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Design and testing of a 9.5 kWe proton exchangemembrane fuel cellesupercapacitor passive hybridsystem

Billy Wu a,b,*, Michael A. Parkes b, Vladimir Yufit b, Luca De Benedetti a,Sven Veismann a, Christian Wirsching a, Felix Vesper a,Ricardo F. Martinez-Botas a, Andrew J. Marquis a, Gregory J. Offer a,Nigel P. Brandon b

aDepartment of Mechanical Engineering, Imperial College London, SW7 2AZ, UKbDepartment of Earth Science and Engineering, Imperial College London, SW7 2AZ, UK

a r t i c l e i n f o

Article history:

Received 4 October 2013

Received in revised form

23 February 2014

Accepted 10 March 2014

Available online 13 April 2014

Keywords:

Proton exchange membrane fuel cell

Supercapacitors

Passive system

Balance of plant

Electric vehicle

* Corresponding author. Department of EarthE-mail address: billy.wu06@imperial.ac.u

http://dx.doi.org/10.1016/j.ijhydene.2014.03.00360-3199/Copyright ª 2014, Hydrogen Ener

a b s t r a c t

The design and test of a 9.5 kWe proton exchange membrane fuel cell passively coupled

with a 33 � 1500 F supercapacitor pack is presented. Experimental results showed that the

system reduced dynamic loads on the fuel cell without the need for additional DC/DC

converters. Fuel efficiency gains of approximately 5% were achieved by passive hybrid-

isation in addition to addressing two main operational degradation mechanisms: no-load

idling and rapid load cycling.

Electrochemical Impedance Spectroscopy measurements indicated that the super-

capacitor capacitance dropped with decreasing cell voltage and suggested that operation

below 1.3 V is not recommended. Knee-frequency measurements suggested little benefit

was gained in using passive systems with load cycles that have frequency components

above 0.19 Hz. Analysis of system sizing suggested using the minimum number of

supercapacitors to match the open circuit voltage of the fuel cell to maximise load

buffering.

Copyright ª 2014, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rights

reserved.

Introduction

Fuel Cells (FC) are electrochemical energy conversion devices

that turn a chemical fuel into electrical energy at an efficiency

typically higher than direct combustion. Of the different

Science and Engineeringk (B. Wu).

83gy Publications, LLC. Publ

varieties of FCs, low temperature proton exchangemembrane

fuel cells (PEMFC) fuelled by hydrogen have received the

greatest attention with respect to their automotive applica-

tions as a possible replacement for the Internal Combustion

Engine (ICE) [1]. Barriers to mainstream adoption of Fuel Cell

Vehicles (FCV) include cost, durability and lack of hydrogen

, Imperial College London, SW7 2AZ, UK.

ished by Elsevier Ltd. All rights reserved.

i n t e rn a t i o n a l j o u r n a l o f h y d r o g e n en e r g y 3 9 ( 2 0 1 4 ) 7 8 8 5e7 8 9 67886

refuelling infrastructure. Addressing the issue of degradation,

the United States Department of Energy (US DOE) stated that

in order for FC systems to compete with ICEs, lifetimes of

more than 5000 h with less than a 10% loss in performance

needs to be achieved [2]. Under automotive applications, the

three main causes of PEMFC degradation have often been

cited as being: no-load idling, rapid power cycling and start-

up/shut-down cycles [3].

No-load idling leads to accelerated rates of carbon corro-

sion, especially so on the Cathode Catalyst Layer (CCL). This

corrosion is mainly associated with the electrochemical

oxidation of the carbon particles supporting the catalyst, and

can lead to a loss of catalytic activity [4,5]. It is accelerated by

the FC running at elevated potentials which exceed the carbon

oxidation potential of 0.518 V. Two mechanisms of carbon

corrosion are shown in Eqns. (1) and (2). Carbon corrosion can

also be accelerated under conditions where the potential

changes rapidly through load cycling or through fuel starva-

tion leading to cell reversal [6].

Cþ 2H2O/CO2 þ 4Hþ þ 4e� E0 ¼ 0:207VRHE 25 �C (1)

CþH2O/COþ 2Hþ þ 2e� E0 ¼ 0:518VRHE 25 �C (2)

Another process affecting FC durability is platinum cata-

lyst dissolution. This leads to the decrease of effective catalyst

surface area and a loss of electrochemical performance. Plat-

inum dissolution may proceed through a number of mecha-

nisms including; coarsening of the platinum particles in the

catalyst layer [7] and Oswald ripening [8], diffusion of plat-

inum into the membrane [9] and a thermodynamically

favourable particle agglomeration [10]. While the actual

mechanism is still debated, platinum dissolution rates have

been linked to FC operating conditions. Mathias et al. [11]

showed that platinum dissolution could be reduced by

lowering the operating voltage of a FC at low relative humid-

ity, while Borup et al. [12] found that load cycling significantly

accelerated platinum dissolution in comparison to steady

state operation.

