Design and Optimization of a Combined Fuel Reforming and Solid Oxide Fuel

7/31/2019 Design and Optimization of a Combined Fuel Reforming and Solid Oxide Fuel

1/13

Design and optimization of a combined fuel reforming and solid oxide fuel

cell system with anode off-gas recycling

Tae Seok Lee a, J.N. Chung a,, Yen-Cho Chen b

a Department of Mechanical and Aerospace Engineering, University of Florida, Gainesville, FL 32611, USAb Department of Energy and Resource Engineering, National United University, Miaoli City 360, Taiwan

a r t i c l e i n f o

Article history:

Received 21 June 2010

Received in revised form 8 May 2011

Accepted 11 May 2011

Available online 22 June 2011

Keywords:

SOFC

Steam reforming

Fuel processor

Anode off-gas

System efficiency

Dodecane

a b s t r a c t

An energy conversion and management concept for a combined system of a solid oxide fuel cell coupled

with a fuel reforming device is developed and analyzed by a thermodynamic and electrochemical model.

The model is verified by an experiment and then used to evaluate the overall system performance and to

further suggest an optimal design strategy. The unique feature of the system is the inclusion of the anode

off-gas recycle that eliminates the need of external water consumption for practical applications. The sys-

tem performance is evaluated as a function of the steam to carbon ratio, fuel cell temperature, anode off

gas recycle ratio and CO2 adsorption percentage. For most of the operating conditions investigated, the

system efficiency starts at around 70% and then monotonically decreases to the average of 50% at the

peak power density before dropping down to zero at the limiting current density point. From an engi-

neering application point of view, the proposed combined fuel reforming and SOFC system with a range

of efficiency between 50%and 70%is considered very attractive. It is suggested that the optimal systemis

the one where the SOFC operates around 900 C with S/Cratio higher than 3, maximum CO2 capture, and

minimum AOG recirculation.

2011 Elsevier Ltd. All rights reserved.

1. Introduction

Solid oxide fuel cells (SOFCs) can provide clean and highly effi-

cient power for a wide spectrum of small to large-scale applica-

tions. SOFCs are expected to be around 5060% efficient at

thermal efficiency. In applications designed to capture and utilize

the high-temperature waste heat through a co-generation system,

overall combined efficiencies could top 8085%. SOFCs operate at

very high temperaturesaround 8001000 C and they are also

the most sulfur-resistant fuel cell type; they can tolerate several

orders of magnitude more of sulfur than other fuel cell types. In

addition, they are not poisoned by carbon monoxide (CO), which

can even be used as fuel [1,2].

Hydrogen, a very versatile fuel, can be produced from various

materials and by several methods. Industrial scale production of

hydrogen has been operational in the oil and gas industry for more

than a century and forms the base of the modern chemical indus-

try. In centralized facilities and/or distributed generation, hydro-

gen has to be supplied to stationary fuel cell applications from

nearby hydrogen mass-production processes. For mobile applica-

tions, it is difficult to store and handle hydrogen directly as an

on-board fuel. Moreover, not only the safe hydrogen storage tech-

nologies, but also its infrastructures are needed. However, the lack

of an infrastructure, storage difficulties and related safety issues

can only be solved unilaterally by introducing an on-board fuel

processor which converts a commercially available fuel, such as

gasoline or diesel, into hydrogen such that it is available on de-

mand [3,4].

The development of portable fuel processors has however for

several years received significant attention in relation to the in-

crease in worldwide fuel cell development activities. The research

and development work on fuel processors for fuel cells could be

classified according to different approaches (or methodologies)

and fuel selection. The most considered fuels are alcohols [5,6],

gasoline [7], diesel [3,4], and single or mixture of heavy-hydrocar-

bons as surrogates for (conventional) petroleum based liquid fuel

[6,810] and jet fuel [11]. There are three established methods

for reforming fuels: steam reforming (SR) [5,8,12,13], partial oxida-

tion (POX) [11] and auto-thermal reforming (ATR) [3,4,7,9,1113].

All of the mentioned methods produce a syngas mixture; however

the difference in reaction temperatures and oxidants yields differ-

ent CO concentrations in the syngas mixture. The H2 production

generally decreases in the order of steam reforming, auto-thermal

reforming, and partial oxidation.

