded that the optimum number
of channels was 3-5. Furthermore, they found that the
axial and radial length of the channels had a significant
impact on the pressure drop but less on the temperatures.
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20
Using the tandem water-cooling channels, they selected
the following parameters: five channels, 10 mm of axial
length, 12 mm of radial length, and an inlet velocity
0f 0.8 m/s. This analysis showed a maximum motor
temperature of 95 ◦C at the end windings [96].
Qi et al. [95] and Huang et al. [114] also studied a
tandem cooling channel for their 36 kW machine, with
a maximum power of 56 kW. This motor also used five
cooling channels with a width of 10 mm. With an inlet
flow rate of 12 LPM and an ambient temperature of
60 ◦C, the maximum winding temperature is 109 ◦C.
However, after operating at maximum power for one
minute, the maximum winding temperature increased to
171 ◦C. They concluded that the insulation grade is rated
at 200 ◦C, so there is still a safety margin [95], [114].
Additionally, Chai et al. [112] and Bi et al. [113]
studied the impact of the thermal resistance heat path
on the machine temperatures using a single channel end
cap for each of the two stators. Extending the end cap
towards the end windings (removing the epoxy in its
place) decreased the thermal resistance from the stator
yoke and windings to the cooling system. Two principal
factors were studied to determine the effects on the
thermal performance, the epoxy thermal conductivity and
the stator yoke thickness. The increase in epoxy thermal
conductivity improved the average winding temperatures
of both end cap designs. The extended end cap saw
almost a 6 ◦C temperature reduction at the lowest epoxy
thermal conductivity (0.4 W/m-K). However, this tem-
perature difference was reduced with increasing epoxy
thermal conductivity. A more considerable difference
was seen with the changing stator yoke thickness. While
the temperatures increased in both designs with thicker
stator yokes, the extended end cap saw a 10 ◦C decrease
in average winding temperature at the thickest yoke (25
mm) [113].
Furthermore, Bi et al. [113] studied both designs at
an overload condition (700 Nm at 100 RPM) for 180
s. They concluded that the extended end cap design
could keep the maximum machine temperature below the
limit (130 ◦C), while 5.3 ◦C lower than the conventional
design. The result was in an additional 23 s, past the
requirement for the modified end cap [113].
B. Heat Transfer Fins
The cooling performance of liquid jackets can also
be improved with extended protrusions or fins. The
addition can improve the thermal resistance from the
heat-generating components by increasing the contact
surface area of the cooled jacket material. Le et al. [91]
used a spiral OD water jacket combined with twelve
rectangular fins protruding in towards the windings on
the inner diameter of the housing, as shown in Fig. 11a.
By varying the width and depth of these fins, an almost
15 ◦C reduction in maximum rotor temperature was
seen compared to the spiral jacket alone, without fins.
They found the rotor temperature was stable (no longer
decreasing) at a width and depth of 15 mm and 6 mm,
respectively. The resulting maximum rotor temperature
was 121 ◦C for the 60 kW machine.
Fins have been used in axial flux motors for long-term
testing. Rahman et al. [94] designed a 25 kW AFPM
with an aluminum liquid cooling ring to be placed in
the OD of the machine stator. A schematic of a cooling
ring with aluminum fins is below in Fig. 11b. This
cooling ring is equipped with fins between the winding
turns and potted with a high thermal conductivity epoxy.
The developed wheel motor and cooling system were
patented, placed in a GM S10 mule vehicle, and driven
over 1000 miles.
Vansompel et al. [117] and Mohamed et al. [103]
have published some papers on their yokeless and seg-
mented armature AFPM machine designs. Each machine
is cooled with a different coolant flow path and is
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21
adjusted whether the power converter is integrated with
the machine. However, the stator heat extracting fins
have stayed the same between these designs and are
shown in Fig. 11c. The cooling design is made from
a stack of 2 mm aluminum alloy sheets to make a stator
housing with fins. This allows for a high thermal con-
ductivity parallel to laminations and winding strands and
low in the perpendicular direction. In the non-integrated
motor, the cooling effect of the heat extracting fins is
analyzed by Vansompel et al. [117]. Here, a coolant
is not specified, but a CFD model and a constant heat
transfer coefficient (convective boundary condition of 40
W/m2-K, with a reference temperature of 25 ◦C) are
applied to the fins. While they found the fins increased
the losses from 65 W to 80 W, the motor temperature
was much cooler. Due to the low thermal conductivity
of the potting material, without the fins, there is a large
temperature gradient over the epoxy. The introduction
of these fins allowed for a much smaller temperature
gradient within the stator [117].
On the other hand, Mohamed et al. [103] studied a
similar aluminum fin structure with an integrated inverter
module and axial flux machine. While the fins were the
same design as Vansompel et al. [117], a single cooling
channel was added per module, with one inverter module
per stator coil. Several design changes were studied,
including swapping the aluminum housing for copper, a
new thermal interface material (TIM) with an improved
thermal conductivity (between the cooling and power
modules), and three channels per module instead of one.
