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Back-iron Extension Thermal Benefits for
Electrical Machines with Concentrated
Windings
Fengyu Zhang, David Gerada, Zeyuan Xu, Xiaochen Zhang, Chris
Tighe, He Zhang and Chris Gerada
Faculty of Science and Engineering, University of Nottingham Ningbo
China, 199 Taikang East Road, Ningbo, 315100, Zhejiang, China.
First published 2019
This work is made available under the terms of the Creative Commons
Attribution 4.0 International License:
http://creativecommons.org/licenses/by/4.0
The work is licenced to the University of Nottingham Ningbo China
under the Global University Publication Licence:
https://www.nottingham.edu.cn/en/library/documents/research-
support/global-university-publications-licence.pdf
IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS
Abstract— This paper proposes a novel, low-cost,
effective way to improve the thermal performance of
electrical machines by extending a part of the back-iron
into the slot. This modification helps in reducing the
thermal resistance path from the center of the slot to the
coolant, however its thermal benefits must be clearly
evaluated in conjunction with the electromagnetic
aspects, due to the higher iron losses and flux-leakage,
and furthermore such an extension occupies space which
would otherwise be allocated to the copper itself. Taking a
case study involving an existing 75kW electric vehicle (EV)
traction motor, the tradeoffs involving the losses, flux-
leakage, output torque, torque-quality and the peak
winding temperature with back-iron extension (BIE) and
without are compared. Finally, experimental segments of
the aforesaid motor are tested, verifying a significant
26.7% peak winding temperature reduction for the same
output power with the proposed modification.
Index Terms—Thermal management, machine cooling,
power density, thermal analysis, thermal resistance
network, slot cooling
I. I
NTRODUCTION
ITH
the increasingly stringent emissions legislations,
there is an unprecedented demand for transport
electrification, be it for rail, marine, aerospace or
automotive. For some applications this involves hybridization
of the conventional engine-based systems, while for other
applications all-electric architectures are developed [1, 2].
The core component in these electrified architectures is the
electrical machine, and hence its performance improvement
merits detailed research in order to achieve important
Manuscript received November 9, 2018; revised January 6, 2019;
accepted February 19, 2019. This work was supported by China NSFC
under Grant 51607099 and by the Ningbo Science & Technology
Beauro under Grant 2017D10029.
F. Zhang, D. Gerada, H. Zhang (corresponding author), and C.
Gerada are with the Power Electronics, Machines and Control group,
University of Nottingham Ningbo China, Ningbo 315100, China (e-mail:
fengyu.zhang@nottingham.edu.cn; David.Gerada@nottingham.ac.uk;
he.zhang@nottingham.edu.cn; Chris.Gerada@nottingham.edu.cn ).
D. Gerada, Z. Xu, X. Zhang and C. Gerada are with University of
Nottingham, NG7 2RD, UK (e-mail: David.Gerada@nottingham.ac.uk;
Zeyuan.Xu@nottingham.ac.uk; Xiaochen.Zhang@nottingham.ac.uk;
Chris.Gerada@nottingham.ac.uk ).
C. Tighe is with Electrical Cooling Solutions Ltd, UK (email:
chris.tighe@electricalcoolingsolutions.com).
performance metrics such as high efficiency, high power
density (kW/kg, or kW/L), and equally important a low cost
($/kW) in order to make a strong business case and increase
market proliferation [3]. While material developments are an
enabler for improved performance, often new materials come
with an increased cost premium. It has been shown that
improvements in thermal management are a key enabler to
push machines’ technological boundaries [2].
Automotive and railway traction machines are required to
specifically have a high torque density, low inertia, wide
constant power -speed ranges and high reliability as demanded
by the market. Various ambitious cost and performance targets
have been set by related governmental bodies, such as the
Department of Energy (DoE) in the USA, which has set the
FreedomCar 2020 targets, and the Advanced Propulsion
Centre (APC) in the UK [4]. The maximum allowable
temperature in an electrical machine’s stator is typically
determined by the constituent insulation materials’ thermal
limit pertaining to the windings, with various industrial classes
of insulation set, such as class H (180
o
C) and class C (220
o
C). Thus improved cooling enables higher power densities in
electrical machines, which helps to reduce weight, volume and
cost. Improved thermal management pushes winding current
density higher before reaching the maximum winding
temperature limit, which is one critical factor determining the
motor torque production [5]. Reducing losses generated in the
machine is of course another alternative way from heat
sources view to improve the thermal condition of the motor. In
the ideal case, the two methodologies of cooling improvement
and losses reduction are coupled, in order to find the optimum
tradeoff between the two in light of the specific optimization
targets.