In this work, we investigate a combined fuel reforming and

SOFC system and present an in depth analysis and evaluation of

this system that utilizes n-dodecane as the surrogate for diesel

fuels. Most previous studies [35,12] consider low-temperature

fuel cells due to its vehicle-oriented purpose, such as the Auxiliary

0196-8904/$ - see front matter 2011 Elsevier Ltd. All rights reserved.doi:10.1016/j.enconman.2011.05.009

Corresponding author.

E-mail address: [email protected] (J.N. Chung).

Energy Conversion and Management 52 (2011) 32143226

Contents lists available at ScienceDirect

Energy Conversion and Management

j o u r n a l h o m e p a g e : w w w . e l s e v i e r . c o m / l o c a t e / e n c o n m a n
http://dx.doi.org/10.1016/j.enconman.2011.05.009mailto:[email protected]://dx.doi.org/10.1016/j.enconman.2011.05.009http://www.sciencedirect.com/science/journal/01968904http://www.elsevier.com/locate/enconmanhttp://www.elsevier.com/locate/enconmanhttp://www.sciencedirect.com/science/journal/01968904http://dx.doi.org/10.1016/j.enconman.2011.05.009mailto:[email protected]://dx.doi.org/10.1016/j.enconman.2011.05.009


2/13

Power Unit for heavy trucks. However, our application is not re-

stricted by the operating temperature of the fuel cell system. Car-

bon monoxide, the second most produced chemical species in thereformate, should be either converted into hydrogen by watergas

shift (WGS) reaction or removed by a purification process for Pro-

ton exchange membrane fuel cells (PEMFC) to prevent CO poison-

ing of the anode catalyst [1]. On the contrary, it is unnecessary to

have another CO converting or purifying process for a SOFC. Carbon

monoxide can be utilized in the anode of a SOFC as a fuel by either

direct oxidation or WGS reaction to produce more hydrogen. In this

paper, we focus on the application of a SOFC in an air-deficient

environment such as in the aerospace and under-water applica-

tions, rather than on lowering the operating temperatures. It is

noted that for aerospace and under-water applications, the avail-

able space is a main limiting factor. Selecting the most suitable

reforming technology depends on both the application require-

ments and the available design options. The steam reforming is

adopted here as no oxygen is required and it facilitates the highest

hydrogen production [12,13].

In this work, an analytical, parametric study is performed to

evaluate the feasibility and performance of a combined fuel

reforming and SOFC system that utilizes n-dodecane as the surro-

gate for diesel fuels. Specifically the effects of adding the anode off-

gas recycling and recirculation components and the CO2 absorbent

unit are investigated.

2. Thermodynamic and electro-chemical model

2.1. Combined fuel reforming and SOFC system

Fig. 1 shows the process flow diagram of the combined system.The two major components, an external-reformer unit and a SOFC

unit, are coupled to convert the liquid hydrocarbon fuels to elec-

trical power. The entire reforming unit consists of a direct inter-

nal reforming component inside the SOFC, an external reformer,four heat exchangers, a carbon dioxide adsorbent, a flue gas con-

denser and an anode off gas (AOG) recycling unit. The AOG has a

very high temperature and generally also contains high water va-

por content due to the fact that water vapor is the only product of

the electrochemical reaction which produces the electrical power

in a SOFC. The management of the AOG is of high importance to

the overall system efficiency and maintaining the designed steam

to carbon ratio (S/C). With an appropriate amount of recycled

AOG, it is possible to maintain the prescribed steam to carbon ra-

tio value at the inlet of the external reformer. To provide a given

steam to carbon ratio (S/C) at the inlet of the external reformer,

the total amount of steam required is solely supplied by the recy-

cled AOG. So the S/C is closely related to AOG recycle ratio. It is

important to note that AOG recycling also accomplish the purpose

of conserving water because only when the AOG does not contain

enough water vapor, water would then be added with the fuel to

maintain the desired steam to carbon ratio. It is obvious that a

compressor is required to accomplish the AOG recycle process.

