1

Thermal management of high temperature polymer electrolyte membrane fuel cell

stacks in the power range of 1 to 10kWe

E. Harikishan Reddya, b,

hariehkr@gmail.com

,

Sreenivas Jayantia, ∗sjayanti@iitm.ac.in

,

Dayadeep S. Monder b,

C, dmonder@iitb.ac.in

aDepartment of Chemical Engineering, IIT Madras, Chennai 600036, India

bDepartment of Chemical Engineering, IIT Hyderabad, Yeddumailaram 502205, India

CDepartment of Energy Science and Engineering, Indian Institute of Technology Bombay, Mumbai 400076, India

∗Corresponding author. Tel.: +91 44 22574168; fax: +91 44 22574152.

Originally published in Journal of Power Sources, Elsevier B.V.

http://dx.doi.org/10.1016/j.ijhydene.2014.09.132

Abstract

Maintaining optimal temperature of the stack is critical for efficient operation of high

temperature polymer electrolyte membrane fuel cells. While a number of possibilities of

thermal management exist for small stacks, the problem becomes more complicated for larger

stacks. In the present study, the thermal management of stacks in the power range of 1-

10 kWe is considered through computational fluid dynamics simulations. It is shown that

large stacks need to have dedicated cooling plates through which a coolant is circulated.

Further, stacks of the size of 10 kWe can have reasonably low cell temperature variations

(∼20 K) only by passing pre-heated liquid coolant through the coolant plates. Estimates show

that the concomitant increase in the coolant flow rate induces large pressure drops, of the

order of 30 bar, if a four-parallel serpentine is used on an active cell area of 30 cm × 30 cm. It

is therefore necessary to use parallel channel flow fields with carefully designed feeder

manifolds to maintain optimal cell temperatures and reasonably low coolant pressure drops in

large stacks.

Keywords: High temperature PEM fuel cells; Thermal management; Scale-up; Coolant

circulation system; Pressure drop; Computational fluid dynamics

2

Introduction

High temperature proton exchange membrane fuel cells (HT-PEMFCs) can be

operated up to a maximum temperature in the range of 180-200 °C, and thereby offer

advantages such as simplified water management due to the release of product water in

vapour form and high carbon monoxide (CO) tolerance, thereby allowing a wider choice of

fuels [1–4]. Due to their good dynamic response and high efficiency compared to internal

combustion engines, PEM fuel cells are well-suited to serve as power trains in vehicles such

as scooters and cars. Several aspects of HT-PEMFCs have been investigated in the past

couple of decades [5–16]. As a result, the principal features of HT-PEMFCs in terms of the

materials, processes and performance characteristics are well-understood at a cell level. A

recent review of advances in high temperature PEM fuel cells is given by Bose et al. [17].

One of the topics that has been receiving attention lately is the issue of thermal management

of HT-PEMFCs [18]. Currently available HT-PEM fuel cells offer lower current/power

densities than low temperature proton exchange membrane fuel cells (LT-PEMFCs) and

being able to maintain a uniformly high temperature in HT-PEMFC stacks is expected to

reduce this performance gap (see for example, [7,19]). A number of strategies are adopted for

thermal management of fuel cell stacks [20,21] and these have been critically reviewed

recently by Zhang and Kandlikar [22].

The cooling of the fuel cell stack can be achieved in a number of ways. These include

active cooling with air or liquid coolants, passive cooling with cooling fins and heat

spreaders, evaporative cooling and cooling with a separate air flow [22]. Passive cooling

methods have the limitation that they can be used only for very small stacks. Active cooling

methods, in which the coolant fluid is pumped through cooling passages within the stack, are

capable of greater heat removal from the stack and have received attention for cooling of LT-

PEMFCs [23–26]. Thermal management of PEMFC stacks requires two considerations,

namely, the initial heating up of the stack to the operating temperature and later maintaining

it at a constant operating temperature. Andreasen et al. [18] studied various cooling strategies

for an HT-PEMFC stack and concluded that the start-up time could be reduced by sending

pre-heated cathode air rather than using direct electrical heating. A number of stack thermal

management studies have been reported in the literature. These include the study of

Andreasen et al. [27] for an air-cooled HT-PEMFC stack designed for a hybrid electrical

vehicle; the studies of Scholta et al. [28,29] for a 5-cell, air/liquid cooled stack and a 10-cell

stack cooled using heat pipes; the study of Song et al. [30] for a natural circulation driven

water cooling system; and Kvesicetal and the work of Kvesic et al. [31,32] who developed a

multi-scale, three-dimensional model for a stack containing one coolant channel in each

bipolar plate. Kvesic et al. showed that with pre-heating of the reactants and coolant, the cell