Out of the three improvements, the TIM significantly
impacted the increased power density (at the same tem-
perature). The improved thermal conductivity increased
the normalized power density by 52%. Additionally, the
copper housing increased the winding losses from 73 W,
per module, to 83 W. This resulted in a hotspot of 115
◦C or a reduction of 3 ◦C winding hotspot temperature
compared to aluminum. This hotspot was seen with an
inlet water temperature of 25 ◦C, 1 LPM flow rate,
and 7.45 W in switching losses. Furthermore, combining
all three improvements, the normalized power density
increased by 91% [103].
Alternatively, Polikarpova et al. [118] and Pyrhonen
et al. [119] have implemented copper bars extending
from the liquid-cooled frame into the stator iron for
the double-stator AFPM machine. An example with two
copper bars shown is displayed in Fig. 11d. Most of
the winding heat was transferred through the stator iron
towards the cooling system. The copper bars can extract
the stator heat combined with the stator teeth. A total of
three bars per tooth were selected, and it was stated that
any more would cause mechanical stability problems. No
significant losses are generated since these copper bars
are parallel to the magnetic flux path. However, a slight
increase in iron losses is seen due to the increased flux
density (more current in a smaller area) but is deemed
insignificant. They found that just a single copper bar
in each stator tooth contributed to a reduction of 6 ◦C
of the maximum stator slot winding and end winding
temperatures, compared to just the jacket. It also had a 10
◦C reduction in the maximum rotor embedded magnets.
In addition to the copper bars, a potting material is
included in the stator. They see further reductions in
maximum temperatures with both thermal improvements
included compared to just the jacket. It is found to have a
21 ◦C, 21 ◦C, and 25 ◦C reduction in the slot windings,
end windings, and magnets, respectively [118].
C. Internal Flow Channels
Due to sizeable copper losses in the stator, the most
effective cooling methods reduce the thermal resistance
from the coolant to the windings. One popular cooling
enhancement includes cooling channels that come as
close to the coils as possible.
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22
(a) (b) (c)
Liquid cooled
frame
Windings
Stator iron
Copper bars
(d)
Fig. 11: Stator cooling fins a) connected to the cooling jacket at the outer diameter of the stator [71], b) at the inner diameter
of the stator spiral jacket [89] c) [98], [105] and d) copper rods extruding into the stator [106], [107].
By reducing the thermal resistance from the cooling
system to the heat-generating components, the output
power can be increased, as well as other key performance
metrics. Chang et al. [87] studied the motor performance
improvement of a conventional water-cooling jacket by
adding channels connected to the same jacket. These
channels flowed in between adjacent windings. This
analysis found that the peak power could be increased
from 50 kW to 65 kW with the jacket and the intro-
duction of the channels, respectively. It was found that
the peak torque, rated power, and rated torque could be
increased by 14.3%, 33.3%, and 9.1%, respectively [87].
Examples of different cooling channel designs are
shown in Fig. 12. The configurations include full inter-
winding channels, for indirect and direct cooling, in
Fig. 12a and Fig. 12b, respectively. Additionally, copper
has been used to make channels or pipes that meander
adjacent windings. This is done with a single copper
pipe and several, as shown in Fig. 12c, Fig. 12d, and
Fig. 12e. Also, cooling channels have been embedded in
the stator core, as shown in [93].
By including channels in the stator cooling system, the
cooling configurations expand. Most commonly, chan-
nels are connected to a manifold, similarly designed
to the OD cooling jackets, especially when the coolant
flows through each channel parallel to one another.
However, channels can have an in-series coolant flow
path, and a manifold is less common.
For parallel cooling channels connected by a manifold,
most commonly, the channels wrap around the entire
outer diameter of the coils. These are described as full
inter-winding designs since they cover the full surface
between the neighboring windings. However, there are
also partial inter-winding designs. By reducing the space
of the cooling system, the machine design can become
more compact [100]. Care must be made since reducing
the cooling capacity can reduce the motor performance.
To further decrease the thermal resistance between the
coils and the coolant, these channels can be cooled
with dielectric coolants, such as oil. Dielectric coolant
can allow the liquid to contact the copper windings
directly. However, water-cooling channels must have
an insulating buffer between the fluid and windings.
Although water has a higher thermal conductivity than
oil, this insulating layer can seriously reduce the cooling
capacity [100].
Jones-Jackson et al. [100] studied the stator perfor-
mance between a full inter-winding design and two par-
tial inter-winding designs. The full inter-winding cooling
design had the lowest thermal resistance due to the high
cooling surface area. When comparing two very similar
designs, one as a partial inter-winding design and one
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23
Coolant flow p
ath
Conductors
(a) (b)
Inlet
Outlet
(c) (d) (e)
Fig. 12: Different channel cooling configurations for stator cooling, such as a) full inter-winding for indirect cooling, b) full
inter-winding for direct cooling, c) a single copper pipe, d) multiple copper pipes, and e) alternating copper pipes.
as a full inter-winding design, the thermal resistance
decreased by about 19% with the full coverage design
[100].