During start-up/shut-down events, a hydrogen/air bound-

ary forms which can result in large internal currents; accel-

erating the degradation of FC electrodes. Maranzana et al. [13]

showed that these internal currents can be as high a 1 A/cm2

and vary depending on the flow rate of gases during start-up.

Rapid load cycling of PEMFCs also leads to inefficiencies

due to increased mass transport losses. Gas manifolding ef-

fects, compounded with compressor inertia can result in a

delay time on the order of seconds before a requested increase

in the mass flow rate of air is provided. By comparison, the

rate of electrochemical reaction of the Oxygen Reduction Re-

action (ORR) is several orders of magnitude faster. Therefore,

unless the system is operated with a high oxygen stoichiom-

etry, or the load ramp rate is limited, rapid increases in the

load will result in increased mass transport losses.

To alleviate these detrimental effects and allow the

downsizing of the FC stack, hybridisation with batteries and/

or supercapacitors has been suggested. Yufit and Brandon [14]

developed a FCebattery hybrid system which was actively

controlled using a DC/DC converter where the lithium-ion

battery was used to meet peak power requirements whilst

the DC/DC converter maintained the FC under a constant

current load. Hosseinzadeh et al. [15] investigated the theo-

retical optimisation of a FCelead acid battery system for a

forklift truck powertrain. An optimal design was found to give

theoretical reductions in hydrogen consumption by up to 26%

while Garcia et al. [16] modelled the performance of a

FCebatteryeSC powertrain. Using lithium-ion batteries and

supercapacitors connected via DC/DC converters, the hybrid

powertrain was found to smooth load cycling and enable the

FC to operate at conditions close to steady state. Efficiencies of

the FC were predicted at 60.9% while the hybrid vehicle had a

system efficiency of 55.6%. Hwang et al. [17] investigated the

performance of a FCelithium-ion battery hybrid for use in a

light electric vehicle. This powertrain was found to have an

efficiency of 46%. Hwang also reported on the performance of

a FCelithium-ion battery generator [18]. This system was able

to smooth rapid changes in load while achieving FC effi-

ciencies between 52 and 58%.

The addition of the DC/DC converters in all examples,

however, introduces an extra level of power conversion,

which typically ranges from between 90 and 99%, depending

on the load and construction of the converter [19,20]. The

additional cost of this extra component can also be significant,

with 2010 costs in the region of 25.9 $/kW�1 [21] and predicted

2030 costs of 6e19 $/kW�1 [22]. The concept of passive

hybridisation has therefore been proposed as a means of

gaining the benefits of hybridisation without the addition of

large costly power electronics. In this configuration, the FC

and secondary energy storage device are connected in parallel

directly. During operation the voltages of both devices are

therefore intrinsically balanced.

Several studies have previously examined the effect of

passive hybrid systems. Kuperman et al. [23] used a frequency

domain based analysis to investigate the potential benefits of

a batteryesupercapacitor passive hybrid. It was shown that

supercapacitors act as a low-pass filter for the battery under

pulse loading, leading to reduced battery stress. The disad-

vantage of passive hybridisation was highlighted as being the

loss of active power regulation. Bernard et al. [24] showed that

active pressure control of a FC anode and cathode can be used

to regulate power flows in a FCebattery passive hybrid sys-

tem. Nishizawa et al. [25] developed a FCebattery passive

hybrid system for aircraft applications and showed that there

was an increase in efficiency compared to equivalent systems

using DC/DC converters and that the lower impedance

response of the battery resulted in more mild transient loads

for the FC. Keranen et al. [26] developed a FCesupercapaci-

torebattery triple hybrid and showed that benefits of passive

hybridisation could be achieved with a relatively small

supercapacitor pack and that system voltage should be near

the maximum voltage of the supercapacitor pack due to small

voltage variations.

Within the academic literature, there have been no sys-

tems that have studied passive hybridisation between a FC

and supercapacitor. Therefore, this paper presents electro-

chemical characterisation of supercapacitor cells and FC

stacks to provide an unique insight into the coupled perfor-

mance of a FCesupercapacitor passive hybrid system and

propose a sizing methodology that can be used in automotive

applications.

Fig. 1 e Schematic of balance of plant system for a 9.5 kWe PEMFC stack. Adapted from Wu et al. [27].

i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n en e r g y 3 9 ( 2 0 1 4 ) 7 8 8 5e7 8 9 6 7887

Fuel cell system

The FC stack used in this study was a Nedstack P9.5-75, which

comprised of 75 cells (Johnson Matthey), each with an active

area of 200 cm2. The stack of 75 cells was rated to supply

9.5 kWe of power.

Balance of plant

A schematic of the supporting Balance of Plant (BOP) system,

designed and built in-house for the 9.5 kWe Nedstack, is

shown in Fig. 1.