3

temperature variation can be reduced to within the range of 3-6 K for a hydrogen-fed

stack [31] and to about 9 to 10 K for a stack fed with reformate gas [32]. This is in agreement

with the results of Scholta et al. [29] who reported a cell temperature variation of 56 K for an

inlet coolant temperature of 373 K. Modeling studies by the present authors of cooling of a

1 kWe stack by using cathode air [33] and a separate liquid coolant circuit [34] confirmed the

importance of coolant inlet temperature. The modeling results from the last study [34] are

broadly in agreement with the experimentally measured cell to cell variation of 6 to 8 K by

Supra et al. [14].

Most of the detailed studies reported hitherto have been performed for small stacks,

typically of 1 kWe or less. This is the power requirement of a scooter or a golf-

cart [35,36] and is far smaller than what is required for a motor car, where a fuel cell stack in

the range of 40 to 100 kWe [37] may be required. There is increasing interest in stacks of kW

range as evidenced by the recent studies of Jansen et al. [38] and Samsun et al. [4]. Samsun

et al. studied a 5 kWe HT-PEMFC stack designed as part of an auxiliary power unit using an

onboard reformer fueled by diesel and kerosene. The electrical power output from a stack can

be increased either by increasing the number of cells, or by increasing the active area of each

cell, or by using a modular construction of stacks of small size. Having a large number of

cells may demand more mechanical strength from the cell materials as significant

compression is required to ensure leak proofing. Also, the tie rods become longer and weight

of the other accessories increases. Increasing the active cell area poses thermal management

problems and uniform distribution of reactants may also become an appreciable problem.

One would envisage therefore a modular construction of stacks of reasonably small sizes. In

the present work, we consider issues related to the scaling up of a 1 kWe stack to about

10 kWe by using larger cells.

Some of the issues involved in the scale-up can be understood from the following

considerations. Increase in the stack size reduces the surface area (per unit volume) available

for heat transfer by an external air flow over the stack; the distance from the surface to the

core of the stack also increases. Our estimates of natural convective heat transfer show that,

due to these two factors, the fraction of total heat produced removed through the edges by

natural convection is reduced from 4.67% to 2.35% with an increase in active cell size from

10 × 10 cm (100 cm2) to 25 × 25 cm (625 cm

2). In the case of forced draft, an air velocity of

25 m/s (90 kmph) leads to a temperature difference of 50 K between the surface and the core

for a 10 × 10 cm cell. If the cell size is increased to 25 × 25 cm, the temperature difference

for the same air velocity increases to 120 K, i.e., parts of the cell along the outer edge will

operate at cell temperatures as low as 80 °C, which will be disastrous for a PBI-based

membrane. Therefore, operation of a HT-PEMFC stack with larger cells using forced draft

4

cooling alone is not feasible. A large stack needs to be cooled by passing a coolant through

the stack through specially-made passages. This brings in additional constraints of parasitic

power consumption for coolant circulation and the need for maintaining the coolant

temperature at optimal levels. Increase in the cell size from 10 × 10 cm to 20 × 20 cm

increases the total length of traverse of a single serpentine flow field by a factor of four and

may therefore quadruple the pressure drop for coolant circulation. The pressure drop for

coolant circulation also becomes important when cell temperature variations are to be kept as

low as possible, as will be shown later.

In the present work, we study these issues by evaluating the optimal conditions for

coolant circulation for 24-cell stacks of active cell areas of 10 × 10 cm, 20 × 20 cm and

30 × 30 cm, corresponding to a nominal stack power of 1, 4 and 9 kWe when the HT-PEMFC

operates at 0.5 V and 0.9 A/cm2 [32]. A stack model previously used for analysis of

integrated cathode air cooling [33] and external coolant circulation [34] is used to investigate

the pressure drop and cell temperature variations for the three cells with different cooling

strategies. Details of these studies and the results obtained are discussed below.

Description of the mathematical model

The mathematical model corresponds to a 24-cell stack, shown schematically

in Fig. 1, in which one coolant plate is provided for every four cells. Since the focus of the

present work is on thermal management and thus primarily on heat transfer, a simplified

model of the cell is used. The resolved features of the stack include six regions (see Fig. 2a).