For higher power density machines, using water or oil as a
cooling medium, yields better thermal performance with
respect to the air-cooled machines. CFD and experimental
results indicate that wet stator cooling [6, 7] leads to markedly
better thermal conditions compared to the indirect cooling
methods and the traditional water-jacket cooling [8]. In [9] the
thermal performance of a machine with both the stator and
rotor flushed directly by coolant is investigated. However,
with either the wet stator or totally wet cooling method, lots of
accessory components have to be mounted, which increase the
system complexity and reduce reliability. Also having stator
and rotor wet cooling adds extra rotational friction losses and
is thus only suitable for very specific circumstances such as in
Back-iron Extension Thermal Benefits for
Electrical Machines with Concentrated
Windings
Fengyu Zhang, David Gerada, Zeyuan Xu, Xiaochen Zhang, Member, IEEE ,Chris Tighe,
He Zhang, Senior Member, IEEE and Chris Gerada, Senior Member, IEEE
W
1
IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS
the case of lower speed machines. Therefore, in industrial
machines (for example those used in automotive traction), the
external cooling jacket is widely adopted as a robust cooling
methodology.
Often, one major bottleneck in the thermal improvement of
motors with indirect cooling is the equivalent thermal
conductivity perpendicular to the winding orientation in the
slot which is poor due to the multiple layers of insulation
involved (typically the insulation build up consists of wire
enamel, slot impregnation resin, air-pockets, and the slot liner)
[10, 11]. In light of this some novel cooling methods targeting
directly the heat directly inside the slot have been proposed.
As described in [12-17], slot water jacket cooling provides an
efficient way to reduce the winding hot spot temperature,
albeit the copper losses are increased due to the space
occupied by the slot jacket. Moreover due to electrical
conductivity of water, it is imperative that sealing is ensured
within such a configuration. A novel heat path has been
proposed and experimentally validated in[5], where a piece of
thermally conductive material is inserted into the slot of a low
frequency machine, benefiting approximately from a 40% hot-
spot temperature reduction for the same current loading.
However, this method adds extra eddy current losses, which
might nullify the effect the heat path brings, especially with
higher frequency machines as in the case of automotive
traction. In [18], the authors study the effects of different
winding design on the motor loss distribution and the
balancing effects between stator iron loss and rotor loss.
Electromagnetic and thermal aspects are closely coupled
and interlinked in the machine design process, which makes it
irrelevant to modify the motor geometry from the thermal
point of view only. For thermal engineers, it is thus critical to
ensure that the benefit in any proposed novel thermal
technique is not outweighed by the deterioration in other
aspects such as the torque (and its quality), or the machine
losses. In [19], flux-barriers are inserted in the stator
lamination and simulation results are compared to the more
conventional indirect water cooling jacket, with benefits of
10% temperature reduction in the slot region of the machine
observed. However, the aforementioned arrangement needs
extra pump power and sealing due to possible water leakage.
The effects of parallel and trapezoidal slot shapes on the loss
distribution and the heat transfer characteristics are compared
and researched in [20]. It is concluded that parallel slots are
more advantageous, with a 35% improvement in the heat
transfer path between the slot and the stator core pack noted,
which yields a net 8% improvement in the output per active
weight capability of the machine. The benefits noted however
can only be achieved with open-slot modular stator windings.
This paper proposes and experimentally verifies a novel
way to decrease the equivalent thermal resistance between the
hot spot and the coolant by extending a small part of back-iron
into the middle of the slot, where the hot spot is located. This
paper is organized as follows: Section II introduces the
principle behind the back-iron extension (BIE) and discusses
multi-domain models and considerations involved in
optimizing the geometry of the back-iron extension. The
results of the optimization are experimentally verified and
further discussed in Section III. Finally the conclusions of this
research are summarized in Section IV.
II.
BACK
-
IRON EXTENSION
As mentioned in the introduction, this paper proposes a part
of the back-iron extended from the middle point of the slot
bottom along the centreline of the slot to the slot opening, as
shown in Fig. 1. In water-jacket cooled machines, most of the
heat generated in the slot is dissipated to the stator lamination,
through the slot wall, to the stator tooth, or directly to the
stator back-iron, and then to the jacket coolant. Back-iron
extension provides another alternative heat transfer path for
heat to be removed to the back-iron, increasing the contact
area between the slot and stator lamination, and shortening the
equivalent thermal resistance between the hot spot in the slot
and coolant, which is typically located in the middle of the
slot.
BIE depth
BIE width
Fig. 1. Back-iron extension geometry
This section analyses the thermal benefits of the back-iron
extension, taking as a case study an existing 75kW 12slot, 8-
pole Permanent Magnet machine, with concentrated winding
used for an electric vehicle shown in Fig. 2 and the relevant
data listed in Table 1. The electromagnetic effects of the back-
iron extension are also considered, due to the flux leakage
through back-iron extension and additional iron-losses
generated within. The optimized back-iron extension geometry
is generated based on the rated power of the traction machine
under investigation, with guidelines given about the optimized
BIE geometry on machines of different size.