Li et al. [101] took a different approach by using cop-
per pipes to make two meandering parallel paths inserted
into the slot at the stator tooth’s bottom. The pipe has an
OD of 16 mm and is clamped to a support structure that
connects the stator tooth to the end cap. Additionally, this
support structure acts as a barrier between the copper
pipe and the windings. The copper pipe was selected
due to its high thermal properties, which allow for more
effective heat dissipation. In the analysis, the copper
windings are assumed to be solid bodies with a realistic
winding thickness (instead of modeling each turn). The
maximum coil temperature is seen in the end windings.
Using an inlet temperature of 30 ◦C and 9.5 kW of losses
in the motor, the maximum temperature rise is almost
60 ◦C, at a maximum end winding temperature of 89
◦C. These results were also validated with the 130 kW
prototype, showing a maximum end winding temperature
of 90 ◦C [101].
Similarly, Le et al. [110] used five copper water pipes
to direct flow in the stator between adjacent windings.
They found that the placement of the inlet and outlet
with respect to one another can impact the temperature
distribution. Since an uneven temperature distribution
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24
in the motor can lead to an unbalanced three-phase
resistance and degrade the electromagnetic performance,
this cooling system is inadequate. However, by placing
the inlet and outlet 180◦from each other, the coolant
temperature rise, and the stator temperature gradient
is reduced. The offset inlet and outlet had the lowest
pressure drop, 25 kPa, compared to 28.6 kPa in the
neighboring inlet and outlet designs. Lowered pressure
drop resulted in a 13% pump energy reduction compared
to the other designs [110].
On the other hand, Odvarka et al. [93] created a
cooling channel through the stator core. It is made from
aluminum, sandwiched between two stator core halves,
and contains fins to guide the flow. With flow at 10 LPM,
1000 W of stator losses, and an inlet temperature of
65 ◦C, they found only a 2.8 ◦C increase in coolant
temperature from the inlet to the outlet.
D. Immersion Cooling
Direct cooling by immersion includes complete motor
flooding, total stator flooding, or semi-flooded concepts.
Due to the high speeds of this application and increased
rotor windage losses, fully flooded solutions are typically
not chosen. In AFPM motors, total stator immersion
cooling is the most common direct cooling method.
Innovations in this design implement baffles, fins, flow
reducers, heat sinks, and intricate coil designs to create
an even flow distribution from inlet to outlet and maxi-
mize the contact area between the coils and coolant.
Camilleri et al. [104] characterized the temperature
distribution of the YASA-750 AFPM motor, which im-
plements a flooded stator design, and has a peak power of
100 kW and a peak efficiency of 94%. They created a test
setup representing the motor’s pole, winding, and flow
passages. The passages, winding, and core were fitted
with 22 thermocouples, and the exterior of the setup
was constructed to minimize heat loss to the ambient.
Tests were conducted at a motor operating point of 1500
RPM and 300 Nm, equivalent to 50 kW in power. Flow
rates between 4 and 10 LPM were tested, resulting in a
laminar flow regime. Decreased flow rate showed more
significant hot spots and poor temperature distribution.
The highest temperature was measured at the inner-mid
winding, away from the coolant and iron core. As the
flow rate increases, the imbalance between the different
flow paths increases, creating a high temperature gradient
in the stator.
Improving on the YASA ltd. design, Camilleri et
al. [108] demonstrated an immersion cooling design
implementing baffles that divide the stator cavity into
four sections creating an even flow distribution between
the stator coils, as shown in Fig. 13. The design reduced
hot spot temperature by 13 ◦C and increased current
density by 7%. The inlet flow rate was 6 LPM, and
the temperature was 80 ◦C. A quarter stator setup was
created to test the design. A total of eight tests were
completed at different flow rates, inlet temperatures and
heat losses which verified simulation results within 6%
of the measured temperature. Authors noted this design
could be enhanced by decreasing the width of the outer
race, and adding flow reducers in the inner and outer
races to increase the head loss in the system [108].
Further improvements to the design are shown in [107],
[120], [121].
Camilleri et al. [120] added a heat sink to the baffled
design, which increased current density by 140% and
decreased hotspot temperature by 87 ◦C compared to
the initial YASA ltd. construction. The design used tra-
ditional concentrated winding AWG14 and the synthetic
dielectric coolant Opticool. The heat sink is made of
0.1 mm sheet copper and is placed between the inner
and middle winding extending out into the stator cavity,
as shown in Fig. 14. The design removes heat from
the inner-mid winding, the hot spot in previous designs
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25
Coolant
Pole Piece
Outlet Inlet
Baffles
Fig. 13: YASA AFPM motor immersion baffle design [108].
[104], [108], and moves it to the iron core. A single coil
thermal setup of a prototype heat sink was completed
and found models were within 3% of experimental
results. The design offers a significantly reduced coil
temperature leading to increased winding life, reduced
resistive losses, and increased current density at the
maximum operating condition. While this design shows
significant improvement in thermal management com-
pared to the baffle design, the electromagnetic effects and
eddy current losses were not assessed with the addition
of the integrated heat sink [120]. Additional losses
induced by the heat sink could impact the efficiency and
effectiveness of this cooling solution.
Iron Three winding
layers
Wrap around heat sink
with fins
Gaps in heat sink
to prevent eddy currents
Extended fins to increase
heat transfer coefficient
Fig. 14: YASA AFPM motor immersion heat sink design [120].