Starting from the bottom of the figure, these are: 1) the coolant plate, 2) the anode monopolar

plate, 3) a thin low conductivity region consisting of anode gas diffusion layer (GDL), the

anode catalyst layer, and the membrane, 4) the cathode catalyst layer where the entirety of the

cell's heat generation is assumed to occur, 5) the cathode side GDL, and 6) the bipolar plate.

The complete details of the air flow channels in the bipolar plate and the coolant flow

channels in the cooling plate are also included in the model geometry. It may be noted that in

this model, the anode side flow field is not resolved and its role in the overall heat transfer is

assumed to be negligible [34]. The model shown in Fig. 2 accounts for one half of the self-

repeating unit of sandwich of layers centered around a typical coolant plate shown in Fig. 1.

Accordingly, only one-half of the cooling plate is included in the computational domain. The

length and breadth of the computational domain change with the size of the active area of the

cell; however, the thicknesses of the various layers are assumed to be constant. The

dimensions of each feature in the repeating unit for a 10 cm × 10 cm cell are also given

in Fig. 2a. In all cases, the height of the air channels and the coolant channels is maintained

constant at 1 mm.

5

Fig. 1 Schematic diagram of a 24-cell HT-PEM fuel cell stack with one cooling plate for

every four cells.

6

7

Fig. 2 Schematic diagram of (a) the resolved features of the computational domain, and (b)

boundary conditions.

In this computational domain, the velocity and pressure fields in the cathode side of

the bipolar plate and the flow of the coolant in the coolant channels are resolved by solving

the Navier-Stokes equations using well-established techniques of computational fluid

dynamics (CFD). In each case, the flow is assumed to be incompressible, and the

corresponding material properties, such as density and viscosity, are specified. There is one

inlet and one outlet for both fluids. At the inlets the average normal velocity into the

computational domain is specified and at the outlets, the flow is assumed to be fully

developed. A no-slip boundary condition is used at the walls. The temperature field in all the

layers is obtained by solving the heat conduction equation with appropriate source and

boundary conditions. In the cathode catalyst layers, the following local volumetric heat

generation term, including both the reversible and irreversible terms [33,34], is added as a

heat source to the heat balance equation:

Q = (−ΔH/2F−Vcell) i

Here, Q is the heat released per unit volume, ΔH is the enthalpy of water formation in

the gaseous phase, F is Faraday's constant, Vcell is the electric potential at which the fuel cell

is operated and i is the local current density.

The local current density is a function of a number of electrochemical parameters

which cannot be resolved in a stack-level model. In view of this, the dependence of the

current density on the local temperature of the cell and the operating potential is captured

using the following correlations [33] using the empirical data of Korsgaard et al. [7]:

at Vcell = 0.5 V

at Vcell = 0.6 V

where T is in K and i is in A/cm2.

Within the stack, the heat conduction in the solid layers is solved as a conjugate

problem with the convective flow of cathode air and/or coolant (as appropriate) assuming no

contact resistances at the interfaces of the various layers. The boundary conditions for the

heat transfer problem consist of specified inlet temperature at the inlet to the domain; the

fully developed flow boundary condition at the outlets; and a constant natural convective heat

transfer coefficient of 9 W/m2 K at all exposed walls with an ambient temperature of

300 K [34].

All the simulations have been done using the commercial CFD code ANSYS-

FLUENT. Second order accuracy was maintained in the discretization of the governing

8

equations. Hexahedral grid elements were used to discretize the domain. An optimal grid

density, previously established by comparing with analytical and experimental results from

the literature [34], has been used in the present study. Therefore, the number of grid elements

increased with increasing domain size; the total number of cells was in the range of 0.8 to 4

million. Based on previous studies [33,34,39], a four-parallel serpentine field was used for

the cathode side. For the coolant flow, three flow fields, namely, parallel, wide-parallel and

four-parallel serpentine flow fields have been used; these are shown schematically in Fig. 3.

A number of calculations, covering a range of air and coolant flow rates, inlet temperatures,

and geometrical dimensions of the cell, have been performed using the above computational

model. In each case, for a given cooling strategy, geometry, cell stoichiometry and inlet fluid

temperatures, the coolant flow rates were adjusted until the maximum temperature in the

model geometry reached 473 K. This would then constitute the thermal management strategy

for maintaining the given stack at as high a temperature as possible.

Fig. 3 Schematic diagram of the three flow fields used on the cooling plates.