Fig. 2. Traction machine
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A. Thermal modelling of BIE
Lumped Parameter Thermal Network (LPTN), is a widely
used tool in electrical machine thermal analysis which
provides quick answers for cooling parametrized studies for a
selected cooling methodologies. A LPTN thermal model is
developed to investigate the thermal performances of the BIE
cooling concept for the machine described in previous section.
Heat transfer paths in both the radial and axial directions are
simulated in the 3D thermal model, with three longitudinally
divided parts in the machine core and one part for each end-
winding. The longitudinal lengths of the core machine parts
are 20 mm, 100 mm, and 20 mm, adding up to the net 140mm
active length detailed in Table I. Exploiting symmetry 1/24
th
of the stator and 1/16
th
of the rotor are simulated. The half slot
includes 25 nodes, with a 5×5 distribution to achieve a
detailed temperature distribution. Fig. 3 (a) shows the radial
thermal network of the motor with BIE in the center of the slot
and (b) displays the corresponding axial thermal network,
where the 25 nodes in the slot in (a) are not shown on purpose
to give a clear view of the heat flow longitudinally.
Heat losses from the water jacket of the traction machine to
ambient air are neglected, as they are very small compared to
the heat transfer to the coolant. An isothermal boundary
condition is applied to the coolant based on an assumption that
the water jacket is sufficiently effective to remove all the heat
generated. The two sides where the BIE and stator tooth are
connected to other parts in Fig. 3 (a) are assumed adiabatic
due to the geometrical symmetry.
A heat conservative equation for any node and all its
neighboring nodes, shown in Fig. 3, can be expressed by the
following (1):
0
nT T
j i
j i Rij
qi
(1)
where node j is an adjacent node to node i in the thermal
network, q
i
is the loss generated in node i; T
i
, T
j
are the
temperatures in node i and node j respectively, while R
ij
is the
thermal resistance between node i and node j.
The losses considered in the modelling are I
2
R copper
losses, stator iron losses, rotor iron losses and magnet losses.
All losses are assumed uniformly generated in the
corresponding parts of the machine, added to the nodes as a
heat source, respectively. Transient temperature analysis can
be determined by adding thermal mass to each node.
TABLE
I
TRACTION MACHINE PARAMETERS
Machine type Three – phase PMSM
Machine Rating:
Voltage 384 V 480 V
Rated/ peak
power 37/74 kW
Rated/ peak
torque 126/382 N.m
Maximum speed 10000 rpm
Working time at
peak power ≥60 s Working
temperature -40 ~ 105
℃
Efficiency ≥95% Coolant:
Ethylene
water glycol
– EWG 50/50
Geometry: Materials:
Outer diameter 245 mm Magnet N38EH
Active length 140 mm Iron sheets M235 – 35 A
Winding copper
back iron
extension
Heat
source
Thermal
resistance
water coolant channel
aluminium housing
stator back-iron
stator tooth
air gap
magnets
shaft
rotor back-iron
(a)
water coolant channel
aluminium housing
stator back-iron
tooth
air gap
magnets
rotor back-iron
shaft
rotor fluid
stator fluid
end-winding end-winding
(b)
Fig. 3. Thermal network of developed motor with back-iron extension
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Only conduction heat transfer and convection heat transfer
are included in the LPTN thermal model. The thermal
resistances by conduction between neighboring nodes applied
in (1) can be determined from [21-23]:
=
⁄ (2)
where is the distance between the nodes along the heat flow,
is the cross-sectional area (perpendicular to the heat
flow), and is the thermal conductivity of the node material.
Different values of thermal conductivity are applied to the
radial and axial directions of the winding in the slot, stator and
rotor iron. Effective thermal conductivity
is used in
calculating the thermal conductivity of the winding in the slot
[11], taking into account the various constituent materials.
Equation (2) is the classic equation for conductive thermal
resistance. An alternative method is presented in [24, 25]to
calculate the thermal conductive resistance, which suggests a
more accurate temperature distribution considering internal
heat generations, such as copper and iron losses.
Convective heat transfer occurs primarily in three places
where fluid flows over the machine solid parts: (i) between
coolant and housing, (ii) the air-gap between rotor and stator
core, and (iii) machine end region where end-winding is
directly cooled by air in the region. For convection, the
thermal resistance in (1) is calculated from (3):
= 1 ℎ
⁄ (3)
where
is the surface area of solid to the fluids and h is the
heat transfer coefficient and can be calculated as follows.
