energies

Article
Thermo-Electro-Fluidic Simulation Study of Impact of
Blower Motor Heat on Performance of Peltier Cooler
for Protective Clothing
Kwon Joong Son

Department of Mechanical and Design Engineering, Hongik University, Seoul 30016, Republic of Korea;
kjson@hongik.ac.kr

Abstract: The necessity for portable cooling devices to prevent thermal-related diseases in workers
wearing protective clothing in hot outdoor weather conditions, such as COVID-19 quarantine sites,
is increasing. Coolers for such purposes require a compact design and low-power consumption
characteristics to maximize wearability and operating time. Therefore, a thermoelectric device based
on the Peltier effect has been widely used rather than a relatively bulky system based on a refrigeration
cycle accompanying the phase change of a refrigerant. Despite a number of previous experimental
and numerical studies on the Peltier cooling device, there remains much research to be conducted on
the effect and removal of motor-related internal heat sources deteriorating the cooling performance.
Specifically, this paper presents thermo-electro-fluidic simulations on the impact of heat from an air
blower on the coefficient of performance of a Peltier cooler. In addition, a numerical study on the
outcome of heat source removal is also evaluated and discussed to draw an improved design of the
cooler in terms of cooling capacity and coefficient of performance. The simulation results predicted
that the coefficient of performance could be raised by 10.6% due to the suppression of heat generation
from a blower motor. Accordingly, the cooling capacity of the specific Peltier cooler investigated in
this study was expected to be considerably improved by 80.6% from 4.68 W to 8.45 W through the
design change.

Keywords: thermoelectric cooling; Peltier effect; multiphysics simulation; computational fluid

Citation: Son, K.J. Thermo-Electro-
dynamics; conjugate heat transfer
Fluidic Simulation Study of Impact of
Blower Motor Heat on Performance
of Peltier Cooler for Protective
Clothing. Energies 2023, 16, 4052. 1. Introduction
https://doi.org/10.3390/en16104052 There exists a great risk of heat-related diseases when working in hot outdoor working
Academic Editor: Mahmoud conditions, such as at COVID-19 quarantine sites where wearing protective clothing is
Bourouis mandatory [1]. This is because protective clothing acts as a barrier to prevent the convection
of inner air and body heat to the ambient environment. Therefore, a technical need for
Received: 12 April 2023 portable cooling devices suitable for protective clothing has emerged to enhance comfort
Revised: 6 May 2023
while preventing the risk of heat-related illness, such as heat syncope, in hot weather.
Accepted: 10 May 2023
Furthermore, such portable cooling systems should be lightweight, compact, and low-
Published: 12 May 2023
power-consuming for long-lasting serviceability without interfering with quarantine tasks
and charging during shift hours.
Depending on the energy source, protective clothing cooling devices can be classified
Copyright: © 2023 by the author.
into passive systems that use cold heat sources, such as ice packs, and active systems that
Licensee MDPI, Basel, Switzerland. use refrigeration cycles or thermoelectric cooling. Passive systems bear disadvantages
This article is an open access article such as difficulties regarding temperature control and air circulation. In addition, there
distributed under the terms and is a risk of frostbite and poor comfort due to dew formation from cooling sources at very
conditions of the Creative Commons cold temperatures compared to body temperature [2,3]. The active system can control
Attribution (CC BY) license (https:// the temperature and flow rate of air during its operation. However, there may be the
creativecommons.org/licenses/by/ disadvantage of poor portability because high-density components such as batteries and
4.0/). motors must be equipped. In particular, refrigeration cycle-based coolers are unsuitable for

Energies 2023, 16, 4052. https://doi.org/10.3390/en16104052 https://www.mdpi.com/journal/energies