Wanjiku et al. [107] designed a prototype motor that
implements a flooded stator using a fiberglass casing
and baffles to direct flow. The peak power was 50 kW
at 7,500 RPM. The authors studied three different full
inter-winding inlet and outlet channel designs. While
the designs were similar, two had a single channels
per coil, with different inlet and outlet placements. The
third design had two channels per individual coil. The
double cooling design saw the lowest maximum coil
temperature at 175 ◦C, with an inlet temperature of
105 ◦C, 10 LPM, and ambient at 50 ◦C, with WEG.
The rated current for this analysis was 120 A at 2000
RPM. The single-channel designs saw a maximum coil
temperature of 178 ◦C and 205 ◦C for the neighboring
inlet/outlet versus the dispersed inlet/outlet. However,
while the dispersed inlet and outlet design saw the hottest
coil temperature, the pressure drop was the lowest at
5 kPa, compared to 21 kPa and 22 kPa for the other
single and double cooling channels, respectively. They
also found that the coolant significantly impacted the
coil temperature and pressure drop. In the same study,
using the dual cooling channels with engine oil, they
saw both performance metrics increase at 241 ◦C and
31 kPa, respectively. The motor losses were over 15 kW
at the low speed condition of 900 RPM, predominantly
coming from the windings [107].
While Camilleri et al. and Wanjuku et al. implemented
baffles to optimize their direct cooling designs, Liu et al.
[49] used fin-like structures to direct free-flowing oil.
Here, an ironless stator AFPM machine is developed
with concentrated Litz wires. These oil fins are also
used as the winding supporter due to removing the stator
core. By removing the stator iron, the only losses in the
stator are the AC (alternating current) and DC (direct
current) losses. While the rotor losses (4.6 W of the
50 kW machine) and proximity losses are negligible,
the major contributors are DC copper and eddy-current
losses. With an inlet flow rate of 12 LPM (1.2 m/s),
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26
the flow velocity in the stator was found to be about
0.5 - 1 m/s. This is due to the compact, concentrated
winding design, making it challenging to direct the flow.
Assuming adiabatic wall conditions, the maximum end
winding temperature was 100 ◦C, with the rest of the
windings at about 70 ◦C for the 50 kW output. However,
during the experimental testing, a temperature sensor was
placed on the outer end winding and saw a steady-state
temperature of about 80 ◦C. This was caused by natural
convection, which was not present in the simulation [49].
Immersion cooling has a high coolant volume and low
thermal resistance between the coolant and windings,
making it an extremely effective solution. The system
pressure drop is similar to in-direct solutions such as
cooling jackets, end covers, and internal flow channels,
so the pumping power requirements are similar to a
conventional cooling system [96], [107], [110]. Note
immersion solutions use a dielectric oil coolant while in-
direct system typically use WEG [49], [87], [99], [100],
[107], [120]. Dielectric oil coolant has poor properties
and is more expensive, which will be discussed further
in Section VI-A.
Few papers discuss the integration challenges of im-
mersion cooling, including the stator case construction,
structural analysis of the stator bars and casing, flow
channel construction, and pressure drop requirements. In
some high-power, high-RPM applications, the structural
rigidity of the stator may require potting, making this
solution infeasible.
E. Hollow Coils
Many conventional motor windings are manufactured
with copper bar or extruded conductors. Square cross-
section windings are favorable for manufacturing and
achieving a high slot copper fill factor. However, these
compact windings typically compromise cooling inte-
gration. Hollow coils address this issue by placing the
cooling mechanism directly in the heat source [122].
Nitsche et al. [123] indicated that using internal flow
coils allows for eliminating a water jacket leading to
up to 50% more compact designs. Due to the large
conductor cross-section, direct cooling via hollow coils is
applied in generators for power generation [124]–[126].
However, the availability of hollow coil extrusion and the
development of additive manufacturing (AM) has made
this solution more viable.
Essential system design considerations for internal
flow coils include pressure drop, coil geometry, seal-
ing integration, and electrical connections. When using
hollow wire extrusion, considerations must be made for
bending into shape and attaching material for electrical
and thermal connections. Manufacturing via AM has
also been explored due to the increased flexibility in
design. Unlike extrusion, the cross-section does not need
to be consistent through the coils. Further, AM presents
the opportunity to prototype unique cross-sections and
integration geometry quicker and easier.
Lindh et al. [50] designed a prototype 60 kW DSSR
AFPM motor with direct cooled hollow windings. The
windings consisted of hybrid hollow wire made of a
stainless steel tube, with an inner diameter of 3 mm
and a wall thickness of 1 mm, surrounded with Litz
wire. The copper conductor and conduit were split at the
terminals, and the tube was connected to insulating tubes.
The copper fill factor was 37% due to the stainless tubes.
The authors noted that optimization of the conductor
cross-section would achieve a better fill factor of the
conductor. The machine consisted of two coils in parallel
from a continuous 6 m length of the hybrid hollow wire.
Galvanic isolation between the cooling and electrical
system was achieved using a dielectric coolant and
connecting nylon tubes to the stainless tubes at the end
of each parallel pair. All cooling connections were made
outside the stator so leaks could be dealt with easily and
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27
would not harm the machine.