The flow inside the water cooling jacket is considered
turbulent. The heat transfer coefficient between the coolant
and aluminum housing applied in (3) can be calculated by (4)
to (6) [26]:
=
(
⁄ )×(
)×
.×(
⁄ )
.
×
⁄
(4)
ℎ =
⁄;
=
⁄;
= 4
⁄ (5)
= (0.79 ×(
) − 1.64)
-2
(6)
where f is friction factor, is the thermal conductivity of the
fluid (water), R
e
is the Reylolds number, P
r
is the coolant
Prandtl number, D
h
is the hydraulic diameter, is the water
velocity, and is the dynamic viscosity of water. In the above
equations S and P are the cross section area and peripheral
length of the water channel respectively.
The air flow across the air-gap is calculated from (7) to (8)
[21] :
ℎ = 0.386
0.5
0.27
/
(7)
=
⁄;
=
⁄;
= 2 (8)
where is the thermal conductivity of the air, T
a
is the Taylor
number, P
r
is air Prandtl nmber, D
h
is the hydraulic diameter,
R
e
is the Reylolds number, l is air-gap length, R
r
is the rotor
radius, is the air density, is the air velocity, and is the
dynamic viscosity of air.
In order to calculate the heat transfer coefficient between the
end-winding and stator/rotor fluid in 0 (b), equation (9) is used
[21]:
=
×(1 +
()
) (9)
where is the rotor velocity while k
1
, k
2
and k
3
are curve
fitting coefficients. In the thermal network, k
1
=15,
k
2
=0.15,k
3
=1.
In this paper, both the width and depth of the back-iron
extension are optimized to provide the best cooling benefits.
The back-iron extension width ranges from 0 to 4 mm, while
the BIE depth is varied from 0 to the entire slot depth of
21.7mm, as described by Table 2. The introduction of BIE,
reduces the slot area available for winding (winding area),
resulting in a decreased copper fill factor as shown in Fig. 4. It
is worth noting that in this paper, the ‘copper fill factor’ in
Fig. 4, is defined as the ratio of copper area for each motor
with the proposed BIE divided by the original motor slot area,
maintaining the same ratio of copper area to the available
winding area. The available winding area reduces with the
introduction of the BIE, hence the amount of copper in the slot
reduces with respect to the original design without the BIE.
The thermal resistance formulas used with the circuit of the
original motor are similar to those used with the motor using
BIE, which are based on node geometry and material thermal
conductivity, except for the resistances on the added heat
transfer path(i.e. BIE). The loss sources are also different for
the two thermal models, as will be discussed in more detail in
the subsequent section.
TABLE
2
BACK
-
IRON EXTENSION GEOMETRY VARIATIONS
Simulation
Back-iron Extension
Depth
Back-iron
Extension Width
mm
BIED1 4.34 (1/5 of slot depth)
0.1 – 4 (0.1 mm)
interval gap
BIED2 8.48 (2/5 of slot depth)
BIED3 10.32 (3/5 of slot depth)
BIED4 17.36 (4/5 of slot depth)
BIED5 21.7 (full slot depth)
0 1 2 3 4
0.36
0.38
0.40
0.42
0.44
0.46
Fill Factor
Back-iron extension width (mm)
BIED1
BIED2
BIED3
BIED4
BIED5
Fig. 4. Copper fill factor with BIE geometry variations
B. Electromagnetic aspects of back-iron extension
Finite Element Analysis (FEA) is used to analyze the non-
linear electromagnetic effects and to combine with the thermal
model of the previous section so that any geometric change
can be instantly assessed from both the electromagnetic and
the thermal domains. It is worth noting that adequate mesh
refinement is conducted in the back-iron extension. The
introduction of the BIE causes additional flux leakage through
the extension, especially when the back-iron extension depth
is large which in turn reduces the output torque of the motor.
Fig. 5 shows the flux density contour for the same BIE
dimension drawn in Fig. 3(a). Because of the additional flux
leakage introduced by the BIE, higher current will be
demanded to generate the same output torque, resulting in
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IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS
increased copper losses generated in the slot and higher iron
loss generated in the stator core itself. Furthermore, additional
iron losses are generated within the BIE itself.
Flux density (T)
Fig. 5. Flux density contour at rated torque with BIE
Considering BIE depth and width variations, analysis is
conducted with an input current of 90A (this current
corresponding to the rated current of the original motor) firstly
and output torque is calculated, as shown in Fig. 6 (a). In this
figure, the different BIE depths (BIED 1-5) corresponding to
Table 2 are represented by the different lines, while for each
depth, different BIE widths are considered on the x-axis.