Energies 2023, 16, 4052 2 of 16

portable use in protective clothing since they must be equipped with pumps, compressors,
and evaporators for refrigerant circulation, evaporation, and condensation [4,5]. On the
other hand, thermoelectric cooling does not require complex refrigerant processes, thereby
achieving miniaturization and weight reduction of the portable cooler [6,7].
Portability significantly differs depending on the cooling range and capacity, even in
the same thermoelectric cooling method. A full-body cooling system has been developed
to allow cooled air to circulate through the upper chest and back areas of the protective
suits [6]. This cooling device also has a dehumidification function against moisture caused
by sweat inside the protective clothing. Despite adopting such advanced technology, the
system is heavy weighing 1.2 kg and equipped with two 80 mm-diameter circulation
fans, making it difficult for workers to wear it for long periods of time. A half-body
cooling system that supplies cold air only to the upper body has also been developed. The
weight of the system, excluding the blower fans and batteries, is 450 g, showing a dramatic
reduction in weight [8]. This paper introduces the design and performance of an even
more lightweight Peltier cooler than the two devices mentioned above. The cooler, with a
compact structure and weighing only 279 g, can blow cooled air into the protective clothing
without the requirement of additional fans in the other parts of the clothing.
Experimental studies, system modeling studies, and numerical analysis studies have
been conducted complementarily to design and evaluate the performance of portable
Peltier coolers. Experimental studies aim to measure cooler prototypes’ performance,
identify technical issues, and draw design improvement plans where necessary. The
testing apparatus for Peltier coolers has been designed and constructed to measure air
flow rate and temperature, surface temperature distribution, fan rotation speed, and power
consumption by using various sensors and data acquisition devices such as anemometers,
thermocouples, thermographic cameras, electrical power sensors, motor encoders, and
so forth [9–12]. For example, experimental studies have been conducted to determine
the performance characteristics of the cooler by measuring the coefficient of performance
(COP), a widely used figure of merit for thermoelectric cooling systems, against the blower
fan speed or by measuring the electric consumption versus COP relation [11,12].
The system modeling approach can derive an initial design that fulfills the performance
requirements from the experimental models or rules of thumb of various components
constituting the Peltier cooler. For example, the performance curve of the blower fan [13],
the heat pumping characteristic of the Peltier element [14], and the heat exchange model of
the cooling fin [13] can be utilized in this method. Model-based design strategy is beneficial
at the initial design and prototyping stages before the detailed development process through
numerical and computational approaches. This can help component selection and system
construction by predicting the device’s cooling performance approximately [15].
When developing thermoelectric cooling devices, computer-aided engineering (CAE)
methods have been widely exploited complementarily with experimental or system model-
ing approaches because computer simulations can verify the design rapidly and reliably
before the prototyping and testing stages. The Peltier cooler operates inherently by cou-
pling thermoelectric effects, fluid flow, and heat transfer. Therefore, a numerical method
for it should have multiphysics characteristics [16,17]. Numerical modeling and analyses of
Peltier coolers have been generally performed three-dimensionally rather than simplifying
to lower dimensions due to the complex geometries of thermoelectric elements, heat ex-
changing fins, and fan blades. In terms of time dependence, both steady-state and transient
analyses have been of great interest in this simulation field. This is because the prediction
of both steady-state response and initial dynamic behavior is vital for the performance
evaluation of a Peltier cooler, which takes a relatively long time of over 200 s to reach the
steady state [18,19]. In the airflow analysis using computational fluid dynamics (CFD), tur-
bulence models have been popularly used instead of laminar models because the associated
Reynolds number is very high due to fast impeller speed and narrow cooling fin chan-
nels [8,9,13]. Comparing three widely used turbulent models, i.e., the standard k − e model,
the realizable k − e model, and the shear stress transport (SST) k − ω model, the SST k − ω
Energies 2023, 16, 4052 3 of 16

model showed the highest accuracy in a three-dimensional steady-state CFD analysis for a
home refrigerator with an integrated Peltier cooling unit [19]. As a material constitutive
relation, linear thermoelectric equations have been most widely used [20,21]. Their mate-
rial coefficients are usually treated as constants, but temperature-dependent nonlinearity
becomes non-negligible if the operating temperature difference is significant [22,23].
Three-dimensional steady-state CFD analysis of thermoelectric air cooling chamber
with a liquid heat exchanger on the high-temperature side was carried out based on the
realizable k − e turbulent model [24]. Design verification of a portable Peltier air cooler
integrated with the work jacket was performed with CFD simulations [8]. Transient CFD
and thermoelectric analyses were conducted in a one-way coupling fashion [25]. While
performing transient CFD analysis based on the k − ω turbulent model and transient
thermoelectric analysis in order, the average wall temperatures on both TEC sides obtained
from the CFD analysis were reflected in the thermoelectric model computation as boundary
conditions. Although the blower fan geometry was included in the CFD analysis in the
above-mentioned simulation studies, any heat transfer through the blower surface was
neglected by treating it as an insulation boundary condition. To the best of the author’s
knowledge, computational research on the effects of fan motor heat on the performance
degradation of thermoelectric coolers (TECs) for protective clothing has yet to be published,
even though the amount of heat generation is considerable.
This paper introduces a numerical study on the effect of heat from a blower motor on
the performance of a portable Peltier cooler for protective clothing. A three-dimensional,
steady-state, coupled-field analysis based on an incompressible turbulent model and
linear thermoelectric constitutive relations is performed for multiphysics phenomena in
which turbulent airflow, thermoelectric cooling, and heat transfer through conduction
and convection co-occur. The simulation results for the case incorporating the motor heat
were verified by comparing them with the experimental results. In addition, a simulation
without the internal heat source was carried out to quantify the improvement of cooling
capacity and efficiency.
The rest of this paper is structured as follows. The consecutive Sections 2 and 3
introduce the design characteristics of the proposed Peltier cooler and its performance
testing results, respectively. Section 4 presents the basic numerical models and analysis
setups for the simulation of the Peltier cooler. Section 5 shows the simulation results with
experimental observations and discusses the effect of an internal heat source on the overall
cooling performance. Lastly, Section 6 presents the conclusions of this computational study.