Two prototypes were evaluated by Lindh et al. [50],
one which implemented direct cooling and one that did
not. Eddy current losses were assessed, showing that
only the tubes near the air gap were affected, and overall
eddy current losses were only 5 W. At the test case of
380 Nm, 1500 RPM, and a flow rate of 6 LPM, losses
for the direct cooled machine were 3.5 kW and 3.8 kW
without. The max winding temperature was 86 ◦C with
direct cooling and 137 ◦C without. Water and Chemlube
221 PAO-oil were compared at the max pressure of 600
kPa. Here, the flow rate of the water was increased
significantly, resulting in a 15.9 ◦C drop in maximum
windings temp and an increase in wire current density
by 3.5 times compared to the indirect prototype [50].
Documentation of prototype AFPM motors using this
technology is limited to work by Lindh et al. [50], [109].
Radial flux motors have implemented this technology
more widely with both extruded and AM coils seen in
literature [123], [127], [128]. This subsection will also
discuss these radial flux designs, which could be applied
to AFPM machines.
Extruded profile and hollow coils were assessed by
Reinap et al. [122] in a simulation radial flux machine
and a single coil test. The benefits of this cooling design
were greatest during high power and overload machine
operation. Analyzed conductors included: a solid rect-
angular, hollow rectangle with a hole and a profiled
’U’ in vertical and horizontal orientations. The hollow
and profiled U conductors had a cross-sectional area
of 8.3 mm2, while the solid conductor, with identical
outer dimensions, had a cross-section of 11.4 mm2.
Copper losses in the windings were the machine’s most
significant losses, and the current density reached 100
A/mm2in some locations. Losses in the solid conductor
were 1912 W compared to 2619 W in the directly cooled
conductor. The authors tested transformer and engine oil
at a flow rate of 0.7 LPM and a heat power of 60 m/cm3,
the max conductor temp being 109 ◦C and 99.8 ◦C,
respectively. Testing showed that continuous operation of
the machine could be safely performed at higher current
densities (40 A/mm2and above) with the direct cooling
of the conductor [122].
Wu et al. [127] presented an AM hollow coil design
and validated it in a radial flux 250 kW prototype
motor for a passenger aircraft hybrid powertrain. The
machine achieved a specific power of 20.17 kW/kg and
an efficiency of 95.89%. The coils are aluminum alloy
AlSi10Mg manufactured through direct metal laser sin-
tering (DMLS) AM. The aluminum alloy is 70% lighter
but 40% more resistive than a copper alloy. The windings
were cooled below 110 ◦C at steady state operation (85
kW) despite a winding loss of 5.7 kW. The coolant
channels run along the coil lengths and are connected
in parallel with a manifold, as shown in Fig. 15. Each
parallel path had a flow rate of 0.057 LPM reducing
the pressure drop significantly as channel openings were
only about 1 mm wide [127]. A challenge with AM
parts is balancing surface roughness for increased heat
transfer and pressure drop through the system. AlSi10Mg
samples can have surface roughness from 6 µm to 94 µm
depending on the surface angle and printing parameters
[129], [130]. High roughness can dramatically increase
the pressure drop through small channels like those
described by Wu et al. [127].
Wohlers et al. [128], created two conductor cross-
sections that optimize current density throughout the
winding and maximized surface area. The first is a ’U’
profile coil which can be casted. The second consists
of holes, curves, and cutouts which can only be made
through AM, as shown in Fig. 16. The shape of the
conductor cross-section increases the homogeneity of
the current density distribution within the conductors,
decreasing losses. The authors compared their first de-
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28
Coolant Channels
Coil to Coil
Connection
Manifold Connection
AM AlSi10Mg
Fig. 15: Additively manufactured hollow conductor for radial
flux machine [127].
sign to a solid square conductor and showed their design
could reduce losses by 50% and achieve a current density
of 100 A/mm2without exceeding the coil max insulation
temperature of 180 ◦C. This reduction in losses allows
for a broader range of operating frequency and current,
increasing the power of the machine. A prototype AM
coil was made through direct metal laser sintering of
AlSi10Mg and tested with deionized water coolant. The
test results were in agreement with the simulation results.
Calculations determined the maximum current density
for this coil to be 130 A/mm2when flood cooled by
deionized water [128].
Fig. 16: Additively manufactured profiled conductor cross-
section [128].
Hollow coils present a direct cooling solution that
allows potting of the stator for increased structural
rigidity and when combined with immersion, can in-
crease current density dramatically. Compared to in-
direct solutions, increased allowable current density and
an integrated thermal management system will decrease
motor size. This concept also has versatility as hollow
coils could be used in all AFPM motor topologies.
However, hollow coils inherently have a poorer copper
fill factor than solid coils, meaning the stator may be
bigger than immersion solutions with concentrated wind-
ings. Further, high system pressure drop can occur due to
the high viscosity of the dielectric coolant and the small
channel size. Lindh et al. [50] recorded pump pressures
up to 600 kPa, requiring a larger pump, increasing sys-
tem mass and parasitic losses. Over the motor’s lifetime,
pumping power requirements may increase significantly
due to fowling in the small channels, although more
research is required.