Results and discussion

This work presents results for simulations of two approaches to stack cooling: 1)

integrated air cooling, and 2) a liquid coolant circuit. Both approaches use the geometry given

in Fig. 2. In the first approach, the air to be fed to the fuel cell stack is circulated through the

coolant channels and thus preheated before it is supplied to the fuel cell cathodes through the

air channels (see Fig. 2a). In the second approach, a liquid coolant is circulated through

9

separate cooling channels provided in the stack; therefore air is supplied to the fuel cells

directly.

Typical results of the temperature and current distributions predicted for integrated air

cooling and cooling using a liquid coolant circuit for a 10 cm × 10 cm active cell are shown

in Figs. 4 and 5 respectively. In these figures, the cathode flow field has a four-parallel

serpentine channel (FPSC) of height/depth of 1 mm and a width of 1 mm each with the

maximum length of the channel being 10 cm (Fig. 6). The total width of the FPSC unit is

24 mm and four such units will make up the total cell width of 10 cm. In case of larger cell

units, the length-wise direction (along the z-direction in Fig. 6) will increase to the cell length

and the width is made up by adding more number of these FPSC units. For example, in a

20 cm × 20 cm cell, eight FPSC units shown in Fig. 6 will make up the width of the cell with

each unit being 20 cm long. Each FPSC has a common inlet header and a common outlet

headed for the four parallel serpentine channels. The cooling plate may have parallel or FPSC

channels as shown inFig. 3.

Fig. 4 Integrated air cooling with the four-parallel serpentine channel (FPSC) flow field for

the bipolar plate and parallel channel flow field for the cooling plate: subplots (a–b) give the

spatial current density variation in the cathode catalyst layers CCL-1 and CCL-2; subplots (c–

d) show the temperature variation in CCL-1 and CCL-2 while subplots (e–f) give the

temperature field in the cathode flow fields CFF-1 and CFF-2; subplot (g) gives the

temperature distribution in the cooling plate.

10

Fig. 5 Same plots as in Fig. 4 but with a separate coolant circuit using a liquid coolant

(Duratherm 600) with an inlet temperature of 450 K, and a cathode air inlet temperature of

400 K (stoichiometric factor of 2).

Fig. 6 Schematic diagram of the four-parallel serpentine channel (FPSC) in the bipolar plate.

The results shown in Figs. 4 and 5 have been obtained for parallel channels in the

cooling plate. These figures show the contours of temperature and current density in the two

catalyst layers, the bipolar plates and the cooling plate for the cases of integrated cathode air

cooling (Fig. 4) and cooling with a liquid coolant (Fig. 5). It can be seen from Fig. 4 for the

integrated air cooling method that cathode air entering the cooling plate (from the bottom

11

in Fig. 4g) at 300 K, gets heated up quickly, and leaves the cooling plate at ∼410 K from the

top. This preheated air subsequently enters the cathode side of the first (Fig. 4e) and the

second bipolar plates (Fig. 4f) at the top left hand corner. The air again gets heated up

quickly, but as it flows through the bottom sections of the FPSC flow field, its temperature

drops due to the effect of the cool air entering the coolant plate. This effect is more prominent

in the first cathode bipolar plate (Fig. 4e) which is closer to the coolant plate than in the

second cathode bipolar plate (Fig. 4f). It can be seen from Fig. 4c and d that this cooling

effect also affects the temperature distribution in the cathode catalyst layers and that as a

consequence the current density (Fig. 4a and b) is lower in the bottom sections of the cells. It

may be noted that the color-map for 4g is different from the color-map for 4e and 4f, which

in turn is different from the one for 4c and 4d. This is done to highlight the details of the

temperature variations in each region.

Fig. 5 shows the corresponding contours for the liquid coolant case where the coolant

(Duratherm oil) enters the cooling plate at a temperature of 450 K at the top (Fig. 5g) and

flows down through parallel channels. While air enters the cathode channels at 300 K, it

quickly heats up and the temperature distribution (Fig. 5e and f) follows primarily that in the

cooling plate. As a result, the first bipolar plate and the first cathode catalyst layer have a

relatively lower temperature (Fig. 5c and d) and lower current density (Fig. 5a) than the

corresponding layers of second cell (Fig. 5b). These results demonstrate that the temperature

distribution in the cooling plate has a strong effect on the temperature and current density

distribution in the cells in the stack.