From the thermodynamics point of view, the high temperature

compression process requires much more work than a low tem-

perature process. Therefore, the AOG is cooled down before the

compressing process. Accordingly, a cooler is used for this pur-

pose. After the compressor, the temperature of the AOG is raised

back by passing it through a recuperator heat exchanger. How-

ever, a certain portion of the AOG is not recycled and is sent di-

rectly to the after-burner instead where it is burned that provides

heat to the reformer in an effort to retrieve more energy. The flue

gases from the after-burner are designed to pre-heat the fuel and

to provide more heat to the reformer.

Nomenclature

AbbreviationAOG anode off gasATR autothermal reformingLHV lower heating valueMCFC molten carbonate fuel cellPEMFC proton exchange membrane fuel cellPOX partial oxidationS/C steam to carbon ratioSOFC solid oxide fuel cellSR steam reformingWGS watergas shiftYSZ yttria-stabilized zirconia

List of symbolsDeff effective gaseous diffusivity (cm

2/s)F Faraday constant (96485.3 C/mol)Fi molar flow rate of labeled i stream (mol/s)G Gibbs free energy (J/mol)i current density (A/cm2)

io exchange current density (A/cm

2

)is limiting current density (A/cm2)

K equilibrium constant (equivalent)L electrode thickness (lm)NC number of carbon in the fuel ()nij number of moles changed for jth species in ith reaction

(equivalent)P pressure (atm)Pi partial pressure of ith species (atm)R gas constant (8.314 J/molK)

Rc recycle ratio ()Rcontact contact resistance (X)r recirculation ratio ()T temperature (K)U fuel utilization factor ()V voltage (Volt)W power (watt)yi molar fraction for ith species ()

Greekei extent of ith reaction (equivalent)e effectiveness ()g efficiency ()mij stoichiometric coefficient for jth species in ith reaction

()q electrical resistivity (X cm)

SubscriptsA anode

AB after-burneract activation overpotentialC cathodeconc concentration overpotentialE electrolyteelect electricityf fuelN Nernst potentialohm Ohmic overpotential

T.S. Lee et al. / Energy Conversion and Management 52 (2011) 32143226 3215


3/13

The molar balance including electro-chemical reactions as well

as chemical equilibrium reactions is considered next. After evaluat-

ing the molar balance, an energy balance is performed for each

component.

2.2. Assumptions and Justification

The following assumptions are made in the analysis:

Steady state operation.

All gaseous phases are ideal gas.

Chemical and thermal equilibrium are achieved due to enough

residence time.

Only hydrogen is electrochemically reacted at the anode inside

the SOFC.

Pressure drop is neglected through the system.

Anode-supported SOFC

Pure oxygen provision for the cathode

No carbon deposition in the entire system.

Complete conversion of the fuel at the external reformer.

Since the current paper is basically dealing with the introduc-

tion of a conceptual energy management system design that is

intended to demonstrate the new concept, functions, and charac-

teristics of the coupled SOFC and reforming system. However, it

is not aimed at the detailed modeling of each component. Each

component is a unit and it is represented by the key characteristics.

So, we have adopted some simplifying ideal conditions as listed

above. The pure oxygen for the cathode is one of them. We believe

that these idealizations would not be likely to change the major

outcomes and conclusion in this paper if more realistic assump-

tions were used. For some special applications where air is not

readily available such as underwater vehicles, and aero and space

applications, pure oxygen is usually used. Since our external refor-

mer utilizes only steam not oxygen, it is the steam reforming of thefuel in our case.

There are possibly some other species formed, but the key and

dominant products are CO and H2 in our reformer. The term an-

ode-supported SOFC denotes the anode is the thickest part among

the anode, electrolyte, and cathode so that the anode can provide

enough mechanical strength for the fuel cell structure. We as-

sumed enough residence time for reactants to react completely.

This is another one of those ideal conditions. We also assumed

no carbon deposition which can be justified for the conditions of

steam to carbon ratio investigated in this work [14,15].