When the BIE depth is small (under 8.48mm), the output
torque is not really affected no matter how wide the back-iron
extension is. This is within expectation, as it is difficult for
flux leakage to be significant with a shallow back-iron
extension as verified by Fig. 6 (b), which plots the maximum
flux density in the air-gap corresponding to the lines in Fig. 6
(a) for a BIE width of 4mm. The maximum flux density in the
air-gap of the BIE motor is almost same with that in the
original motor when the output torque is not affected by the
BIE introduction. On the other hand, with BIE depths beyond
10.32mm (BIED4, BIED5), the maximum flux density
decreases and torque starts to show a reducing tendency and
drops by up to 1.59 % from 107.3Nm down to 105.6Nm when
the back-iron extension width is 4 mm, and the depth is equal
to the slot height (i.e. BIED5).
Fig. 6 (c) shows the required increase in current (beyond the
original 90A) to keep the output torque constant as with the
original machine. The maximum increase in current is 2.17 %
when the torque drops to its lowest value of 105.6 Nm.
Torque quality is also another important aspect which needs
assessment when proposing such a geometrical modification.
Fig. 7 shows the increase in torque ripple for various BIE
geometries with a maximum increase in torque ripple of up to
10% for the same output torque.
0 1 2 3 4
105.0
105.5
106.0
106.5
107.0
107.5
Output Torque (Nm)
Back-iron Extension Width (mm)
BIED1
BIED2
BIED3
BIED4
BIED5
(a)
1.4148 1.4146 1.4141 1.413
1.4104
1.4005
Original BIED1 BIED2 BIED3 BIED4 BIED5
1.39
1.40
1.41
1.42
Maximum Flux Density in the Airgap (T)
(b)
0 1 2 3 4
90
91
92
Back-iron Extension Width (mm)
Current (A)
BIED1
BIED2
BIED3
BIED4
BIED5
(c)
Fig. 6. (a) Output torque reduction with back-iron extension, (b)
maximum air-gap flux density with back-iron extension and (c)
required current increase to maintain the same output torque
0 1 2 3 4
0
3
6
9
12
BIED1
BIED2
BIED3
BIED4
BIED5
Percentage (%)
Back-iron Extension Width (mm)
Fig. 7. Torque ripple increase for various BIE geometries
To maintain the same output torque, the effect of current
requirement increase on the thermal performance of the
machine reflects in both the loss magnitude as well as the loss
distribution. Higher current increases the flux density within
the stator core due to the stronger armature field, and hence
the iron losses therein increase. Furthermore, additional iron
losses are generated within the back iron extension itself. The
copper I
2
R losses increase as well, both due to the aforesaid
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increased current requirement, as well as due to the increased
resistance since the BIE occupies space which would
otherwise be available for copper. Therefore, with increased
BIE width at constant BIE depth, copper losses I
2
R increase
due to increased electrical resistance resulting from the
reduced copper fill factor, within periods when current
remains same in Fig. 6 (c). Fig. 8 summarizes the increase in
stator iron losses (which also include iron losses generated
within the BIE itself), and copper losses with respect to the
original design. For the copper losses (P
cu
) the loss increase is
up to 26%, while the iron losses (P
Fe
) increase by up to nearly
9%.
Fig. 8. Loss increase with back-iron extension
C. Thermal benefits of back-iron extension
From the foregoing discussions it is clear that the selection
of the BIE geometry is a multiphysics problem since on one
side the thermal path for the losses is improved, but on the
other hand, the additional copper and iron losses can increase
significantly.
The balance between these two aspects determines whether
BIE brings about any thermal benefits to the motor.
Combining the thermal tool of section A with the analysis of
section B in a single iterative design environment, where
thermal conductivities (and losses) are adjusted with
temperature until a defined convergence is achieved, Fig. 9 is
obtained. In Fig. 9 (a) the peak winding temperature at the
rated torque is shown for the full sweep of BIE geometries
considered, with the line at 173°C representing the peak
temperature for the original machine. Fig. 9 (b) shows the net
peak temperature difference as a percentage (%) with respect
to the original machine.
For the case of a very short BIE (BIED1), the increased
losses outweigh the thermal benefit. Thermal improvements
are observed for BIED2 and BIED3, especially when the BIE
width is small, with winding temperature reductions of up to
3% and 8% respectively. For deeper back iron extensions,
BIED4 and BIED5, representing radial BIE projections of
80% and full slot depth, the reduced slot thermal resistance is
in largely dominant over the increasing losses and improved
thermal performance is observed for all the investigated BIE
widths. Peak winding temperature reductions of up to 14%
and 19.3% are observed when the BIE width is around 1.6 mm
and 2.2mm respectively. It is worth noting that peak winding
temperature is more sensitive to BIE width value up to 0.7
mm, with the peak winding temperature dropping by down to
8 °C in this region. When BIE width increases beyond the
aforementioned values no significant temperature reduction is
observed. Table 3 summarizes the optimized BIE width for the
different BIED considered, together with the corresponding
peak winding temperature reduction.