Thus, for each cooling strategy and flow condition, the temperature distributions in

the cathode catalyst layers, and the local current density can be evaluated thus giving the

stack-average current density. Simultaneously, the CFD calculations also give as output the

pressure drops on the cathode side and the coolant side. The variation of these quantities with

cell size and the implications on scaling up are discussed below.

Thermal management using cathode air cooling

With integrated cathode air cooling, it is possible to keep the cell temperatures below

200 °C by operating at a high stoichiometric factor of about 8 for a cell area of

10 cm × 10 cm [33]. These calculations have been repeated for cell sizes of 20 cm × 20 cm

and 30 × 30 cm by increasing the length of the flow domain appropriately. In the present case

where one cooling plate is used for every four cells, the second cathode catalyst layer is

hotter than the first catalyst layer because it is further from the cooling plate (see Figs. 4 and

5). Therefore the temperature contours on the second cathode catalyst layer are used to assess

the thermal management issues.

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Fig. 7 presents the predicted temperature and current density contours in the second

cathode catalyst layer for the three cell sizes. It may be noted that these images are

foreshortened to bring them to the same size and hence the distance should be seen in relative

terms, i.e., percentage of the total length or width. The temperature and current density values

are unaffected by this post-processing of computed results. It can be seen that the pattern of

temperature and current distribution is similar in all the three cases. This is attributable to the

temperature of the cooling fluid which enters through parallel channels located at the bottom

and exits at the top (see Fig. 4g). In this case of cathode air cooling, the air is already hot

(∼450 K) by the time it enters the FPSC channels of the bipolar plates at the top (Fig. 4e and

f). It then gets cooled and heated up again as it flows through the cathode channels depending

on the difference in the temperatures between the gas flowing in the cathode channels and

that flowing in the cooling plate channels. As a result, a fairly complicated evolution of the

gas temperature occurs in the flow channels (Fig. 4e and f) as well as in the cathode catalyst

layers (Fig. 4c and d). The same dynamics is evident in Fig. 7 for larger cells and the relative

temperature distribution on the second cathode catalyst layer is similar for all the three sizes.

The range of variation appears to increase as the cell size increases; this is a consequence of

the relative effectiveness of convective cooling and conduction across different layers and

over different lengths resulting in different stoichiometric air requirements to limit the

maximum temperature to 473 K.

Fig. 7 Current density distribution in CCL-2 with integrated cathode air cooling for a cell

dimension of (a) 10 cm, (b) 20 cm, (c) 30 cm, and the corresponding spatial variation in

temperature CCL-2 for a cell dimension of (d) 10 cm, (e) 20 cm, (f) 30 cm.

These are quantitatively summarized in Table 1 in terms of the average current

density and maximum temperature difference over the catalyst layer in each case. It can be

13

seen that as the cell size increases, the air stoichiometric factor rate required to keep the

maximum cell temperature below 473 K increases slightly. While the average current density

decreases slightly with increasing cell size, the minimum catalyst layer temperature shows a

more pronounced decrease with increasing cell size. This may lead to a cold spot on the cell

locally which may lead to a reduction in performance. For higher current densities (at a cell

operating voltage of 0.5 V), the drop in performance is more pronounced. For example, at a

cell voltage of 0.5 V, the average current density deviation from the ideal value (obtained

assuming that the entire cell is operating at 473 K) is about 11% for a cell size of 10 cm, 14%

for a cell size of 20 cm and 15% for a cell size of 30 cm. With increasing cell size, the air

flow requirements will also be higher and thus the cell pressure drop will increase, especially

for the high stoichiometric factors required (see Table 1).

Table 1 Integrated air cooling of stacks: key temperatures and average current density

obtained from CFD simulations. Maximum temperature on the second catalyst layer in all

cases is 473 K.

Dimensions of

the active cell

area

(mm × mm)

Cathode air

stoichiometric

factor

Cathode air

outlet

temperature (K)

Minimum

temperature (K)

on cathode

catalyst layer

Average

current

density

(A/cm2)

First

cell

Second

cell

First

cell

Second

cell

100 × 100 7.85 410 413 425 441 0.794

200 × 200 8.88 399 403 407 424 0.766

300 × 300 9.09 395 399 399 418 0.755

Thermal management using an external coolant circuit

Similar calculations for increased cell size have been performed for the separated

liquid coolant circuit case. As noted previously [34], the effectiveness of this configuration

depends significantly on the coolant inlet temperature. The predicted temperature and current

density contours in the second cathode catalyst layer for thermic oil coolant with an inlet

temperature of 450 K for cell sizes of 10 × 10, 20 × 20 and 30 × 30 cm are shown in Fig. 8.