2.3. Molar balance and chemical equilibrium

As shown in Fig. 1, the reformer unit includes two parts, an

external reformer and an after-burner. The pre-heated mixture of

fuel and recycled AOG, mostly water, is passing through the

reforming channels of the external reformer. A reasonable way of

approaching the liquid hydrocarbon reforming modeling is by pre-

dicting the thermodynamic equilibrium composition through

Gibbs free energy minimization calculation under the presence of

catalysis, and given pressure and temperature. In the presence of

catalysis, steam reforming reaction, which is shown in Eq. (1),

takes place and the readjustment of chemical species is achieved

by chemical reaction equilibrium under the given temperature

and pressure.

CnHm n SCH2O ! mCOCO mH2 H2 mCO2 CO2 mCH4 CH4

mH2OH2O 1

Hydrocarbon reforming reaction, Eq. (1), is assumed irreversible

and relatively fast under the certain catalyst as well as completely

converted. In the hydrocarbon reforming chemical reaction, stoi-

chiometry coefficients for the product chemicals are arbitrary at

this time. Even though stoichiometry coefficients are arbitrary,

we are not dealing with nuclear reaction so each coefficient should

satisfy three atomic balances, i.e. carbon, hydrogen and oxygen

atomic balances. With three atomic balances, we still have an

underdetermined condition, i.e. five unknowns and three equa-tions. Here, chemical equilibrium assumption plays a key role.

Fig. 1. Process flow diagram.

3216 T.S. Lee et al. / Energy Conversion and Management 52 (2011) 32143226


4/13

From the classical thermodynamics, two independent chemical

reaction equations can be composed from one phase (gaseous

phase), five chemical species (CO, H2, CO2, CH4, and H2O) and three

component atoms (C, H, and O) and those reactions are as follow,

CH4 H2OCO 3H2 2

CO H2OH2 CO2 3

The first reaction, Eq. (2), is the well known methane steam

reforming reaction and the second one, Eq. (3), is the well known

watergas shift reaction. This pair of chemical reactions is uniquely

set for the given condition, one phase, five chemical species, and

three atomic components. If we consider carbon deposition then

we should add solid carbon as a deposited material on the catalyst

surface. With solid carbon formation, we need one more chemical

reaction to solve the stoichiometry of steam reforming reaction, Eq.

(1). However, carbonhydrogenoxygen (CHO) species diagram

shows that carbon deposition could be eliminated by providing en-

ough water and/or oxygen into the system [14,15]. For our lowest

steam to carbon ratio operating condition the products are located

underneath the carbon deposition boundary line in the CHO

phase diagram which means that the above no carbon depositionassumption is valid theoretically. Therefore we do not need to con-

sider carbon deposition, so we have a unique set of independent

equilibrium reactions, Eqs. (2) and (3). Let the extents of reactions,

which are defined as dei = dnij/mij, shown in Eqs. (2), (3) be eR1 and

eR2, respectively. Then equilibrium molar fractions are expressed as

follows:

yH2 O Fo;H2O eR1 eR2

Fo 2eR14a

yCH4 Fo;CH4 eR1Fo 2eR1

4b

yH2

Fo;H2 3eR1 eR2

Fo 2eR1 4c

yCO Fo;CO eR1 eR2

Fo 2eR14d

yCO2 Fo;CO2 eR2Fo 2eR1

4e

where Fo and Fo,i denote the inlet total molar flow rate, labeled 3b in

Fig. 1, and the inlet molar flow rate of ith component, respectively.

Once the extents of reactions are obtained, evaluation of the equi-

librium molar fraction is straightforward. Therefore, we need two

equations to be solved simultaneously to find eR1, and eR2 under

the equilibrium condition. These are the chemical equilibrium

equations corresponding to the steam reforming and watergasshift reaction, as given below:

KR1T P

Po

2 y3H2yCOyCH4yH2 O

expDGR1

RT

5a

KR2T P

Po

0 yH2yCO2yH2OyCO

expDGR2

RT

5b

The temperature-dependent equilibrium constant is obtained

by the classical method in which the change in Gibbs free energy

of the reactions is used. After the Gibbs energy difference is ob-

tained, equilibrium molar fractions, Eq. (4), are substituted into

Eq. (5) and then the system of equations, i.e. three atomic balances

and two extent of reactions, is solved by the NewtonRahpsonmethod with a tolerance of 107.