0 1 2 3 4
140
150
160
170
180
Peak Temperature (
o
C)
Back-iron Extension Width (mm)
BIED1
BIED2
BIED3
BIED4
BIED5
(a)
0 1 2 3 4
-10
0
10
20
30
Temperature Reduction (%)
Back-iron Extension Width (mm)
BIED1
BIED2
BIED3
BIED4
BIED5
(b)
Fig. 9 (a) Peak winding temperature (b) Peak winding temperature
variation with BIE dimensions for rated torque condition
TABLE
3
OPTIMIZED BIE WIDTH AND IMPROVED WINDING TEMPERATURE
Simulations
Optimized BIE
width
Temperature
reduction
Winding
temperature
reduction
percentage
mm
0
C %
BIED2 0.7 2.28 2.68
BIED3 1 6.45 8
BIED4 2 12 14
BIED5 2.2 16 19.3
The BIE reduces the peak winding temperature mainly by
shortening the heat transfer path between the hot spot in the
slot and the coolant. The temperature distributions of half the
slot in the original motor and in the motor with the optimized
BIE for the same output torque are displayed in Fig. 10. The
hot-spot (173°C) inside the slot of the original motor (i.e.
without BIE) is located at the left hand side (i.e. centre of slot)
as shown in Fig. 10 (a). For the motor with the optimized BIE
the hotspot temperature is reduced down to 156°C, and the
hotspot location is shifted to the centre of Fig. 10 (b) (i.e.to the
mid-position of the half slot).
In light of the analysis performed, an optimized BIE width
of 2.2 mm for a BIE depth 21.7 mm is identified as the
optimal selection to pursue with for the experimental
validation. The copper fill factor with the proposed motor
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decreases by 11.28% from 0.4583 in the original motor to
0.4066, as shown in Fig. 4.
0 5 10
5
10
15
20
166
138
153
137
150
170
160
Slot depth (mm)
Slot width (mm)
120
130
140
150
160
170
Temperature
168
173 163
172 169 163 152 137
162 160 155 146 133
139 138 135 130 122
(a)
0 5 10
5
10
15
20
139148
153
155
153
137
146
150
152
151
128
136
141
146
120
130
140
150
160
170
Temperature
Slot depth (mm)
Slot width (mm)
145
153 156 155 149 13 9
148 152 151 147 138
137 139 138 136 130
(b)
Fig. 10. Temperature contour (a) original motor. (b) motor with BIE
D. Generalization aspects
The previous sections have shown the benefit of the
proposed BIE. To generalize the relationship of back-iron
extension geometry with motor slot geometry with the aim of
achieving the best thermal performance, a series of scaled
motors based on the studied traction motor are analyzed. In
this exercise the dimensions of the machine are scaled from
1/4
th
to 4 times the dimensions of the original motor. The ratio
of the back-iron extension width 's
1
' to the slot bottom width
's
2
' is plotted versus the slot bottom width in Fig. 11,
considering the back-iron extension depth 'd
1
' to be always
equal the slot depth 'd
2
'. From Fig. 11, it can be seen that the
optimal ratio of 's
1
/s
2
' increases in an almost linear
relationship with the slot bottom width when the slot bottom
width is small, increasing from 6.2 % to 11% when the slot
width increases from 10mm to 21mm. As the slot bottom
width increases beyond 40mm, the optimal ratio of 's
1
/s
2
'
flattens out to around 12-13%.
Slot bottom width S
2
Slot depth d
2
s
1
d
1
(a)
0 20 40 60 80 100
9
10
11
12
13
BIE Width S
1
/Slot Bottom Width S
2
(%)
Slot Bottom Width S
2
(mm)
Fig. 11. Optimized back-iron extension geometry
III. E
XPERIMENTAL VALIDATION
A. Experimental set up
Two stator segments with a water-cooled housing are
designed and built purposely for the back-iron extension
thermal benefit validation. One segment represents the
original stator configuration, while the other segment features
the optimized back-iron extension dimension. The segments
are of the same length as with the original traction motor. Iron
losses are neglected in the validation process as they are small
compared to the copper losses generated in the slot. Therefore,
to simplify the manufacturing procedure, both segments are
made from solid steel.
As discussed in section II, the introduction of the back-iron
extension in the slot reduces the slot area available for
winding, thus resulting in a lower copper fill factor and higher
copper losses. The number of strands per turn is thus reduced
accordingly for the segment with the BIE.