2. Peltier Cooler Design

According to the opinions of the medical staff at COVID-19 quarantine sites, the main
technical specifications of the cooler required for outdoor tasks in hot weather conditions
can be summarized into four main items. First, the cooler for protective clothing should
prevent potential heat diseases by supplying conditioned air at least 5 ◦ C lower than the
outside temperature in hot weather. Second, the cooler should operate continuously for at
least two hours before the work shift without battery replacement or charging. Third, the
cooler structure should be simple and compact, less than 10 cm in thickness, to facilitate
its attachment and detachment to and from protective clothing as well as to avoid any
interferences with tasks during operation. Fourth, the system’s total weight, excluding the
battery, should be less than 300 g so as to mitigate muscle fatigue in outdoor workers.
A portable Peltier cooler was designed to satisfy the four requirements mentioned
above, and its prototype was built for immediate use in the field. Figure 1 shows the CAD
image of the Peltier cooler (Cyro™ designed and manufactured by NK Innovation, Inc.,
Sejong, Korea) used in this computational study. The overall dimension of the cooler is
172 mm × 49 mm × 61 mm, and its weight is 279 g. The power consumption of the cooler
is 30 W, which can be used for more than 2 h with a 11.1 V 8700 mAh rechargeable battery
supplying 96.57 Wh at maximum. Structural materials should facilitate rapid prototyping
and enable effective thermal insulation between cold and hot air ducts. Therefore, a 3-D
Energies 2023, 16, 4052 4 of 16

printable photoactive polymer resin (Formlabs™ White Resin) with low thermal conduc-
tivity and lightweight was used as the primary structural material. The thermoelectric
element adopted for the cooling system is TEM TB-127-1.4-2.0 manufactured by Kyrotherm.
Heat sinks were attached to both sides of the Peltier element, and the insulation wall be-
tween them separated cold and hot air ducts. Heat sinks are made of aluminum alloy with
lightness and high thermal conductivity. Due to its high thermal conductivity and excellent
adhesion characteristics, TSE3941 silicone adhesive from Momentive Performance Materi-
als Inc. was used for heat sink attachment. In addition, each flow channel is equipped with
a blower fan to ventilate air and induce forced convection around the heat sink.
Air inlets
Air inlets

Cold air outlet

Blowers

Peler element
with heat sinks
Air inlet

Hot air outlet

Figure 1. Design of portable Pelter cooler Cyro™ for protective clothing (courtesy of NK Innova-
tion, Inc.).

Figure 2 shows the constructed prototype and an example of the protective clothing
with the Peltier cooler installed. The Peltier cooler design in this study is similar to that
of [9] in that the cooling and heating parts are separated for insulation but different in
that the airflows are in the same direction, not in the opposite direction. As shown in
Figure 2a, the device frame consists of three 3D-printed parts: the middle substrate and
two duct covers with the air inlet and outlet. Figure 2b shows an example of the Peltier
cooler installed inside a protective suit. The two holes are the hot air duct’s inlet and outlet.
The cooler and the battery pack are fixed by an inner chest harness, which accomplishes
structural and installation simplicity.

(a) (b)
Figure 2. (a) Constructed cooler prototype, (b) protective clothing with Peltier cooler attached to
inner chest harness (courtesy of NK Innovation, Inc.).
Energies 2023, 16, 4052 5 of 16

The blowers in the air ducts have an integrated structure of the motor and the impeller.
They inhale stagnant outer air in the direction of the impeller’s axis of rotation and blow
it toward the heat sink for heat exchange. In this process, the heat from the motor is
unwantedly transferred to the air, causing a preheating effect before reaching the heat sink.
As a result, the blower improves heat-exchanging efficiency through forced convection
at the low-temperature heat sink while also lowering cooling performance by the air-
preheating effect. As the motor output boosts to enhance blowing capacity, the degradation
in cooling efficiency due to preheating becomes more significant. Similarly, the preheating
effect from the hot side blower also deteriorates the heat dissipation capability at the
high-temperature heat sink. Therefore, this paper analyzes the impact of motor heat on
the thermoelectric cooling performance using multiphysics simulations and predicts the
potential performance improvement by removing such internal heat sources.

3. Performance Test Results

An experimental study was conducted to evaluate the performance of the Peltier
cooler prototype and obtain experimental data to validate the simulation results. Figure 3
shows the experimental apparatus for testing the cooler’s performance. Figure 3a,b present
an experimental setup for the cold and the other side duct, respectively. The apparatus
comprises a structural frame, various sensors, and a data collection board with a flat-panel
display. The main frame for fixing the cooling device and the sensing components is
the assembled structure of aluminum extrusion profiles. A pair of thermal anemometers
measured the velocity and temperature of the air in both the inlet and outlet. Although
not shown in Figure 3, the non-contact thermal camera module (Letpton 3.5 manufactured
by Teledyne FLIR LLC, Wilsonville, OR, USA) and a portable thermal camera (GTC-600C
supplied by Robert Bosch GmbH, Gerlingen, Germany) monitored the surface temperature
of the subject device. The opaque plastic cover was partially cut and replaced with a
transparent film, as shown in Figure 3a, to capture the surface temperature of the heat sink
and the blower inside the cold air duct with thermal cameras. Power digital wattmeters
(Model SEN0291 by Zhiwei Robotics Corp., Shanghai, China) were used to measure power
consumption by sensing the shunt voltage and the electric current for electric components
such as the thermoelectric element and the blower motors. A single-board computer
(Raspberry Pi 4 Model B made by Raspberry Pi Foundation, Cambridge, UK) processed the
wattmeter data to be displayed on the monitoring panel.