Here, the coolant enters the stack (in a counter current direction with respect to the air) at a

temperature close to the stack operating temperature. Hence, the range of temperature

variation across the layer is rather small and is primarily determined by the cathode air

temperature. Due to the high coolant temperature and relatively low stoichiometric factor on

the cathode side, the air gets heated up quickly as it flows through the FPSC giving rise to a

small low temperature region at the entry. The extent of this region can be reduced by pre-

14

heating the cathode air. Increasing the stoichiometric factor for the cathode air increases the

maximum temperature variation within this layer [39]. Calculations show that these

temperature differences increase with increasing cell size and can reach as high a value

as ∼65 to 70 K for a 30 × 30 cm cell for a stoichiometric factor of three on the cathode side.

However, since this low temperature region is small, the overall effect on the average current

density is rather small.

Fig. 8 The same variables as in Fig. 7 but with a separate coolant circuit using a liquid

coolant (Duratherm 600) with an inlet temperature of 450 K, and a cathode air inlet

temperature of 400 K (stoichiometric factor of 2).

The results of these calculations are summarized in Table 2. Here, the computed

average current density is listed for two coolant temperatures, two cathode air inlet

temperatures, two stoichiometric factors and three cell sizes. The current density is a strong

function of temperature, and the calculated average current density values are thus an

indication of the temperature variation over the cell. Low values of the current density

indicate large deviation in temperature across the cell. The ideal scenario would be a uniform

temperature over the cell at the highest sustainable operating temperature, which is taken as

200 °C or 473 K for the purposes of this work. In the liquid coolant circuit case, there is a

pronounced decrease in the average current density with increasing cell size when the coolant

inlet temperature and the air inlet temperature are both low. When both inlet temperatures are

high, there is very little effect of size on the average current density and a high current/power

density (amounting to nearly 92% of the ideal, uniform temperature value) can be achieved

for all sizes. In the case of integrated cathode air cooling, a significant size effect can be seen

with the penalty on current density (power output) increasing from 13% to 17% as the cell

size is increased from 10 × 10 cm to 30 × 30 cm.

15

Table 2 Computed average current density for stacks with liquid cooling at a cell operating

volage of 0.5 V. The last row corresponds to the case of integrated air cooling (see Table 1).

Coolant inlet

temperature

(K)

Cathode air

inlet

temperature

(K)

Cathode air

stoichiometric

factor

Average current density (A/cm2) for

cell size (mm × mm) of

100 × 100 200 × 200 300 × 300

400 300 1 0.737 0.667 0.649

400 300 3 0.769 0.720 0.706

450 400 1 0.837 0.834 0.832

450 400 3 0.839 0.837 0.839

300 7.85–9.09 0.794 0.766 0.755

The above results show that with a liquid coolant, the cell temperature variation over

the catalyst can be controlled within a tight range by having a high coolant inlet temperature.

This results in a fairly uniform temperature and current density even for large cells. Thus, a

liquid coolant system would be the preferred method for thermal management of a high

temperature PEMFC stack of power output in the range of 5 to 15 kWe.

Pressure drop considerations

The choice of the flow field is important because for optimal performance, the

coolant should be able to provide effective cooling to all parts of the cell. A more uniform

coolant distribution can be assured by having, for example, a four-parallel serpentine channel

configuration for the coolant plate as well. However, as shown above, for high average

current densities, it is necessary to maintain high coolant inlet temperatures. Since the

maximum temperature anywhere in the cell is limited to 473 K, increasing the coolant inlet

temperature reduces the allowed coolant temperature increase from the inlet to the outlet.

Therefore, the coolant flow rate must increase to remove the same amount of heat (the

amount of heat removal required is fixed by the operating current and voltage). Due to this,

the coolant flow rate increases drastically when the coolant inlet temperature approaches the

cell operating temperature. This leads to very high pressure losses in the coolant channels.

The problem becomes even more severe as the cell size increases because the length of

traverse varies as ∼l2. Also the length of serpentine channels (l) is much greater, thus the

pressure drop required to maintain coolant circulation will be much higher.

In view of this difficulty–of needing to maintain high coolant inlet temperatures while

keeping the pressure losses low– three possible coolant flow field configurations have been

16

considered in the present study. These are schematically shown in Fig. 3 and can be described

as follows:

•

A parallel flow field with a channel width of 1 mm and a land width also of 1 mm.