A plurality of thermocouples (TCs) are placed in the
manufactured segments to accurately characterize and
compare the thermal performance. These are placed in the slot
(TC1-TC4), end-winding and stator back-iron (TC5) as shown
in Fig. 12. For the stator segment representing the original
stator, there are three TCs in each slot, with TC3 placed in the
location where the hotspot is expected. TC1 and TC2 are
placed between the slot liner and the winding. On the other
hand, in the segment with the back-iron extension, there are
four TCs, and in this case TC4 corresponds to the location
where the hotspot is expected. TC3, similarly to TC1 and TC2,
is located between the winding and slot liner, which is used
for ground insulation.
7
IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS
25%
50%
TC1
TC2
TC3
TC5
(a)
25%
50%
TC1
TC4
TC2
TC3
TC5
(b)
Fig. 12. Layout and thermocouple locations (a) original stator
segment. (b) stator segment with back iron extension
The segments are fed by a DC current in order to generate
heat in the windings. The segments are thermally insulated
from air hence the heat generated in the segments can be
considered as being transferred to the coolant in the housing
and pumped out to the water chiller. The experimental setup is
shown in Fig. 13, including the segment under test, DC power
supply, temperature logger and water chiller.
Segment
Thermocouple
Flow meter
Chiller
Pico logger
PC
Power supply
Segment
Thermal insulation
Housing
Fig. 13. Experimental setup
B. Thermal experiment results
As discussed in section II, the back-iron extension reduces
the winding temperature by improving the problem of low
equivalent thermal conductivity inside the slot. The equivalent
thermal conductivity inside the slot affects the degree of
thermal benefits the back-iron extension can bring. Varnish,
the thermal conductivity of which is around 0.2 – 0.3 W/(mK),
depending on the grade, plays an important role in
determining the equivalent thermal conductivity inside the
slot. Therefore, the two segments were varnished by vacuum
pressure impregnation (VPI) which ensures a good fill quality
and consistency by the impregnation resin.
Fig. 14 shows the power/losses as a function of the input
current for the varnished segments. As shown in this figure,
the power output from the power supply is higher than the heat
generated in the segments, as there are losses in the connection
wires and wire connection junctions between the power supply
and the segments. Furthermore, from Fig. 14, it is evident that
the decreased copper fill factor due to the space occupied by
the back-iron extension results into a higher electrical
resistance in the segment with BIE and thus results in a higher
losses for the same current.
Fig. 14. Power comparison in the varnished segments
Fig. 15 shows the measured and predicted peak
temperatures for the varnished segments as a function of the
input current. The peak winding temperature as predicted in
LPTN modelling and from the experimental results of each
segment for different currents are plotted and compared. The
experimental tests agree well with the predictions, with
deviations which can be attributed to things such as
manufacturing processes and precise TC placement.
From Fig. 15, as expected, it is noted that the peak
temperature for both stator segments increases with current,
however the rate of rise varies between the two, with the one
featuring the BIE slowing down the rise and resulting into a
lower increasing rate.
For the segment representing the original stator, an input
current of 30A translates to a hotspot of 180
°
C, while for the
8
IEEE TRANSACTIONS ON INDUSTRIAL ELECTRONICS
varnished BIE segment, for the same current the peak
temperature is 132
°
C or around 26.7% lower.
Alternatively, for the same peak temperature of 180
°
C, the
current in the varnished BIE segment can be increased from
30.4A (8.87 A/mm
2
) to 36.5A (11.71 A/mm
2
).
Fig. 15 Back-iron extension thermal benefits
Fig. 15 shows significant benefits resulting from the
adaptation of BIE. It is worth noting that the BIE benefit is
even more significant if the equivalent thermal conductivity in
the slot is small, since the equivalent thermal resistance inside
the slot occupies a large percentage of the total equivalent
thermal resistance between hot spot and coolant.
IV.
CONCLUSION
For next generation electrical machines with a step change
in performance metrics, the improvement in thermal
management is one of the most important enablers in pushing
operational boundaries. Often improvements in thermal
management are obtained at the cost of more intensive
cooling, such as by flooding the stator or using costly high
thermal conductivity materials. This paper proposes a simple,
novel way to improve the motor thermal performance with
concentrated winding by extending part of the back-iron in the
slot, shortening the heat removal path between the hot spot in
the slot and coolant. Such a modification is low-cost to
implement as it doesn’t involve any new additional material,
and can be achieved by the appropriate design of the
lamination punching tool in a motor volume manufacturing
environment. It translates to marked thermal benefits,
especially in electrical machines with wider slot, such as
concentrated-wound machines. Moreover, it is also applicable
to other types of machines and/or winding configurations. For
single layer concentrated windings, where the coil section is
usually thicker, the thermal benefit could be even more
significant with respect to the double-layer case presented in
this paper. For traditional distributed winding, implementing
the proposed method needs careful consideration. The impact
on the copper fill factor and end-winding length is significant.