Pel"er cooler Aluminum frame Power sensor Thermal anemometer

LiPo ba!ery

LED display with SBC

(a) (b)
Figure 3. Experimental apparatus: configuration for performance test (a) on the cold duct side and
(b) on the hot duct side.

Coefficient of performance (COP) is a widely used performance figure for Peltier

coolers [11,12,26]. COP is defined by the ratio of cooling capacity Qc to total power
consumption Pt and can be expressed as

Qc Qc
COP = = (1)
Pt PP + Pb
Energies 2023, 16, 4052 6 of 16

where PP and Pb are the power consumption of the Peltier element and the blowers, respectively.
For the Peltier cooler of the design shown in Figure 1, the cooling capacity Qc repre-
sents the net thermal energy reduction rate in the cold air duct [24], so it can be expressed as

Qc = (ṁC p T )inlet − (ṁC p T )outlet (2)

where ṁ is the mass flow rate, C p is the heat capacity at constant pressure, and T is the
absolute temperature.
Figure 4 shows the temperature distribution plots for the portable Peltier cooler in
operation obtained by the infrared thermal imaging camera. The minimum temperature on
the upper surface of the heat sink in the cold air duct was measured as 20.8 ◦ C. Meanwhile,
the maximum temperature on the blower surface was 44.7 ◦ C and 47.8 ◦ C in the cold and
hot air ducts, respectively. These measured surface temperatures were used as the essential
boundary conditions imposed on the blower surfaces for heat transfer analysis.

(a) (b)
Figure 4. Color-mapped infrared thermal images (a) on the cold duct side and (b) on the hot duct side.

Table 1 shows the experimental results and the calculated performance indices. Each
measured value in this table is the mean of three data points in a steady state with a low
data fluctuation. The COP of the cooler, including the internal heat source, was evaluated
as 10%. The air temperature could be reduced by 1.9 ◦ C at most, falling short of the design
target of 5 ◦ C.

Table 1. Experimental results and calculated performance values.

Item Value
Ambient temperature 26.0 ◦ C
Power consumption (Peltier element) 22.67 W
Power consumption (cold duct blower) 2.91 W
Power consumption (hot duct blower) 3.77 W
Cold duct outlet velocity 4.06 m/s
Cold duct outlet temperature 24.10 ◦ C
Hot duct duct outlet velocity 4.05 m/s
Hot duct outlet temperature 36.06 ◦ C
Cold duct heat sink base temperature 20.8 ◦ C
Cold duct blower surface temperature 44.7 ◦ C
Hot duct blower surface temperature 47.8 ◦ C
Cooler capacity 1 5.11 W
COP 2 17.4%
1The cooling capacity Qc was obtained from Equation (2). 2 The coefficient of performance was calculated from
Equation (1).

4. Computational Model and Analysis Setup

This section presents the fundamental theories and simulation setup for carrying out a
three-dimensional steady-state thermo-electro-fluidic analysis of the portable Peltier cooler
introduced previously.
Energies 2023, 16, 4052 7 of 16

4.1. CFD and Thermoelectricity Model

The governing equations for the Peltier cooler analysis, which simultaneously incorpo-
rates the fluid flow, electrical current, and temperature fields, comprise the mass continuity
equation, the electrical continuity equation, the momentum conservation equation, and
the energy conservation law. Without describing those well-known fundamental gov-
erning equations, this section only introduces other basic theories for understanding the
coupled-field simulation approach.
The specific turbulence model used in this research is the shear stress transport
(SST) k − ω model, which combines the advantages of the k − e model and the k − ω
model [27,28].
The SST k − ω model consisting of two equations for the specific turbulent kinetic
energy k and its dissipation rate ω can be expressed in the steady-state form [29] as
" #
∂
∂ui ∗ ∂ ∂k
ρku j = τij − β ρωk + (µ + σk µt ) (3)
∂x j ∂x j ∂x j ∂x j
" #
∂
γ ∂ui 2 ∂ ∂ω 1 ∂k ∂ω
ρωu j = τij − βρω + (µ + σω1 µt ) + 2(1 − F1 )ρσω2 (4)
∂x j vt ∂x j ∂x j ∂x j ω ∂x j ∂x j

where ρ is the mass density, u is the fluid velocity component, τ is the Reynolds stress
tensor, µ is the dynamic viscosity, µt is the turbulent viscosity, and νt is the eddy viscosity.
In Equations (3) and (4), β∗ , β, σk , σω1 , σω2 and γ are constants computed based on two
sets of model constants, i.e., one for the Wilcox k − ω model and the standard k − e model
using a lever rule with the weight factor F1 . For those interested in further details on
the derivation of the two-equation SST k − ω turbulent model and the associated model
parameters, it is highly recommended to see the reference article [27].
The Prandtl number, defined as the kinematic viscosity to thermal diffusivity ratio,
can compute heat transfer through interfaces between air and heat sinks. In this study, the
Kays–Crawford turbulent Prandtl number model in the following form [30] was chosen for
the conjugate heat transfer (CHT) analysis of the Peltier cooler.
" !#−1
ρD
√
 