The air subsystem draws in ambient air using a 24 VDC low

pressure blower followed by a membrane humidifier before

entering the cathode inlet of the stack. Oxygen depleted air

from the cathode outlet was then channelled back into the

humidifier which recovered heat and water. The air was then

released to the atmosphere through a back pressure regulator.

Air stoichiometry was maintained during normal operation at

a target value of 2 through a feedforward controller.

Hydrogen was supplied at 1.5 bar at the inlet. The dry

hydrogen entered into an in-house designed bubble humidi-

fier, through de-ionised water, that was also used to recover

heat from the cooling loop through a heat exchanger. In order

to prevent an uneven distribution of hydrogen across the in-

dividual cells, the exhaust hydrogen was recirculated back to

the inlet with a diaphragm pump to increase the working

stoichiometry to >1.5. Uneven hydrogen distribution across a

single cell can occur as result of the flow plate design and

reduction in partial pressure of hydrogen due to the con-

sumption of reactants. The anode gas loopwas also connected

to a purge valve which periodically removed the mixture of

nitrogen crossover, water and hydrogen, minimising the

cause of localised cell flooding.

In order to minimise the effect of internal currents, caused

by the hydrogen front propagation across the flow channels

during start-up events, a nitrogen supplywas used to purge air

in the anode manifolds and Gas Diffusion Layer (GDL). In

addition, the nitrogen supply was used to remove excess

hydrogen in the anode after use.

To remove excess heat and maintain the FC stack at a

constant operating temperature of 65 �C, a cooling loop which

consisted of a centrifugal pump, radiator and PID-controlled

fan was implemented.

Control and monitoring of the BOP components was real-

ised with a National Instruments (NI) compactRIO (cRIO)

running LabVIEW�. A detailed description of the sensors and

equipment used for the FC can be found in Wu et al. [27] and

Cordner et al. [28].

Fuel cell stack test

The currentevoltage characteristics of the FC stack will

essentially define the operating region of the supercapacitors

in a passive hybrid system, since the potentials are linked. For

a FC, the voltage is a function of temperature, hydration and a

range of other external conditions. The optimum operating

temperature of a conventional PEMFC is generally regarded as

being in the range of 60e80 �C [29]. Higher temperatures risk

dehydration of the Nafion� membrane due to water evapo-

ration, subsequently increasing the likelihood of mechanical

failure of the membrane and pin-hole formation [30]. Lower

temperature operation results in slower electrochemical ki-

netics and increased likelihood of flooding due to the reduced

capacity of the air to remove produced water. Fig. 2(a) shows a

low temperature (30 �C) polarisation curve for the Nedstack

P9.5-75 stack before the membrane was hydrated (left idle for

>1 week). The stoichiometry of the cathode and anode were

maintained at 2 and 1.5 respectively, with the relative hu-

midity of the cathode and anode at >99% and 70% respec-

tively. This represents the typical operating conditions of the

designed system. The achieved peak power was only 4 kWe

due to lower membrane conductivity (low hydration) and

slower rates of electrochemical kinetics (low temperature). At

an operating temperature of 65 �C and fully humidified con-

ditions, the peak power reached 8 kWe. The full 9.5 kWe power

was not achieved due to the low pressure operating

Fig. 2 e Comparison of ambient temperature/dry start and

operational temperature polarisation curves (a) and stack

resistance values (b).

Fig. 3 e Bode plot of total capacitance measured for a 1500 F

Maxwell supercapacitor at different OCP at a temperature

of 20 �C.

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limitations of the blower used,whichwas not able to achieve a

pressure greater than 1.2 bar (absolute). This limited both the

anode and cathode operating pressure.

The resistance of the FC stack varies with load current.

Fig. 2(b) shows the resistance of the FC stack which was

calculated based on Eqn. (3), where R, V and I represents the

resistance, voltage and current, respectively. Here VOCV is the

no-load voltage of the FC stack.

R ¼ DVDI

¼��VOCV � Vworking

��jIj (3)

The resistance of the FC and supercapacitor will define the

instantaneous power distribution between the devices in a

coupled system. At higher currents, the stack resistance de-

creases due to a combination of increased water production at

higher currents, leading to higher membrane conductivity,

and lower charge transfer resistance as suggested by the

exponential component in the ButlereVolmer equation

shown in Eqn. (4) for a 1 electron reaction. Here i represents

the current density, i0 the exchange current density, aa/ac the

anodic and cathodic charge transfer coefficients, F is Faraday’s

constant, R is the universal gas constant, T is temperature and

h is the overpotential. Operation at even higher currents

showed an increase in resistance, which is likely due to a

combination mass transport effects and dehydration of the

anode due to osmotic drag effects.

i ¼ i0

�exp

�aaFhRT

�� exp

�acFhRT

��(4)

Supercapacitor characterisation

Direct coupling necessitates the balancing between the po-

tentials of a supercapacitor pack and the FC stack. Therefore,

in order to appropriately size the system and understand its’

limitations, the dynamic performance of supercapacitors, in

terms of impedance and capacitance behaviour at different

frequencies and State-Of-Charges (SOC), were characterised.