F. Heat Pipes
Heat pipes (HPs) are efficient heat transport devices
that use a phase change fluid. Their equivalent thermal
conductivity can reach 100,000 W/mK [131], making
them ideal for thermal management in high-power elec-
tric machines. Prototype machines implementing this
technology have seen improvements in thermal manage-
ment and power density [110], [132]–[137]. However,
Le and Mueller [110], [136] present the only use in
AFPM motors seen in literature. Despite their high
heat transfer efficiency, the use of HPs in commercial
electric machines has been limited due to cost and
reliability [131], [132]. Studies noted challenges with
increased power losses when adding HPs to the stator.
However, recent developments in additive manufacturing
and advanced materials could soon make this technology
a more practical cooling solution.
Wicked HPs are the most popular for use in electric
machines. There are many types of wicks; however,
conventionally manufactured meshed wicks are usually
chosen for their compromise between cost and perfor-
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29
Evaporator Condenser
Wick
Fig. 17: Copper-water HP diagram [131].
mance [131], [132]. In an HP, there are two sides: the
hot side, the evaporator, and the cold side, the condenser,
as shown in Fig. 17. Ambient heat evaporates the fluid,
forcing it down the pipe to the condenser. The vapor
then condenses into the wick and travels back to the
evaporator. Wicks must be carefully designed for the
intended operating limits to prevent ’dry-out’. ’Dry-
out’ can occur when the heat flux on the hot side is
too high or when the pumping pressure drop is less
than the gravitational, liquid, and vapor pressure [131].
Alternatively, if the heat flux is too low for the phase
change fluid to vaporize, the HP will exhibit the thermal
conductivity of the shell material [110]. For this reason,
it is critical to carefully choose the phase change fluid,
HP casing material, and wick type for the intended
operating conditions and temperatures.
Failure in HPs is typically caused by issues dur-
ing manufacturing or improper design for the intended
application. These issues include insufficient cleaning
of the interior, improper filling, leakage, incompatible
materials, dry-out, and incorrect orientation with respect
to gravity. These issues could lead to poor reliability of
the HP and, therefore, motor [138].
Le et al. [110] designed an AFPM motor for electric
vehicles. The thermal management system consists of a
water jacket, as described above, and flat HPs situated
between the stator slots. The HPs are 2 by 8 by 85
mm with a copper casing and water as the working
fluid. The equivalent thermal conductivity achieved was
70,000 W/(mK) within a temperature range of 80 to 110
◦C [110].
Wu et al. [139] work show the integration of HPs
inside an AM coil for a radial flux machine, as shown
in Fig. 18. The addition of the HP improves cooling
capabilities. However, the HP and conductor are separate
parts that cause losses to increase by 8% compared to a
coil without the HP [139].
Coil to coil
connection
Heat pipes
AM AlSi10Mg
Fig. 18: Additively manufactured coils with integrated heat
pipe [139].
In the AM space, materials such as polymers and
ceramics have been explored for manufacturing HPs as
they do not induce additional losses in the motor. How-
ever, these materials are far less effective due to reduced
material conductivity and increased contact resistance.
Szymanski et al. [140] developed an aluminum alloy
ammonia HP with a hybrid sintered wick using DMLS.
AM allowed researchers to fine tune the wick geometry
in ways that would be impossible using conventional
manufacturing methods. The optimized wick decreased
pressure drop without compromising capillary pumping
capacity and achieved a 10% increase in heat transfer
compared to the equivalent conventional HP. The study
also tested a variety of wick structures and compared
densities and geometries. The authors noted that pre-
dicting the mechanical properties of these AM HPs
is challenging, and research still needs to be done to
understand these parameters [140].
Further research must be done in the additive man-
ufacturing space to understand the impact of printing
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30
parameters and powder size on the thermal performance
of HP. One future direction of high-performance motors
in heat pipe technology is printing HP geometry in
stator windings. Combining the conductor and HP will
eliminate additional losses created as separate parts and
maximize the thermal management system’s thermal
conductivity.
The induced losses and reliability issues in HPs should
be addressed when considering this thermal management
solution. Further, the HP condenser system will also need
to be designed. Le et al. [110] implemented a water
jacket, increasing the weight and volume of the motor
design.
VI. MATE RI AL S AN D COOLANTS
A. Coolants
Dielectric coolants are typically low viscosity oils
used for direct cooling applications. The benefits of
these coolants are: that they do not form ions, provide
insulation in stator flooding configurations and conduct
heat away from the source. Oil-based coolants have a
lower heat capacity and higher viscosity than water-
based coolants making coolant selection critical for each
use case. Synthetic oils, including those in Table IV
[141]–[147], are less sensitive to sludging or fouling
from thermal oxidization, making them a more favorable
choice in direct cooling systems.
Deionized water can also be used, however, the fluid
contact materials must be carefully assessed as deionized
water can quickly form ions [50]. Water-based coolants
such as WEG are typically used for indirect cooling
due to their high thermal conductivity and low viscosity.
These applications usually do not need a dielectric fluid.
Table IV lists available coolants for liquid cooling and
their properties. This is in no way a comprehensive
list of all available coolants. Other coolants reported in
literature include Galden HT 135 [141], and Syltherm™
800 [142], as well as engine and compressor oil in
experimental setups [50], [128].