1 0.3 µt 0.3µt −
Prt = +√ − 1−e 0.3µt Prt∞
(5)
2Prt∞ Prt∞ ρD ρD

In the above Equation (5), Prt is the turbulent Prandtl number, Prt∞ is its far-wall value, and
D is the diffusion coefficient. Readers interested in further details of the Kays–Crawford
model used in this simulation are suggested to review the paper [31].
The heat flux and temperature associated with Peltier devices are mostly computed
analytically by the simple heat flux model or numerically by the finite element method
(FEM) with thermoelectric constitutive relation [21,26,32]. Thermoelectric constitutive
equations for the electric current density J and the heat flux q can be written as a function
of the electric field intensity E and the absolute temperature T in a vector form [33],
respectively, as

J = σ(E − α∇ T ) (6)
q = πE − κ ∇ T (7)

where σ is the electric conductivity, α is the Seebeck coefficient, ∇ is the spacial gradient
operator, π = αT is the Peltier coefficient, and κ is the thermal conductivity. For prob-
lems with a steep temperature gradient in the thermoelectric element, the temperature
dependency of the material constants in Equations (6) and (7) should be considered [22].

4.2. Simulation Setup

As the first step in the pre-processing, simulation domains subject to electro-thermo-
fluidic analysis were selectively extracted from the CAD model illustrated in Figure 1.
Energies 2023, 16, 4052 8 of 16

Among shape details, small-sized features were eliminated, such as small rounds or steps,
since they may cause poor computational efficiency due to unnecessary element refinement
without significantly affecting computational accuracy. Subsequently, the connectivity
between CAD components was checked and modified where necessary to extract the three-
dimensional volumes corresponding to the fluid domains so that each cavity space had a
well-defined continuous and closed surface. In total, two fluid domains corresponding to
cold and hot air ducts were created from the cavity of the CAD assembly model after small
feature removal and continuity check. After the fluid volume creation, solid parts, such as
the plastic covers and fan blades, not included in the computational analysis were removed.
Thin adhesive layers between the thermoelectric part and the heat sinks were modeled as
shell elements, not three-dimensional elements. The thickness was set to 500 µm to calculate
the adhesive layer’s thermal resistance and temperature gradient using the shell element.
Figure 5 depicts the finalized computational domain through the above-mentioned
process together with the mesh elements created by COMSOL Multiphysics software. To
clearly show the internal features of the computational domain, the geometry was treated
translucently, and the meshed elements were marked with lines only on the outer sur-
face. The fluid regions consist of 14,575,305 elements. In addition, the solid volumes
corresponding to the thermoelectric element and heat sinks were discretized into 865,592 el-
ements. For the silicon adhesive thin film, 6463 surface elements were allocated. The entire
computational domain has 3,533,910 associated nodes.

Figure 5. Simulation domains with meshed elements.

The thermoelectric element has a composite structure made of bismuth telluride

(Bi2 Te3 ), copper, and alumina, whose fundamental properties are shown in Table 2. Unlike
other materials, in the case of bismuth telluride, which has a relatively large temperature
gradient due to the Peltier effect, the temperature-dependent property was adopted as
in Table 3. This is the data look-up table contained in the material database of COMSOL
Multiphysics to account for such nonlinear material properties. Seebeck coefficient, electric
conductivity, and thermal conductivity can be computed using a cubic spline interpo-
lation scheme on the five data listed in Table 3. The physical properties of air for fluid
domains, aluminum for heat sinks, and silicone for an adhesive applied to the heat sink
and thermoelectric element interfaces are shown in Table 4.
The fluid analysis domain consists of two parts, i.e., cold and hot air ducts, according
to the cooler’s structure. Each duct has a blower fan for forced convection, and the fan
motor operates at a high speed of about 9000 rpm, which generates rotational turbulence.
The rotational flow simulation itself, which incurs a very high computing cost, is beyond the
interest and scope of this study. Instead, for computational efficiency, a flow field under the
same volume flow rate without consideration of the fan motion was computed to account
for the forced convection effect around the heat sink. In this flow analysis process, the inlet
Energies 2023, 16, 4052 9 of 16

velocity condition was derived from the measured volume flow rate at each duct, and the
outlet condition was set as the atmospheric pressure. The remaining boundaries, except
the inlet and outlet surfaces, were all treated as a no-slip wall. The fan motor heat was
assumed to outflux through the core cylindrical surface, excluding blades. Although the
blower input power can be measured with the experimental apparatus shown in Figure 3,
the heat flux cannot be estimated through accurate motor efficiency measurement as in [34].
The internal heat source effect was reflected in the simulation by assigning the measured
surface temperature of the blower to the relevant boundaries instead of imposing the
motor’s energy dissipation rate. All boundary conditions applied to the fluid domain
described above are summarized in Table 5.