Supercapacitors store energy via the formation of a

charged double layer at the interface between its porous car-

bon electrodes and an electrolyte. Due to the absence of the

charge transfer process, supercapacitors are characterised by

relatively long lifetime, low internal impedance and fast

response time. However, their energy density is limited, as is

their low operating potential (2.7e2.85 V) due to the possible

onset of electrolyte decomposition.

The supercapacitor pack constructed for this study con-

sisted of 33 � 1500 F Maxwell supercapacitors connected in

series giving a maximum operating potential of 89.1 V and

rated capacitance of 45.45 F. It should be noted that the

maximum pack voltage was never reached during operation

because of the limitation imposed by open circuit voltage

(OCV) of the FC which was 72 V. Redundancy was added to

compensate for the absence of a cell balancing system.

In order to characterise a single supercapacitor cell, Elec-

trochemical Impedance Spectroscopy (EIS) measurements at

different OCP were recorded in galvanostatic mode with a 1 A

current amplitude on a Biologic VSP multichannel potentio-

stat/FRA equippedwith 5 A booster. Prior to themeasurement,

the cell was charged or discharged at constant current to a

target potential and then held at this potential until the cur-

rent dropped to <1mA, after which a 30min settling time was

allowed before the measurement was taken. By extracting the

Fig. 4 e (a) Nyquist plot of impedance and (b) fitted series,

pores, total polarisation resistances and DL capacitance for

a 1500 F Maxwell supercapacitor at 20 �C.

Fig. 5 e Bode plot of imaginary capacitance for a 1500 F

Maxwell supercapacitor at different cell voltages. Inset

shows the knee-frequency dependency with cell voltage at

20 �C.

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real (ZRe) and imaginary (ZIm) components of the total

impedance (jZj) from the EIS measurements, it was possible to

describe the complex capacitance as shown in Eqns. (5)e(7)

[31]. Where Ctot, CRe and CIm are the total complex, real and

imaginary parts of the capacitance respectively.

Ctot ¼ CRe � jCIm (5)

CRe ¼ �ZIm

ujZj2 (6)

CIm ¼ ZRe

ujZj2 (7)

A bode plot of the total capacitance at different OCP values

is shown in Fig. 3. Capacitance was observed as being fre-

quency dependent; at higher frequencies under AC current

oscillations, ions do not have sufficient time to diffuse into

mesopores of the electrode, thus not utilising the available

surface area for ion adsorption resulting in a lower accessible

capacitance. At lower frequency, there is sufficient time for

ions to diffuse into the mesopores and access the full surface

area [32]. At higher potentials, the measured capacitance is

greater as result of several physical phenomena such as

reduction of the solvent layer thickness, increase of the

solvent dielectric constant [33], increase in the electronic state

density in the carbon pore walls [34] and higher ion penetra-

tion in mesopores [35]. The available capacitance will directly

impact the dynamic response of the coupled FCesupercapa-

citor system.

The impedance of the cell also varies with the cell voltage.

Fig. 4(a) shows the Nyquist plot of the impedance and Fig. 4(b)

shows the fitted EIS results according to the equivalent circuit

presented in the inset. The equivalent circuit contains high

frequency elements L (induction), Rs (series resistance) and an

impedance ZP of porous electrode as described by de Levie

equation as shown in Eqn. (8) [36]. The inductive element was

added to give a better fit at high frequencies rather than infer

any physical meaning, as this was a function of the experi-

mental set-up.

ZP ¼ffiffiffiffiffiffiffiffiffiRiZi

pcoth

ffiffiffiffiffiffiffiffiffiffiffiRi=Zi

p(8)

In Eqn. (8), Ri represents ionic resistance inside the pores

and Zi is the interfacial resistance which, in absence of any

Faradaic process, is simply a capacitance of double layer (DL):

Zi ¼ 1/(juCDL). It then follows that the equation can be re-

arranged to give Eqns. (9) and (10)

ZP ¼ Ricoth

ffiffiffiffiffiffiffijus

pffiffiffiffiffiffiffijus

p (9)

s ¼ RiCDL (10)

The total resistance (Rs þ Ri) of a single 1500 F super-

capacitor was therefore found to vary between 0.68 and

0.92 mU as SOC increased, resulting in a pack resistance of

22.4e30.4 mU, neglecting contact resistances between

different cells.

Although DL capacitance does not vary linearly with

voltage, it appears there are regions when the variation can be

approximated as linear: 0.5e1.3 V, 1.3e2.1 V and 2.1e2.7 V.

The capacitanceevoltage gradients within these regions were

134, 42 and 360 F/V respectively. Therefore, moving the

Fig. 6 e Nyquist plot of complex capacitance for a 1500 F

Maxwell supercapacitor for different potentials at 20 �C.

Fig. 7 e Wiring diagram for powertrain components in the

9.5 kWe PEMFCesupercapacitor passive hybrid system.