•

A parallel flow field with a channel width of 2 mm and a land width of 0.5 mm with

the land width being 1 mm for every third channel.

•

A four-parallel serpentine channel with channel and land widths of 1 mm each.

While the first and the third flow fields require no further elaboration, the second one

is designed to reduce the pressure drop by increasing the width of the flow channels.

However, the rib width decreases as the flow channel width is increased, and this can lead to

buckling of the interconnect/flow-plate. Another factor to keep in mind is that the effective

resistance of the interconnect will increase as the rib width decreases. Thus one cannot

increase the ratio of flow channel width to rib width indefinitely, and the larger land area for

every third channel provides additional mechanical support as well as additional area for

current collection required for the second design above. One could perform a detailed

structural analysis to come up with an ‘optimal’ ratio of flow channel width to rib width but

this is beyond the scope of the current study. Increasing the channel width also reduces the

number of parallel channels for a given cell face area and thus reduces the flow

maldistribution problem associated with parallel flow channels [40].

Flow and temperature calculations have been done for different coolant plate

configurations for a cell size of 10 cm × 10 cm. The results are summarized in Table 3 where

the pressure drop, the coolant flow, maximum ΔT (across the volume of one cell), and

average current density are given for each case. It can be seen that while the thermal and

electrical efficiency of the cell improves with increasing coolant inlet temperature, the

coolant flow rate requirement increases by an order of magnitude when the coolant inlet

temperature is increased from 400 K to 450 K (for a maximum cell temperature of 473 K)

leading to a more than 20-fold increase in the pressure drop across the cell. The pressure drop

with the FPSC flow field for the cooling plate is nearly 50 times higher and is of the order of

3 bar. The effect of cooling duty can also be seen in the fact that the pressure drop decreases

by about 30% when the cathode air stoichiometric factor is increased from 1 to 3, which

lowers the cooling requirement from the liquid coolant circuit.

Table 3 Comparison of different flow field configurations in the cooling plate: PC = parallel

channel; FPSC = four parallel channel; CCL = cathode catalyst layer.

17

Inlet

temperature

of coolant

(K)

Flow field

configuration

(Channel

width)

Cathode air

stoichiometric

factor λ

Maximum

temperature

variation (K)

across

Average

current

density

(A/cm2)

Pressure

drop

across

coolant

plate

(Pa)

Coolant

flow

rate

(kg/s)

CCL-

1

CCL-

2

400 PC (1 mm) 1 50 42 0.721 303 1.38E-

04

400 PC (2 mm) 1 49 40 0.737 100 1.56E-

04

400 FPSC

(1 mm)

1 34 27 0.805 13,336 1.96E-

04

400 PC (1 mm) 3 43 41 0.771 173 7.92E-

05

400 PC (2 mm) 3 41 41 0.759 81 1.26E-

04

400 FPSC

(1 mm)

3 33 44 0.825 10,508 1.55E-

04

450 PC (1 mm) 1 24 28 0.835 5714 3.42E-

03

450 PC (2 mm) 1 19 25 0.839 2608 3.81E-

03

450 FPSC

(1 mm)

1 21 28 0.837 463,327 2.23E-

03

450 PC (1 mm) 3 32 45 0.826 2615 2.44E-

03

450 PC (2 mm) 3 30 43 0.826 1422 2.15E-

03

450 FPSC

(1 mm)

3 31 45 0.822 315,366 1.37E-

03

If the results are extrapolated to a cell size of 30 cm × 30 cm at the same average

current density, then the coolant pressure drop increases by a factor of 10. As such high

pressure drops (∼30 bar) are clearly unacceptable, it is necessary to use parallel flow

channels on the coolant side for large cells. Since simple parallel channel flow fields (with a

non-contoured header) can give rise to severe maldistribution of flow among the parallel

18

channels [40], and given that the number of cells is rather small in the present stack, we

propose a branching manifold of the type shown in Fig. 9. Here, the incoming pipe is

branched into 2, 3 or 4 parallel branches, each of which is further subdivided into further

branches. The diameters of the branching tubes can be adjusted to ensure that the pressure

drop at entrance to each is the same through all the paths (in actual practice, this branched

circuit can be etched on to one side of the stack). Fig. 9a shows the branched inlet manifold

for the 24 cells of the stack with four levels of branching and Fig. 9b shows the inlet manifold

for the seven coolant plates with two levels of branching. At each stage of splitting, the

diameters of the branches can be fixed such that the pressure drop across the split is the same.