Table 2. Properties of constituting materials thermoelectric element (from material database of

COMSOL Multiphysics).

Heat Thermal Electrical Seeback

Density
Material Capacity Conductivity Conductivity Coefficiency
(kg/m3 )
(J/kg·K) (W/m·K) (S/m) (V/K)
Bi2 Te3 7700 154 κ (T ) 1 σ(T ) 1 α( T ) 1
Copper 8960 385 400 5.998 × 107 -
Alumina 3900 900 27 - -
1 Refer to Table 3 for temperature-dependent material data.

Table 3. Temperature-dependent properties of Bi2 Te3 (from material database of COMSOL Multiphysics).

Temperature (K) κ( T ) (W/m·K) σ ( T ) (×105 S/m) α( T ) (×10−4 V/K)

200 2.4 1.4286 1.68
250 1.9 1.1111 1.92
300 1.6 0.86957 2.10
350 1.6 0.71429 2.25
400 1.75 0.58824 2.37

Table 4. Material properties of air and structural components (from material database of COM-
SOL Multiphysics).

Thermal
Heat Capacity
Material Density (kg/m3 ) Conductivity Viscosity (m2 /s)
(J/kg·K)
(W/m·K)
Air 1.184 1007 0.02551 1.563 × 10−5
Aluminium 2700 900 238 -
Silicone 1650 1460 0.83 -

This study adopted a partial coupling approach for selected energy domains in some
computation stages instead of a full coupling scheme. The partially coupled solver is
highly advantageous regarding computing time, although the calculation accuracy is
slightly less than that of the fully coupled solver in a tolerable range [23]. Specifically,
before thermoelectric and conjugate heat transfer analyses, turbulent flow and electric
current analyses were carried out to obtain the velocity field in the fluid domain and
the current density field in the electric domain, respectively. Then, the computed field
data were transferred to the subsequent computing steps in a one-way coupling fashion.
Based on the shared flow and electric field data, in the remaining steps, the temperature
field was calculated in a fully coupled manner which simultaneously incorporates the
temperature-dependent Peltier effect, Joule heating, and heat transfer through conduction
and convection.
Energies 2023, 16, 4052 10 of 16

Table 5. Boundary conditions for simulation study.

Boundary Condition Value Remarks

Cold duct inlet velocity 3.94 m/s Derived from measured 1
Cold duct inlet temperature 26.03 ◦ C Measured
Cold duct blower temperature 44.7 ◦ C Measured
Hot duct inlet velocity 4.05 m/s Derived from measured 1
Hot duct inlet temperature 26.03 ◦ C Measured
Hot duct blower temperature 47.8 ◦ C Measured
Input voltage 10.38 V Measured
1 The inlet velocity was calculated from the continuity equation with incompressibility assumption.

5. Simulation Results and Discussion

Numerical simulations to evaluate and improve the thermoelectric cooling perfor-
mance of the portable Peltier cooler shown in Figure 2 were conducted in a coupled fashion
using the commercial software package COMSOL Multiphysics with its extension modules,
including CFD Module, AC/DC Module, and Heat Transfer Module.
According to the presence of heat sources in the air ducts due to the fan motor
operation, numerical simulation was conducted in two cases: one incorporating the internal
heat and the other suppressing the motor heat. Each simulation was performed through
three stages: turbulent CFD analysis, electric field analysis, and conjugate heat transfer
analysis accompanying thermoelectric cooling and Joule heating effects. As a result, such a
multiphysics simulation yielded the flow and pressure fields of the air, the potential and
current density fields of the electric parts, and temperature fields throughout the entire
computational domain.
Figure 6 shows the fluid velocity fields visualized as vector arrows in the cold and
hot air ducts. The cold duct has a maximum speed of 27.32 m/s in the entire domain and
an average normal velocity of 6.28 m/s through the middle cross-section of the heat sink.
Meanwhile, the highest fluid speed in the hot air duct and the average cross-sectional
velocity in the heat sink was calculated as 19.63 m/s and 7.60 m/s, respectively.

Velocity (m/s) Velocity (m/s)

(a) (b)

Figure 6. Simulation results: air velocity fields (a) in cold air duct and (b) hot air duct.

Figure 7 shows the simulation results in the electric domains, i.e., the copper wiring
layers and semiconductor segments. Figure 7a displays the computed electric potential
distribution by applying 10.38 V to the input port. Since the junctions of p-type and n-type
semiconductors interconnected by copper conductors occur repeatedly in space, a uniform
voltage drop appears across each junction. Figure 7b illustrates the electrical field intensity
distribution with a color map and the current density with vector arrows at 30,000 different
Gauss quadrature points. The length of each arrow was set in proportion to the vector
magnitude of the corresponding current density. Due to the inherent characteristics, highly
conducting copper plates possess nearly zero electric field intensity but relatively high
current density. On the other hand, semiconducting bismuth telluride blocks have a higher
electric field intensity of 23.2 V/m and a relatively lower current density than conductors.
Energies 2023, 16, 4052 11 of 16

Electric potenal (V) Electric field intensity (V/m)

(a) (b)

Figure 7. Simulation results: (a) electric potential distribution and (b) electric field intensity distribu-
tion with current density vectors.