In this work, the direct internal reforming SOFC model is based

on the one proposed by Colpan et al. [15,16]. In this model, CO and

methane are assumed to react with the water vapor for producing

more hydrogen in the anode. However, the total amount of electri-

cal energy produced in the current system comes from watergas

shift reaction and methane steam reforming reaction in addition to

direct oxidation in the anode. Therefore, methane steam reforming

reaction, Eq. (2), watergas shift reaction, Eq. (3), and electrochem-

ical reaction, Eq. (6), occur simultaneously inside the direct inter-

nal reforming section of the SOFC.

H2 1

2O2 ! H2O 6

The extent of electrochemical reaction, eS3, can be expressed

using a molar balance, definition of molar fraction, and recircula-

tion ratio.

eS3 UFo;H2 3eS1 eS2

1 r rU7

Here, ris the recirculation ratio, U is fuel utilization, eSi is reac-

tion coordinates for ith reaction at the SOFC, respectively. Also, Fodenotes the inlet of SOFC anode, labeled 5 in Fig. 1. With the above

assumptions and Eq. (7), molar fractions for all the species at theexit of the anode of the fuel cell are given as below:

yCH4 Fo;CH4 eS1Fo 2eS1

8a

yH2 O Fo;H2O eS1 eS2

Fo;H23eS1 eS2

1rrU

U

h iFo 2eS1

8b

yH2 Fo;H2 3eS1 eS2

Fo 2eS1

1 r1 U

1 r rU8c

yCO Fo;CO eS1 eS2

Fo 2eS18d

yCO2 Fo;CO2 eS2Fo 2eS1

8e

Here, eS1 andeS2 are the only unknowns. With likewise chemical

equilibrium assumed in the external reformer, molar fractions at

the exit of the SOFC anode (Eq. (8)) are evaluated using Eqs. (5a)

and (5b).

2.4. SOFC model

Once the SOFC exit composition is obtained based on thermody-

namic equilibrium, calculation of the amount of energy conversion

from chemical to electrical is straightforward by means of an elec-

tro-chemical relationship. The terminal voltage of the cell can be

obtained after considering all voltage losses; V = VN Vohm

Vact Vconc. Here, the Nernst voltage is calculated as,

VN DG

oTSOFC

2F

RTSOFC2F

lnyH2O;A

yH2 ;A

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiyO2 ;CP=P

oq

0B@

1CA 9

Here, three types of overpotentials, e.g. ohmic, activation and

concentration, are considered and calculated using Eqs. (11),

(13), and (15). The Ohmic overpotential occurs because of resis-

tance to the flow of ions in the electrolyte and resistance to the

flow of electrons through the electrically conductive fuel cell

components.

Vohm iRcontact X

kqkLk 10



5/13

where i is the current density, q is the electrical resistivity (inverse

of electrical conductivity), L is the thickness of electrode, and Rcontactis the contact resistance. Electrical resistivity is given as a function

of temperature for SOFC materials [17] e.g. lanthanum manganite

for cathode, YSZ for electrolyte, and Ni/ZrO2 for anode; qC =

0.008114 exp(600/T) (X cm), qE = 0.00294 exp(10,350/T) (X

cm), qA = 0.00298 exp(1392/T) (X cm). Please note that Ref.

[15] provides the resistivity for Ni/ZrO2

which we used for Ni/YSZ

as our anode material as the two are very close in resistivity values.

Ref. [18] also adopted the same approximation.

Activation overpotential is associated with the sluggish elec-

trode kinetics occurring at the electrodeelectrolyte interfaces.

ButlerVolmer equation with 0.5 transfer coefficient is used for

the activation overpotential.

Vact Vact;A Vact;C

RTSOFC

Fsinh

1 i

2io;A

RTSOFCF

sinh1 i

2io;C

11

where F is the Faraday constant, io,A and io,C are exchange current

density for anode and cathode, respectively. These exchange current

densities are a function of temperature, which are related to the

charge transfer resistance; ioi =Ai exp(Bi/RT). Coefficients Aiand Bi are determined from a curve fitting of data from the litera-

ture [18]; Aa = 1.2903 107 A/cm2, Ac = 3.9255 10

6 A/cm2, Ba =

151.532 kJ/mol and Bc = 149.395 kJ/mol. Concentration overpoten-

tial is caused by concentration gradients due to the resistance to

mass transport through the electrodes and interfaces [19].