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average operating point between 1.3 and 2.1 V will have little

impact on the available capacitance, but varying it between

2.1 and 2.7 V will have more significant improvements in

performance, as a larger capacitance will enable slower load

dynamics for the FC. As the available low frequency capaci-

tance drops off further below 1.3 V, the load buffering capa-

bilities of the supercapacitors decreased. System designers

should therefore aim to size their systems such that the

lowest operating voltage of supercapacitor pack will be below

that at which mass transport limitations occurs in the FC

stack because highly transient loads on the FC during mass

transport limited operation can lead to accelerated

degradation.

The frequency dependency of the total complex capaci-

tance was also analysed by considering the bode plot of its

imaginary part as shown in Fig. 5. The peak value in the

imaginary capacitance occurs at a frequency known as the

knee frequency which represents the transition point above

which the available capacitance becomes frequency inde-

pendent. This frequency also represents the 50% point of the

charge storage efficiency [37]. The inset in Fig. 5 shows the

voltage dependency of the knee frequency. Due to the discrete

measurement frequency of the EIS measurements, fitting of

the curve allowed for a more accurate estimate of the knee

frequency. The exact value for the knee frequency was esti-

mated by fitting the natural log of the frequency-imaginary

capacitance data to a normal distribution.

As the potential of the cell decreased, the knee frequency

increased. This was likely to be caused by a combination of

processes mentioned previously. It also became apparent that

above the knee frequency, the imaginary capacitances

become independent of cell voltage. Again, the measured

knee-frequency dependency is non-linear over the whole

range of operation, however it can be considered as linear

within the operating ranges of 0.5e1.3 V, 1.3e2 V and 2e2.7 V

with knee frequencyevoltage gradients of �29.5, �13.4 and

�57.6 mHz/V respectively. Important aspects to consider, for

system designers, are that the load buffering capabilities of

the supercapacitors are reduced if the applied load has fre-

quency components oscillating above the knee frequency. It is

therefore advised that, for highly dynamic load cycles, the

knee frequency should be the upper limit of load oscillation.

Little benefit is achieved in using a FCesupercapacitor passive

hybrid system for loads cycles with frequency components

above the knee frequency.

As both the real and imaginary components of the capac-

itancewere available, it is also possible to present the complex

capacitance data in the form of a Nyquist plot in a similar

fashion to Nyquist plots of impedance. This is shown in Fig. 6

for different cell voltages. The same trend of increasing

capacitance with voltage can be concluded; however, the

relationship between the complex variations of capacitance is

now more evident, as is the low frequency capacitance which

is the point at which the semi-depressed circle would inter-

cept the real axis at an infinitely small frequency. This

potentially offers an alternative means of estimating the low

frequency capacitance of a supercapacitor without having to

measure to the extremely low frequencies.

Combined FCesupercapacitor passive hybridsystem

The wiring diagram for the combined FCesupercapacitor

passive hybrid system is shown in Fig. 7. Contactors between

the FC, supercapacitor and load provided electrical isolation

for each component. A 310 A fuse, in-line with the FC and

supercapacitors, provided additional passive protection in the

case of large in-rush currents. A diode was placed in front of

the FC to prevent possible charging of the FC stack from the

supercapacitors in case the thermodynamic OCP of the stack

dropped below that of the supercapacitors.

All BOP components were powered by a 24 VDC supply via

a low voltage bus. During start-up/shut-down events, the BOP

was powered by two 12 V lead acid batteries connected in

series. During normal operation, two parallel DC/DC con-

verters delivered power the low voltage bus.

Prior to connecting the FC and supercapacitors to the main

power bus, the supercapacitors were pre-charged to just

below the FC OCP via an external charging circuit to prevent

large in-rush current from the FC to the supercapacitors when

the two devices were initially connected.

Table 1 e Component list for fuel cellesupercapacitorhybrid system.

Product Manufacturer Product number/Name

DC/DC converters RS components 491-257

Sealed lead acid

battery

RS components 727-0391

Lead acid battery

charger

Ideal Power AC 0724A

Blower Domel 497.3.265

Water pump Jabsco 50870 Series

Fan Comex Europe Axial Fan 24V

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A resistor was wired in parallel to the FC and connected

through a contactor. This resistor was used during the shut-

down procedure under autonomous operation. Upon shut-

down, the hydrogen supply was closed however residual

amounts of reagents in the GDLs and flow channels gave rise

to an elevated OCP until the reagents were consumed. As

elevated potentials were detrimental for carbon corrosion, the

resistor was used to apply a load to the FC to facilitate the

consumption of residual hydrogen.

A Computer Aided Design (CAD) drawing of the FCesu-

percapacitor rig is shown in Fig. 8(a) with a photograph of the

rig shown in Fig. 8(b). Table 1 lists the components used.

Fig. 8 e (a) CAD drawing of fuel cellesupercapacitor passive

hybrid test system and (b) photograph of completed

system.