In addition any resulting benefit is likely to be limited as
distributed windings naturally have a higher number of slots
per pole which reduces the issues with winding hotspots.
Taking an existing electric vehicle traction motor, combined
thermal and electromagnetic analysis of different back-iron
extension geometries has been presented, and it has been
experimentally shown that with the appropriate multi-domain
optimization a significant 26.7 % peak winding temperature
reduction can be achieved with the back-iron extension. The
simplicity of the presented solution combined with its
technical effectiveness make it a strongly enabling solution for
next generation improved performance motors.
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Fengyu Zhang received B.E degree in
thermal engineering from Huazhong University
of Science and Technology, Wuhan, China in
2014. She is currently pursuing a Ph.D degree
at University of Nottingham with a focus on
thermal management of electrical machines.
Her main research interests include high
performance motors for transport applications
and their multi-domain optimization.
David Gerada received the Ph.D. degree in
high-speed electrical machines from University
of Nottingham, Nottingham, U.K., in 2012.
From 2007 to 2016, he was with the R&D
Department at Cummins, Stamford, U.K., first
as an Electromagnetic Design Engineer (2007–
2012), and then as a Senior Electromagnetic
Design Engineer and Innovation Leader (2012–
2016). At Cummins, he pioneered the design
and development of high-speed electrical
machines, transforming a challenging
technology into a reliable one suitable for the transportation market,
while establishing industry-wide-used metrics for such machinery. In
2016, he joined the University of Nottingham as a Senior Fellow in
electrical machines, responsible for developing state-of-the-art
electrical machines for future transportation which push existing
technology boundaries, while propelling the new technologies to higher
technology readiness levels.
Dr. Gerada is a Chartered Engineer in the U.K. and a member of
the Institution of Engineering and Technology.
Zeyuan Xu received the Ph.D. degree in
mechanical engineering from the University of
Manchester, Manchester, U.K., in 2002.
He subsequently worked as a Research
Fellow at UMIST, Brunel University, and the
University of Nottingham. He is currently a
Senior Research Fellow in thermo-mechanical
design of high speed electrical machines within
the PEMC group at the University of
Nottingham, Nottingham, U.K. His main
research interests include turbulent thermo-fluid flow, heat transfer
enhancement, and thermal management of advanced electrical
machines and power electronics.
Xiaochen Zhang (S’09-M’12) graduated from
Harbin University of Science and Technology and
received Master’s Degree in 2006. He graduated
from Harbin Institute of Electrical Technology and
received Doctor’s Degree in 2012.
He is with the Department of Electric and
Electronic Engineering, The University of
Nottingham, Nottingham, UK. His research
interests include research on electromagnetic and
thermal analysis on electrical machine, especially
in permanent magnetic machines and high speed
machines.
Chris Tighe received the MEng and Ph.D.
degrees in mechanical engineering from the
University of Nottingham, U.K., in 2007 and 2011,
respectively. After graduating, he spent time
working in the electrical generator industry and in
various research and commercial machine
development positions at the University of
Nottingham. He is now the proprietor of Electrical
Cooling Solutions, an engineering design consultancy specialising in
the thermal management of electrical machines and power electronics.
He Zhang received his B.Eng. degree from
Zhejiang University, China, in 2002. He
obtained the MSc. and Ph.D. degree in
electrical machines from The University of
Nottingham, UK, in 2004 and 2009 respectively.
After this he worked as Research Fellow at the
University and Director of BestMotion
Technology Centre. He moved to University of
Nottingham Ningbo China as Senior Research
Fellow in 2014 and Principal Research Fellow in
2016. Currently he is the Director of Nottingham Electrification Centre
(NEC) within the Power electronics, Machines and Control research
group in University Of Nottingham. His research interests include high
performance electric machines and drives for transport electrification.
Chris Gerada (M’05-SM’12) received the
Ph.D. degree in numerical modelling of
electrical machines from The University of
Nottingham, Nottingham, U.K., in 2005.
He subsequently worked as a Researcher
with The University of Nottingham on high-
performance electrical drives and on the
design and modelling of electromagnetic
actuators for aerospace applications. In 2008,
he was appointed as a Lecturer in electrical
machines; in 2011, as an Associate Professor; and in 2013, as a
Professor at The University of Nottingham. He was awarded a
Research Chair from the Royal Academy of Engineering in 2013 and
his main research interests include the design and modelling of high-
performance electric drives and machines.
Prof. Gerada serves as an Associate Editor for the IEEE
TRANSACTIONS ON INDUSTRY APPLICATIONS and is the past
Chair of the IEEE IES Electrical Machines Committee.
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