Vconc Vconc;A Vconc;C

RTSOFC

2Fln 1

i

is;A

RTSOFC2F

ln 1 PH2 i

PH2Ois;A

RTSOFC

4Fln 1

i

is;C

12

where is,A and is,C are limiting current densities for anode and cath-ode, respectively and shown below:

is;A 2FPH2 Deff;ARTSOFCLA

and is;C 4FPO2 Deff;C

PPO2P

RTSOFCLC

13

where Deff is the effective gaseous diffusivity through the electrode.

In this study, the SOFC is assumed anode supported and cell geom-

etry is provided in Table 1. As shown in Fig. 1, pure oxygen is

supplied to the cathode. It is estimated typical concentration over-

potential for the anode is several orders of magnitude larger than

that of the cathode, therefore, it is a reasonable assumption that

no mass transfer limitation for the cathode. Typical the anode effec-

tive gaseous diffusivity for an anode supported fuel cell is given by

[19]

Deff;A 1:3103 105 T1:5SOFC 0:26382 cm

2=s 14

2.5. Water management and AOG recycle

As a part of the molar balance, the last step for the AOG recycle

unit design is the determination of the recycle ratio. The amount of

recycled AOG is manipulated to maintain the desired S/Cratio. The

AOG recycle ratio is defined as the recycled AOG to pre-recycled

AOG, Rc = F2/F1. In this work, it should be noted that the steam to

carbon ratio (S

/C

) is defined at the inlet of the reformer instead

of at that of the SOFC anode and is given as S/C= F2,H2O/

(F2,CH4 + FfNC) where, Ff and NCdenote molar flow rate of freshly

added fuel, labeled f in Fig. 1, and the number of carbon in the fuel

e.g. for dodecane NC= 12. Substituting the definition of recycle ra-

tio into that of steam to carbon ratio yields,

RcS=CFfNC

F1;H2O S=CF1;CH415

The numerator of Eq. (15) denotes the required amount of

steam due to newly added fuel for the given conditions and the

denominator represents the amount of excess steam available be-

fore recycling. The range of recycle ratio for AOG is between zero

and one. It is obvious that the recycle ratio is always greater than

zero because we manipulate it to regulate the S/C and the S/C isgreater than zero. The recycle ratio should be less than one, the to-

tal recycle case, to prevent mass accumulation on the system. The

un-recycled AOG comprises mostly water vapor but also contains a

small fraction of un-reacted hydrogen as well as carbon monoxide

which can be utilized in the after-burner.

2.6. Energy balance

With the energy balance analysis, the temperature distribution

in the system can be determined. Basically, it is assumed that all

components are adiabatic with the surroundings except the SOFC,

external reformer, and afterburner. Since we are focusing on the

whole system analysis rather than individual components, a single

component temperature is assigned for a unit such as the reformer.Therefore, a detailed model for the reformer was not needed as the

reformer is just a point (a unit) from the system point view, a rep-

resentative temperature would serve the purpose.

First, the first law of thermodynamics is applied to the six com-

ponents (adsorbent, condenser, cooler, injector, pre-heater, and

recuperator) for energy balance calculations. e-NTU method [18]

has been applied for the pre-heater and the recuperator. For the

recuperator and the fuel pre-heater, we assume that the heat

exchangers are cross-flow type with both fluids mixed, so the fol-

lowing effectiveness function is used for the e-NTU applications

[20].

e NTU

NTU1expNTU

CrNTU1expCrNTU

116

where e = q/qmax = q/Cmin(Th,iTc,i), Cr = Cmin/Cmax, and NTU =AUo/

Cmin. For sample calculations, we assume a representative value of

four for the NTU as the overall heat transfer coefficient, Uo, is usu-

ally moderate for gases and the heat exchanger size is a medium

scale for aerospace and underwater applications. Also the kinetic

and potential energy terms are neglected in this work. For calcula-

tion of the enthalpy difference, specific heats for an ideal gas mix-

ture are integrated with respect to temperature. The specific heat

data is taken from the literature as a function of temperature

[21]. Regarding the chemical reaction, heats of reactions for the

SOFC, external reformer, CO2 adsorbent, and after-burner are taken

into account as well as the sensible heat. In this work, the value of

heat of adsorption is obtained from the literature 17000 J/mol

[22]. It is assumed that the heat of combustion in the after-burneris completely transferred to the external reformer to provide the

Table 1

SOFC unit data.