Recirculation pump Thomas 7015Z DC

Contactor 500 A TE Connectivity LEV200A5ANA

Contactor 20 A Finder 22.23.9.024.4000

Diode 190 A IXYS MMD172-08N1

Diode 120 A ON semiconductor STMSTPS24045TVG

80 A battery fuse Pudenz 153.5631.5801

Resistor 100 U

300 W

Farnell 1768254

Fuel cell Nedstack P9.5-75

Supercapacitors Maxwell BCAP 1500

Passive hybrid system test

The concept of the FCesupercapacitor passive hybrid system

is that the supercapacitors act as a low-pass filter to high

frequency oscillating loads which are detrimental to FC effi-

ciency and durability. EIS measurements indicated that the

total polarisation resistance of the 33 cell supercapacitor pack

was 22.4e30.4 mU. FC stack resistance was shown to range

from 1000 to 100 mU depending on the load current. Since the

supercapacitor resistance is an order of magnitude smaller

than the FC, it will initially handle the majority of the load.

Fig. 9 shows the dynamic response of the FCesupercapa-

citor passive hybrid to a 100 A pulse load from no-load con-

ditions. Upon application of the step load, the supercapacitors

initially handled all the load, as the thermodynamic OCP of

both devices was the same under steady state conditions.

Gradually, as the OCP of the supercapacitor dropped due to the

discharge, the operating potential also drops, meaning that

the FC started to handle a higher proportion of the load. When

Fig. 9 e Response of the FCesupercapacitor passive hybrid

system to a 100 A pulse load from no-load conditions.

i n t e rn a t i o n a l j o u r n a l o f h y d r o g e n en e r g y 3 9 ( 2 0 1 4 ) 7 8 8 5e7 8 9 67892

the external load was removed, the FC charged the super-

capacitors as the OCP of the supercapacitors was lower than

that of the FC.

Smoothing of the transient load, imposed on the FC,

effectively translates to reduced losses and increased lifetime.

Fig. 10 compares the dynamic response of the pure FC and

FCesupercapacitor passive hybrid system to step loads of 75 A

from cold start conditions. In pure FC mode, the step loads

lead to an observed voltage drop followed by recovery. This

was due to the finite time required to overcome the blower

impeller inertia as well as gas manifolding effects. The rate of

electrochemical reaction was orders of magnitude faster than

the blower inertia and manifolding dynamics. As a result,

there was a brief period, in the order of seconds, when the

local operating stoichiometry of air was lower than the target

value of 2, causing additional mass transport losses and

reduced operating potentials. Under no-load conditions, the

stack voltage returned to OCP. The gradual recovery was due

to a combination of double layer charging and hydrogen

manifolding effects.

When the same load was applied to the passive hybrid

system, the supercapacitor pack met the peak load before the

Fig. 10 e Comparison between (a) pure FC and (b)

FCesupercapacitor passive hybrid under 75 A step loads

with a frequency of 0.2 Hz. Positive currents represent

discharge of the device.

FC. Having reached a dynamic equilibrium, the supercapacitor

operated around an average current of zero as it removed the

peaks from the load cycle. Consequently, the FC met the

average loadwith sufficient time for the air blower to adjust to

the transient loads. The result was therefore a removal of the

voltage drop caused by the blower inertia and gas

manifolding.

Comparison between the efficiency of the pure FC and

FCesupercapacitor hybrid under a 0.2 Hz square wave load of

varying amplitude can be seen in Fig. 11. The efficiency of the

FC was taken as the energy per mol of hydrogen consumed

against the enthalpy of formation assuming the Higher

Heating Value (HHV) for thewater generation reaction. In both

cases, the efficiency of the FC decreased with increasing pulse

load amplitudes due to the increased losses associated with

higher currents. Comparison between the pure FC and FCeSC

passive hybrid system showed that a 5% efficiency gain was

achieved. This was mainly as a result of operation at a lower

average current, and also reduced transient mass transport

losses associated with blower inertia and manifolding effects.

Pulse loading conditions, however, do not sufficiently

represent real load conditions. Therefore, the system was run

under a 1/10th scale (with respect to power) Highway Fuel

Economy Test (HWFET) drive cycle which is representative of

highway driving [38,39]. A power profile, assuming a vehicle

with a 2000 kg mass, accounting for inertia, aerodynamic and

rolling resistance losses was generated based on the govern-

ing equations given by Ehsani et al. [40]. A constant motor

efficiency of 95% was assumed. Other vehicle parameters

were taken to be the same as the values quoted by Baptista

et al. [41], which represents typical parameters for a London

taxi. Fig. 12 shows the response of the FCesupercapacitor

system to the HWFET drive cycle. Before the cycle started, the

systemwas allowed to equilibrate to 69 Vwhich represented a

charging current of less than 1 A from the FC to the super-

capacitors. After the load cycle, the system was again equili-

brated to the same point to allow for direct comparison

between the FCesupercapacitor results and the pure FCmode.