Figure 8 illustrates the post-processed temperature distribution plots for the two
simulation cases. The outer surfaces were set to be translucent to reveal the internal wall
temperature features clearly.

Temperature ( ) Temperature ( )

(a) (b)

Figure 8. Temperature distribution plots for the Peltier cooler (a) including the internal heat sources
and (b) with the heat generation suppressed.

To verify the reliability of these simulation results, Figure 9 compares the numeri-
cally computed temperature distribution with an experimentally obtained thermographic
image. It should be noted that the thermal image in Figure 9b displays the external wall
temperature, except for the internal wall temperature of selected portions replaced with
the transparent film that can be observed in Figure 3a. On the other hand, Figure 9a shows
the temperature distribution on the fluid boundaries with top surfaces virtually hidden. In
other words, in Figure 9, it is only valid to compare the temperature distribution in two
boundaries, i.e., the blower’s circular top plane and the heat sink surface. Direct compari-
son in temperature of the other surfaces is not compatible. The minimum temperature on
the heat sink surface was 20.8 ◦ C in the experiment, while it was computed as 4.8% lower
at 19.7 ◦ C in the simulation. Therefore, it was confirmed that the simulation results predict
the experimental results precisely within a 5% error.
Figure 10 depicts a bar chart comparing the average temperatures computed at three
locations of interest from the temperature fields numerically obtained for the two simulation
cases shown in Figure 8. The minimum heat sink surface temperature, the cold duct mean
outlet temperature, and the hot duct average outlet temperature were 19.7 ◦ C, 24.26 ◦ C, and
30.04 ◦ C, respectively, in the simulation accounting for the blower heat generation. In the
other simulation with no account for the heat generation effect, meanwhile, the minimum
heat sink surface temperature, the cold duct mean outlet temperature, and the hot duct
average outlet temperature were 17.6 ◦ C, 22.86 ◦ C, and 35.4 ◦ C, respectively, which were
all lower than the simulation results considering blower heat.
Energies 2023, 16, 4052 12 of 16

Temperature ( Temperature (
44.7

20.3
(a) (b)

Figure 9. Temperature distribution comparison of (a) post-processed simulation image versus

(b) experimentally captured thermal image.

36.04 35.4
Simulation with internal heat
35
Simulation without internal heat
)
Temperature (

30

Ambient inlet temperature (26.03)

25 24.26
22.86

19.7
20
17.6

15
Cold duct heat sink Cold duct outlet Hot duct outlet

Figure 10. Temperature comparison chart at three locations according to the presence of internal heat
source in simulation.

Figure 11 depicts the color-contoured temperature distribution only in the electric

domain and the thermal energy transfer represented by vector arrows at 5000 different
Gauss quadrature points. The length of each arrow was set in proportion to the vector
magnitude of the heat transfer. The simulation results illustrate that the Peltier effect pumps
the thermal energy from the top low-temperature region to the bottom high-temperature
region so that the air can refrigerate in the upper chamber. In addition, this numerical
analysis confirmed that the Peltier effect becomes more active due to the removal of the
motor heat, which leads to further lower temperatures in the low-temperature side of the
thermoelectric cooling module.

Temperature ( ) Temperature ( )

(a) (b)

Figure 11. Temperature distribution plot in the electric domain due to the Peltier effect and cor-
responding heat flux vectors represented by arrows for two simulation cases (a) with account for
internal heat sources and (b) without motor heat generation.

Comprehensively summarizing the above-described simulation results, Table 6 lists

the experimental results and the calculated performance figures of merit. In this table,
the total power consumption Pt is obtained as in the denominator of Equation (1) by
Energies 2023, 16, 4052 13 of 16

adding Peltier power consumption PP obtained from the simulation and blower power
consumption Pb measured from the experiment.

Table 6. Performance values derived from simulation results.

Variable Internal Heat Included Internal Heat Excluded

Current 2.8128 A 2.8128 A
Cold duct outlet temperature 24.26 ◦ C 22.86 ◦ C
Heat sink base temperature 19.7 ◦ C 17.6 ◦ C
Hot duct outlet temperature 36.04 ◦ C 35.40 ◦ C
Peltier power consumption 29.20 W 29.20 W
Total power consumption 1 35.88 W 35.88 W
Cooling capacity 2 4.68 W 8.45 W
COP 3 13.0% 23.6%
1It was calculated by adding the simulated Peltier power consumption and measured blower power data. 2 The
cooling capacity Qc was obtained from Equation (2). 3 The coefficient of performance was calculated from
Equation (1).