Coolant selection should consider direct cooling per-
formance characteristics like thermal conductivity and
system level thermal management, for example, weight,
pumping power, material compatibility, and maximum
allowable temperature. YASA Ltd. [18] floods the stator
with OptiCool™-H [143], which has a density less than
water, decreasing the weight of the thermal management
system. However, the maximum coolant temp is only 130
◦C, which could limit the duration of peak power.
It should also be noted that synthetic oil properties
fluctuate significantly depending on temperature. For
example, the viscosity of AmpCool® AC-110 is 41.10
cSt at 0 ◦C, 8.11 at 40 ◦C, and 2.22 at 100 ◦C [146].
High coolant viscosity at low temperature will have
implications for the pumping power required and could
require a system warm up period.
B. Encapsulants
Encapsulating the stator is common in high power
high torque AFPM machines to provide a thermal path
from the conductors to the housing and for structural
rigidity. Typical potting materials have poor thermal
conductivity (less than 1 W/mK) and have maximum
operating temperatures of around 180 ◦C [151]. The
filler materials are added to the potting material matrix
to manipulate the material properties.
Glass is added to increase mechanical properties at
the cost of thermal conductivity. Metallic filler materials
increases thermal conductivity while decreasing mechan-
ical and insulative properties [118]. Metallic materials
typically have a higher density than the matrix material
causing increased machine weight. Further, adding filler
in general increases the potting material viscosity and
could cause challenges during encapsulation, leading
to air pockets [151]. Thermal conductivities of over 4
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31
TABLE IV: Coolants Used for Stator Thermal Management
Product Manufacturer Coolant Base Density
(kg/m3)
Viscosity
(cSt)
Thermal
Conductivity
(W/mK)
Max Temp
(◦C)
Dielectric
Constant
OptiCool-H
[143] DSI Venture, Inc. Hydrotreated
Paraffin Oil
825
@16 ◦C
1.4
@100 ◦C
0.1344
@100 ◦C130 Low
Opteon SF10
[144] Chemours Florinated Fluid 1580 0.71
@25 ◦C
0.077
@25 ◦C110 Low
Novec 7500
[145] 3M™ Fluorocarbon Oil 1630
@20 ◦C
0.71
@25 ◦C0.0591 128 Low
AmpCool AC-110
[146] Engineered Fluids Synthetic Oil 8200
@15 ◦C
2.22
@100 ◦C
0.1325
@100 ◦C180 2.08
THERMINOL 68
[147]
Eastman Chemical
Company Synthetic Oil 969
@100 ◦C
0.26
@100 ◦C
0.117
@100 ◦C360 2.8
@23 ◦C
Ethylene Glycol
[148] —Hygroscopic Liquid 1113
@20 ◦C
19.83
@20 ◦C— 187* 37.7
@25 ◦C
Deionized Water†
[149], [150] —Water 997
@25 ◦C
1
@20 ◦C
0.606
@25 ◦C99.9* 80
@25 ◦C
* Coolant boiling point
†To be used as a reference point
W/mk have been reported, however, above this value
there can be issues with viscosity and diminishing in-
crease in thermal management performance [151], [152].
VII. FUT UR E TR ENDS AND INNOVATI ON S
Our team is currently working on the development
of the additively manufactured (AM) hollow conductor
technology (copper and aluminum). As explained in Sec-
tion V-E, hollow conductors represent a very promising
cooling system approach, and the recent improvements
in AM open new scenarios considered impossible just
a few years ago. Despite its efficiency, the AM hol-
low coils pose several challenges from the electrical,
manufacturing, and fluid-dynamics standpoints. The first
problem to be solved is the electrical conductivity. Most
of the copper powder alloys have an IACS not as high as
extruded copper. Recently some important academic and
industrial effort has been made to support pure copper
printing by means of new green laser technology, but
the large scale availability is still far. Another issue is
the uniformity of the conductivity. The electric machine
phases need to have a well balanced resistance. A highly
controlled manufacturing process is crucial to ensure
consistent physical properties. Again, great progress have
been made in the control of these properties during the
3D-printing and the post-processing such as heat treat-
ments, but the traditional wire extrusion is still a standard
reference. AM, however, has one important advantage
compared to the extruded hollow or solid bar conductors:
the minimization of the welding joints. By enabling
the realization of complex geometries in a single piece
instead of multiple parts, AM makes possible to print sets
of coils and eventually the whole winding without need
of welding. The welding of conventional coils is a source
of uneven electrical extra resistance which can cause
unbalanced phases and hot spots in the winding. Porosity
is also a well known problem of metal 3D-printing,
and in the context of the hollow conductors becomes
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32
particularly important. The porosity, in fact, can not only
raise the resistance of the hollow conductor regardless
of the metal powder utilized, but also compromise the
sealing of the liquid cooling circuit. Highly accurate
laser power settings are therefore essential. Dimensional
stability is finally the last main factor for the AM hollow
coils to succeed. The mechanical tolerances achievable
by means of extrusion are still way better than those of
3D-printing, as well as the surface roughness. In order
to maximize the copper fill factor, the distances between
different coils, or consecutive turns, or between the coil
and the stator bar are usually tight. Without an extremely
precise control of the geometric tolerances, the coils
might not be able to fit in the full machine assembly,
and the scrap rate can become industrially unbearable.