Fig. 11 e Comparison of fuel cell efficiency assuming HHV

between pure FC and FCeSC passive hybrid system with

the fuel cell operating with an air stoichiometry of 2 and

neglecting BOP power losses.

Fig. 12 e Fuel cellesupercapacitor passive hybrid power

distribution between devices with inset zoomed view.

i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n en e r g y 3 9 ( 2 0 1 4 ) 7 8 8 5e7 8 9 6 7893

It can be seen that the FC followed the load profile, however, it

did not experience most of the higher frequency loads which

were buffered by the supercapacitors.

A breakdown of the operating current distribution for the

FC and supercapacitor pack under the HWFET drive cycle can

be seen in Fig. 13(a). It becomes apparent that in the FCesu-

percapacitor hybrid configuration, the peak FC load was

reduced from 76 A to 62 A, resulting in an 18.4% decrease.

There was also a 2% decrease in the average load current of

the FC from 28.5 A for the applied load to 27.9 A. The super-

capacitors operated at an average load of 0 A with a normal

distribution and standard deviation of 11.2 A.

The reduced operation time at no/low load conditions for

the FC, resulted in a lower average operating cell potential for

the FCeSC hybrid system as shown in Fig. 13(b). This therefore

gave an indication of the possible improved durability of the

system from a FC side (near-OCP working potential of FC

corresponds to a higher carbon corrosion rate [11]).

The apparent advantage of this system was that it

addressed two of the main PEMFC degradation modes asso-

ciated with a vehicle operation: rapid power cycling, which

leads to catalyst dissolution, and no-load idling, which in-

duces accelerated rate of carbon corrosion. From an efficiency

point of view, the hybridisation reduced both the transient

voltage drops when operating in pure FC mode, upon the

application of large step loads, and BOP related losses.

Fig. 13 e Histograms of (a) currents and (b) single FC

potentials for a 1/10th scale HWFET drive cycle for the

FCesupercapacitor passive hybrid system.

Estimating buffering effect of the supercapacitors

Quantitatively, the buffering effect of the supercapacitor pack

can be estimated by considering the fundamental equation for

a capacitor as shown in Eqn. (11), where ISC, VSC and C repre-

sents the supercapacitor current, voltage and capacitance

respectively.

ISC ¼ CdVSC

dt(11)

In a passive system, if interconnection losses are neglec-

ted, the voltage of the supercapacitor pack and that of the FC

stack (VFC) are equal i.e. VSC ¼ VFC (see Fig. 10(b)). Applying

Kirchoff’s current law allowed for the expression of the

supercapacitor current in terms of the FC (IFC) and load (ILoad)

currents (ISC ¼ ILoad � IFC) as shown in Eqn. (12).

ILoad � IFC ¼ CdVFC

dt(12)

Assuming that the FC is operated in the ohmic dominated

region, and the FC resistance is largely invariant with current

i n t e rn a t i o n a l j o u r n a l o f h y d r o g e n en e r g y 3 9 ( 2 0 1 4 ) 7 8 8 5e7 8 9 67894

over a certain range (for example, for currents between 40 and

80 A at 30 �C the resistance of the FC is 0.29 U � 0.03 U), it can

be assumed that the FC voltage is related to the FC current as:

VFC ¼ VOCP � IFCROhmic � IFCRCTzVOCP � IFCRFC (13)

where RFC ¼ RCT þ ROhmic.

Here, VOCP represents OCP of FC, RCT is the charge transfer

resistance of the FC, ROhmic is the ohmic resistance of the FC

and RFC is the combined ohmic and charge transfer resistance

of the FC. Assuming that the OCP of the FC and ohmic resis-

tance are time invariant, the current is sufficiently large such

that the charge transfer polarisation losses do not vary

significantly with current, and can therefore be combined

with RFC, and the capacitance is constant at the potential

defined by the average load current and FC polarisation curve,

results in Eqn. (14).

Iload � IFC ¼ �CRFCdIFCdt

(14)

When compared to experimental data, this simple

approximation showed good agreement as demonstrated by

Fig. 14(a). Note that the first pulse and settling period exhibit

Fig. 14 e (a) Comparison between experimental and

simulated data for the FCesupercapacitor passive hybrid

system under a 75 A 0.2 Hz step load and (b) ratio of

supercapacitor to fuel cell resistance under cold/dry start

and nominal operating conditions.

the highest deviation from simulated data due to the larger

non-linear variation of charge transfer resistance at low cur-

rent densities. The validity of assuming a constant charge

transfer resistance when assessing the power split in a pas-

sive hybrid system can be highlighted by Fig. 14(b), which

shows the ratio of supercapacitor to FC resistance. It becomes

apparent that, over the whole operating range of the FC, the

supercapacitor impedance is consistently lower. For a passive

hybrid system, this dictates that variations in the FC resis-

tance will not significantly affect the power split between the

devices due to the dominant effect of the lower supercapacitor

resistance. This therefore allows for the assumption of con-

stant FC resistance in the simplified approximation presented.