Operating temperature (TSOFC) C Various

Pressure of the cell (Pcell) bar 1

Active surface area (A) m2 1

Exchange current density of anode (ioa) A/cm2 Function of T

Exchange current density of cathode (ioc) A/cm2 Function of T

Effective gaseous diffusivity through anode (Daeff) cm2/s Function ofT

Effective gaseous diffusivity through cathode (Dceff) cm2/s N/A

Thickness of anode (La) lm 500

Thickness of electrolyte (Le) lm 10

Thickness of cathode (Lc) lm 50



6/13

heat for the reforming reaction which is strongly endothermic.

Since the AOG recycle ratio is set to maintain the prescribed S/Cra-

tio, so the un-recycled AOG in the after-burner may not be able to

supply enough heat to theexternal reformer. To make up the energy

shortage, additional fuel is added to the after-burner to meet the

demand. Once the energy balance is set, evaluation of temperatures

is carried out by the NewtonRahpson method with a 107

tolerance.

3. Results and discussion

First, we compare our results with those in the open literature

in an effort to verify the current model and methodology. The clos-

est comparison we can find in the literature is the experimental

work reported by Sasaki et al. [23]. Their system does not include

an anode off-gas recycle system but only include an external refor-

mer and a SOFC. So we have modified our simulationcode to match

their system, simply by eliminating the recycle loop and adding

some minor changes (described in Table 2). The system parameters

and operating conditions used for the comparison are given in the

Table 2. The comparison between the current model and the exper-

imental work is based on the SOFC performance that is shown inFig. 2. We found that for the currentvoltage characteristics of

the SOFC, there is a good agreement between our model and the

experimental work by Sasaki et al. [23]. Based on the close compar-

ison, we believe that our model and the methodology are realistic

and correctly implemented.

3.1. SOFC performance

For the purpose of demonstrating the applicability and effec-

tiveness of the proposed combined fuel reforming and SOFC sys-

tem, sample calculations were performed using n-dodecane as a

representative diesel fuel and typical system conditions. As men-

tioned in the previous section, pure oxygen is assumed to be the

oxidizer and supplied to the cathode. The oxygen/fuel (n-dode-cane) molar ratio used in our calculation is 37/2 that is the stoichi-

ometric ratio. First the basic performance characteristics of the

SOFC are examined under various system parameters. To facilitate

a parametric study, the baseline case, defined in Table 3, is used as

a reference.

Using the baseline case, the effects of varying the S/C, SOFC

operating temperature, AOG recirculation ratio and CO2 adsorption

percentage on the SOFC performance are computed and shown in

Figs. 36 for comparison. In Fig. 3, the focus is on the effects of dif-

ferent steam to carbon ratios (S/C) while keeping all other param-

eters equal to those of the baseline case. Three different S/C ratios

of 2, 3 and 4 are evaluated and their effects on the voltage, power

density, and limiting current density are plotted. Based in Fig. 3, ifthe steam to carbon ratio is increased, the terminal voltage, peak

power output, and fuel consumption rate for a constant voltage

all decrease monotonically. The same trend is also reported by

Kang et al. [8] where the authors mentioned that the measured

SOFC terminal voltage decreased from 0.733 to 0.726 and then to

0.713 as the S/C was increased from 0.625 to 1.25 and then to

2.5, respectively. As defined in the previous section, the S/C ratio

in this paper is set at the inlet of the reformer rather than at that

of the SOFC, which facilitates the involvement of the reformer on

the actual S/Cratio to the SOFC. The AOG recycle percentage is ad-

justed to supply enough water vapor for accomplishing the steam

to carbon ratio specified at the inlet of the reformer. The un-recy-

cled AOG is combusted in the after-burner to capture the un-re-

acted hydrogen and methane left in the stream. In general, as theS/C is increased more steam is available in the reformer. So, for

the reformate at the exit of the reformer, the hydrogen molar frac-

tion would drop while that of the water vapor would increase that

causes the Nernst voltage potential to decrease according to Eq. (9).