Fig. 9 Branched inlet manifolds for (a) the cells, and (b) the cooling plates.

Calculations of fluid flow and heat transfer under integrated cathode air cooling and

separate liquid coolant system have been done for a 30 cm × 30 cm cell stack using the

manifolding strategy described above. For a 24-cell stack at an operating current density of

0.89 A/cm2 and cell potential of 0.5 V (due to the higher current density and lower voltaic

efficiency, this would constitute a severe test of thermal management) corresponding to a

nominal stack power of 9.65 kWe, the pressure drops in the inlet manifolds have been

estimated using standard correlations for pipe flow for frictional and bend losses. The

19

pressure drops in the flow fields (FPSC for cathode flow field and parallel flow fields for

coolant channels) have been calculated as part of the CFD simulations.

The results are summarized in Table 4 where the combined pressure drops in the inlet

and outlet manifolds (the outlet manifold pressure drop is assumed to be equal to that for the

inlet manifold) and those in the flow fields are shown as a function of the air and fuel

stoichiometric factors. It can be seen that in all cases, the pressure drop in the manifolds is

lower than that in the active area of the stack. Even though there are a large number of

subdivisions to feed the 24 cells, the manifold pressure drop for fuel (hydrogen) and air is

significantly lower than the pressure drop in the cells. For the coolant, the manifold pressure

drop is at least an order of magnitude lower than the coolant pressure drop in the cell. The

(liquid) coolant pressure drop is quite low for low inlet temperatures for the coolant and air.

The ΔP for the coolant is of the same order of magnitude as the cathode air side ΔP for

higher inlet temperatures and increases as both the coolant as well as air inlet temperatures

are increased. The present calculations show that using parallel channels for the liquid

coolant enables the total pressure drop across the active area as well as the manifolds to be

less than 0.3 bar (air sideλ ∼ 3) for a 10 kWe stack using Duratherm 600 as a coolant.

Table 4 Pressure drops for the reactants and the coolant for a 24-cell, 9 kWe stack with a cell

active area of 30 cm × 30 cm for the case of a liquid coolant circuit.

Medium Inlet temperature

(K) of

Pressure drop (Pa)

ΔP cell ΔP manifold Total

Coolant Cathode

air

λ = 1 λ = 3 λ = 1 λ = 3 λ = 1 λ = 3

Hydrogen 708 210 918

Air 1974 6867 719 4920 2693 11,787

Duratherm

600

400 300 932 782 22 19 954 801

450 300 33,774 11,068 2870 521 36,644 11,589

450 350 39,996 17,503 3573 1062 43,569 18,565

450 400 37,774 20,230 3759 1412 41,533 21,642

Conclusions

A number of scaling issues related to thermal management of HT-PEMFC stacks in

the power range of 1 to 10 kWe have been studied using CFD simulations and analytical

estimates. The following conclusions can be drawn from this study:

•

20

For stacks much larger than 1 kWe, cooling of the stack by convection drawing air

over the external surface of the stack can lead to unacceptable temperature variations within

the stack. Thus cooling must be done using dedicated plates through which a coolant is

circulated.

•

For stacks in the 1 to 5 kWe range (cells that are 10 to 2o cm long), circulation of the

cathode air through the cooling plate channels at a stoichiometric factor of about 10 should be

sufficient to maintain the temperature variation over the cell to within 40 K. For larger cells

and/or a more uniform temperature distribution, a separate coolant circuit using a liquid

coolant such as Duratherm 600 can be used to control the variations to within 20 K.

•

While using a four-parallel serpentine flow field for the bipolar plates appears to be a

good working compromise between the need for ensuring uniform reactant distribution and

minimizing parasitic pumping losses, the pressure drops for large cells can be as high as 10 to

30 bar.

•

It may be therefore necessary to use parallel channel flow fields in the cooling plate

(with an attendant need for careful design of the inlet and outlet headers) to avoid excessive

pressure losses which are associated with the need to operate at high coolant inlet

temperatures. Use of wider parallel channels will have the dual advantage of increasing flow

uniformity (by having fewer numbers of parallel channels) and lower pressure drop (by

increasing the hydraulic diameter of the channel).

In summary, the options for thermal management of HT-PEMFC stacks of the size of

10 kWe are limited. A carefully designed liquid coolant circuit, including identification of

optimal operating conditions, is necessary to maintain uniform and high operating

temperature of the fuel cells in the stack.

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