It was therefore possible to compare the impact of the

passive load filtering offered for different numbers of super-

capacitors in the pack. Fig. 15 shows the rate of change of the

FC current plotted against the applied system load assuming

the FC current starts at zero. A lower rate of change was

preferable in order to reduce the level of dynamic loading on

the FC. It can be seen that at relatively low currents, the

FCesupercapacitor hybrid offers a good level of load buffering

and is largely independent of the number of supercapacitors.

However, as the load current increases, it becomes apparent

that having more cells in series is detrimental to the load

buffering.

The reason for this is evident by considering the voltage

and current dependencies between the FC and super-

capacitors as shown in Eqn. (15). Neglecting losses, the FC

voltage and supercapacitor voltage are equal. The rate of

change of supercapacitor voltage is the same as the rate of

change of the FC voltage. The resulting change in FC voltage

would therefore result in a proportional change in the rate of

change of FC current:

dVSC

dt¼ dVFC

dtfdIFCdt

(15)

By considering Fig. 16, it can be seen that if the same cur-

rent flows through the supercapacitors, increasing the

Fig. 15 e Comparison of different step loads and

supercapacitor pack size with the response time of the FC

current assuming the same 1500 F supercapacitors used in

this study.

Fig. 16 e Diagram showing how increasing the number of

supercapacitors in series increases the rate of change of

voltage for the supercapacitor pack. The faster the rate of

change of voltage for the supercapacitor pack results in

faster rate of change of load experienced by the FC.

i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n en e r g y 3 9 ( 2 0 1 4 ) 7 8 8 5e7 8 9 6 7895

number of supercapacitors in series will increase the rate of

change of voltage of the entire pack. This translates to an

increased rate of change of FC voltage and FC current. Hence,

the maximum buffering effect of the FCesupercapacitor sys-

tem was achieved when the number of supercapacitors in

seriesmatched themaximumOCP of the FC stack tominimise

the number of cells used. This reduced the rate of change of

voltage of the supercapacitor pack for a given current.

Conclusion

Rapid load cycling has been identified as being one of themain

causes of FC inefficiencies as a consequence of compressor

inertia and gas manifolding effects. Passive coupling of a FC

with a supercapacitor pack allowed for the reduction of dy-

namic loading experienced by the FC.

EIS measurements of a 1500 F Maxwell supercapacitor

indicated a proportional relationship between the DL capaci-

tance and cell voltage with 3 pseudo-linear regions. Operation

at potentials lower than 1.3 V resulted in the greatest decrease

in DL capacitance and therefore offered reduced buffering.

The knee frequency was identified as being inversely pro-

portional to cell voltage and ranged from 0.19 to 0.26 Hz for the

given supercapacitor cell. Load cycles with frequency com-

ponents greater than the knee frequency are not suitable for

the FCesupercapacitor passive hybrid system.

Efficiency gains achieved through passive hybridisation

under step loads have been demonstrated to be approximately

5% with respect to fuel utilisation. This was mainly attributed

to reduced transient mass transport losses and a shift in the

average operating current to a more efficient region. Under a

1/10th scaled HWFET drive cycle, it was shown that the

operating current of the supercapacitor has a normal distri-

bution with an average around zero indicating that it is purely

buffering loads. The FC load showed a 18.4% decrease in peak

loads with the average 2% lower. A shift in the average oper-

ating potential suggested lower rates of carbon catalyst sup-

ports will be afforded in the FCesupercapacitor passive hybrid

system.

Analysis showed that in order to gain themost benefit from

the passive system, the number of supercapacitors used

should be kept to a minimumwhilst ensuring the FC does not

overcharge the supercapacitors under no-load conditions.

Derivation of the system equations shows that the FC buff-

ering is determined by the product of the supercapacitor pack

capacitance and FC resistance.

Acknowledgements

The authors would like to thank Richard Silversides for

assistance with advice on the electric systems and previous

fuel cell development team members: Mardit Matian, Ralph

Clague, Mark Cordner, Sam Tippets, Ed Spofforth-Jones, Till

Hanten, Alana Johnson, Laura Harito, Charles Banner-Martin,

Akash Agrawal, Matthew Wong, Dan-Fung Chan, Tanya

Chong, Omar Al Fakir, Nicolas Lee, Robert Bilinski, Nicolas

Higginson, Rebecca Nelson, Michael Squire, Ashwin Suguna-

Balan, Ryan Williams, Jignesh Patel, Olivia Tillbert, Adya Jha,

Xin Miao, Nasrin Shahed-Khah and Sze Li.

The authors would also like to acknowledge the Engi-

neering and Physical Sciences Research Council for funding of

this work, through both a Career Acceleration Fellowship for

Gregory Offer, award number EP/I00422X/1. As well as the in-

kind contributions from: Johnson Matthey, Nedstack, BOC,

Domel, National Instruments, Swagelok and RS.

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