Figure 12 comparatively shows the cooling capacity and COP derived from the perfor-
mance test and two simulation cases as a bar chart. Even considering that the simulation
results slightly underestimate the experimental results, it can be concluded from these
thermo-electro-fluidic simulations that removing the internal heat source can significantly
improve the cooling capacity by 80.6% from 4.68 W to 8.45 W, and accordingly, the COP by
10.6% from 13.0% to 23.6%.
10
8.45 25 23.6
Coefficient of performance (%)

8 80.6 % 10.6 %
Cooling capacity (W)

increase 20 increase
17.4
6
5.11 15
4.68 13
4
10

2 5

0 0
Experiment Simula"on Simula"on Experiment Simula"on Simula"on
with with without with with without
motor heat motor heat motor heat motor heat motor heat motor heat
(a) (b)

Figure 12. Bar charts for Performance figures of merit: (a) cooling capacity from Equation (2) and
(b) coefficient of performance from Equation (1).

Internal heat sources can be technically removed in a way that the electric motors
located outside the air ducts transmit the rotational power inward through the shafts.
Cyro™, the Peltier cooling device for protective clothing shown in Figure 2a, satisfies
three of the four technical requirements specified in Section 2, i.e., operating time, size,
and weight, except for temperature reduction by 5 ◦ C. The next version of Cryo™, with
improved cooling performance through design optimization and heat removal inside the
air duct, is currently under development through industry–academia cooperation.

6. Conclusions
The development of Peltier coolers for protective clothing to prevent the risk of heat-
related diseases in hot working environments has been actively conducted. In particular,
there has been great interest in energy-efficient cooling technology that can increase system
portability and duration time by reducing weight and power consumption. However, little
research effort has been exerted to investigate the effect of blower motor heat on overall
cooling efficiency. Therefore, this paper aims to numerically evaluate the effect of heat
from a blower motor on the performance of a portable Peltier cooler for protective clothing.
Due to the multi-physical characteristics of the thermoelectric coolers, in this study we
conducted the thermo-electro-fluidic analysis in a coupled fashion. Such a coupled-field
Energies 2023, 16, 4052 14 of 16

analysis can computationally evaluate the Peltier cooler’s performance indices, such as
cooling capacity and coefficient of performance, by incorporating turbulent flow analy-
sis, electric current density analysis, Peltier effect analysis, Joule heating analysis, and
conjugate heat transfer analysis. The reliability of the simulation results was validated
through comparison with the performance test results for the Peltier cooler prototype. The
simulation error against the measured data was 4.8% for the minimum heat sink surface
temperature in the cold air duct, guaranteeing simulation accuracy at a tolerable level. The
numerical analysis predicted that the cooling capacity and coefficient of performance could
be improved by 50.6% and 13.0%, respectively, by installing the heat-generating blower
motor outside the air ducts. The research findings of this study are differentiated in that
they quantitatively analyzed the effect of heat generation in blower motors on Peltier cooler
performance, which has not been covered in previous studies. The design and optimization
of a novel Peltier cooler resolving the motor heating issue could present a potential topic of
future research.

Funding: This work was supported by the International Science & Business Belt support program
through the Korea Innovation Foundation funded by the Ministry of Science and ICT (Grant No. 2021-
DD-SB-0288), and was also partially supported by 2023 Hongik University Research Fund.
Institutional Review Board Statement: Not applicable.
Informed Consent Statement: Not applicable.
Data Availability Statement: Not available.
Acknowledgments: The author acknowledges Jong-Bae Kim, CEO of NK Innovation Inc., for provid-
ing the design and prototype of the portable Peltier cooler Cyro™ as well as valuable technical advice.
Conflicts of Interest: The authors declare no conflict of interest.

Nomenclature
Cp heat capacity at constant pressure, J/kg·K
D diffusion coefficient, m2 /s
E electric field intensity, V/m
F1 k − ω model parameter weighting factor
J current density, A/m2
k specific turbulent kinetic energy, J/kg
ṁ mass flow rate, kg/s
Pb power consumption of blowers, W
PP power consumption of Peltier element, W
Pt total power consumption, W
Prt turbulent Prandtl number
Prt∞∞ far-wall turbulent Prandtl number
q heat flux, W/m2
Qc cooling capacity, W
T absolute temperature, K
u fluid velocity component, m/s
Creek symbols
α Seebeck coefficient, V/m
β k − ω model parameter
β∗ k − ω model parameter
γ k − ω model parameter
e specific turbulent energy dissipation rate in k − e model, J/kg·s
κ thermal conductivity, W/m·K
µ dynamic viscosity, Pa·s
µt turbulent viscosity, Pa·s
νt eddy viscosity, m2 /s
π Peltier coefficient, J/C
ρ mass density, kg/m3
Energies 2023, 16, 4052 15 of 16

σ electric conductivity, S/m

σk k − ω model parameter
σω1 k − ω model parameter
σω2 k − ω model parameter
τ Reynolds stress tensor, Pa
ω specific turbulent kinetic energy dissipation rate, J/kg·s
Abbreviations
CAD computer-aided design
CAE computer-aided engineering
CFD computational fluid dynamics
CHT conjugate heat transfer
COP coefficient of performance
FEM finite element method
rpm revolutions per minute
SST shear stress transport
TEC thermoelectric cooler

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