Besides, unexpected variations of the cooling channel
size would have adverse effects to the pressure drop,
already high in this particular cooling system design.
The AM has also another geometry-consistency issue,
the warpage, which in the special case of the electrical
coils results in some sort of twisted shape. Of course,
that is not acceptable but heat treatments can be highly
beneficial from this point of view. On the other hand,
the AM removes the problem of the maximum bending
radius which affects the extruded hollow conductors. In
conclusion, the AM hollow conductor can be one of the
key-technologies to increase the electric current density,
and therefore the power density, of the new generation
AFPMs but the scientific research still needs to find
solutions to several technical challenges.
VIII. CONCLUSION
Thermal management for AFPM is the most chal-
lenging design aspect and is often the limiting factor
in increasing machine capabilities. The rotor magnets
and stator conductors are the two most dominant elec-
tromagnetic loss sources in AFPM motors. Cooling these
components is crucial to achieving the ambitious require-
ments of heavy-duty vehicles and aerospace applications.
The SSDR topology is selected by AFE motors,
Emrax, Evolito, MagnaX, and YASA Ltd., whereas the
DSSR was chosen by AVID, MAGELEC, and Phi Power,
Table I. It would be challenging to compare these electric
motors’ thermal management system performance as it
varies considerably depending on the configuration, loss
distribution between stator and rotor, main operating
point, and the intended application. Nevertheless, some
conclusions can be made about the general trends.
Section II-B explains that the baseline DSSR con-
figuration has a higher mass than the SSDR (YASA)
due to two stator yokes and the additional end-winding
length. For this reason, the companies that use DSSR
tend to increase the maximum speed to compensate
for this disadvantage. The single rotor is theoretically
symmetrical, and the magnetic load in DSSR is also
more balanced than in SSDR, as the rotor is between
two identical stators. The higher speed enables these
companies to reduce the torque needed for the same
power rating, downsizing the machine.
A major challenge of DSSR motors is the cooling
system design. First, air-cooling conditions for the ro-
tor is poorer and generally, the cooling enhancements
discussed in Section IV are not feasible. Magnetic loss
in the rotor is typically higher due to higher electric
frequency caused by higher speeds. Distributed wind-
ings are adopted instead of FSCW to decrease these
losses. However, their small diameter and greater number
of slots make it challenging to utilize direct cooling
methods. The slots are typically too small for cooling
channels, and the stator yokes prevent stator cooling via
immersion. The most prevalent cooling system of DSSR
is indirect cooling by end cover jackets, discussed in
Section V-A2. Here, the thermal contact resistance with
the windings is inherently worse than direct methods,
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33
and the axial length of the motor is increased.
DSSR operating at higher speeds also has conse-
quences on the motor application. Lower speed SSDR
can achieve high power density and high torque density
simultaneously, thus suitable for direct-drive. In contrast,
higher speed DSSR usually requires more complicated
gearboxes with higher ratios, adversely affecting the
vehicle’s overall mass, cost, reliability, and efficiency.
Rotor cooling is essential to prevent reversible and
irreversible demagnetization of the magnets. The double
rotor configuration is typically air-cooled with rotor
blades. Depending on cooling requirements, fins, vents,
and mesh structures can be used to increase the heat
transfer coefficient of the rotor carrier. Studies have
shown that these air-cooling features can be optimized to
minimize windage losses, maximize cooling effect, and
prevent demagnetization.
In addition, stator cooling can be accomplished by
several strategies, including jackets, fins, channels, im-
mersion cooling, hollow coils, and heat pipes. To im-
prove the cooling performance of the design and increase
the power density, the chosen solution should consider
integration between two or more of these thermal tech-
nologies. For example, a water jacket, one of the most
popular solutions, significantly increases motor size and
may not provide sufficient cooling depending on the
machine’s power density and losses. Depending on the
machine topology, the most successful strategy seen
in commercial AFPM machines is immersion cooling.
Enhancements to these flooded concepts, such as baffles,
extended heat transfer surfaces and hollow coils, further
improve cooling capabilities.
With increasing demand for high-power electric ma-
chines, innovative thermal management technologies
must be thoroughly investigated. Rotor blades and sta-
tor immersion cooling in SSDR configurations show
promising results and have started to see commercial
success. Additively manufactured hollow or profiled
conductors show promise in increasing current density,
and, therefore power density indicating machines for new
stage of transportation electrification will be achievable
in the future. Further research should investigate testing,
validation, and long term reliability of these emerging
solutions in prototype machines. The challenges associ-
ated with integration should also be studied to understand
which strategies may see better success in commercial
high-performance AFPM motors. Several technologies,
however, show promising results, indicating machines
for the new stage of transportation electrification will
be achievable.
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This article has been accepted for publication in IEEE Transactions on Transportation Electrification. This is the author's version which has not been fully edited and
content may change prior to final publication. Citation information: DOI 10.1109/TTE.2023.3242698
This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/