This effect is clearly demonstrated in Fig. 2 with the highest open

circuit voltage corresponding to the lowest S/C. When the S/Cratio

is raised, more hydrogen would be produced in the reformer and

fed into the SOFC due to an increase in the amount of water vapor

available for the water gas shift reaction. Further more, according

to Eq. (13), a lower hydrogen molar fraction at the SOFC anode inlet

would result in a lower limiting current density. This explains why

the limiting current density is lowered when the S/C is increased.

As shown in Fig. 3, the terminal voltage curves are relatively close

among the three cases when the current density is less than 0.35 A/cm2. While the limiting current density really sets the three cases

Table 2

System parameters used for comparison with experiment of Sasaki et al. [8].

Fuel n-Dodecane

Cathode flow Air

Fuel cell

Type SOFC

Anode Lanthanum manganite (140 lm)

Electrolyte YSZ (10 lm)

Cathode Ni-YSZ (50 lm)

Fuel cell temperature (C) 1000

Reformer Steam reforming

S/C 2.5

Reformer temperature (C) 750

Fuel utilization factor 0.1

Oxygen utilization factor 0.02

Exchange current density of anode (ioa)

(A/cm2)

0.1

Exchange current density of cathode (ioc)

(A/cm2)

0.07

Current density, i [A/cm2]

Voltage,V[V

olt]

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.40.5

0.55

0.6

0.65

0.7

0.75

0.8

0.85

0.9

0.95

1

Literature (Sasaki et al.)This work

n-dodecane fuel

S/C = 2.5

T = 1000oC

Fig. 2. Currentvoltage characteristics comparison with experimental results bySasaki et al. [8].

Table 3

System parameters and operating conditions for the baseline case.

SOFC operating temperature (C) 850

Steam to carbon ratio (S/C) () 3

AOG recirculation ratio () 0

CO2 adsorption percentage (%) 0

External reformer temperature (C) 550

Fuel utilization factor () 0.85

SOFC active area (m2) 1



7/13

apart that results in quite different peak power densities. The low-

er steam to carbon ratio case has the higher peak power density

due to the higher limiting current density.

Fig. 4 shows how the SOFC operating temperature affects theterminal voltage as well as the current density. First, we found that

the maximum power density increases with increasing SOFC oper-

ating temperature and so does the limiting current density. Similar

to the previous case (Fig. 3), the peak power density is primarily

dependent on the limiting current density that is proportional tothe SOFC operating temperature. However, the trend of open-cir-

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.0

0.2

0.4

0.6

0.8

1.0

S/C = 2

S/C = 3

S/C = 4

Voltage

Power density

Voltage,

V[volts]

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8

Powerdensity,

[W/cm

2]


Fig. 3. Polarization curves for different steam to carbon ratios, SOFC temperature 850 C, reformer temperature 550 C, no recirculation, no CO2 adsorption, 0.85 fuel

utilization factor, and 1 m2 for active area.

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.80.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.0

0.2

0.4

0.6

0.8

1.0

TSOFC

= 950 [oC]

TSOFC

= 850 [oC]

TSOFC

= 750 [oC]

Voltage

Power density

Voltage,

V[volts]

Powerdensity,

[W/cm

2]


Fig. 4. Polarization curves for different SOFC temperatures, S/C3, reformer temperature 550 C, no recirculation, no CO2 adsorption, 0.85 fuel utilization factor, and 1 m2 for

active area.



8/13

cuit voltages differs from that of the previous case. The open-cir-

cuit voltage or the Nernst voltage, Eq. (9), is inversely proportional

to the SOFC operating temperature that results in the highest open-circuit voltage for the lowest SOFC operating temperature. Based

on Eq. (9), we also note that all the overpotentials (voltage losses

due to irreversibilities) are also inversely proportional to the SOFC

operating temperature. This is the reason why the terminal voltagedecreases faster for a lower SOFC operating temperature which re-

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.0

0.2

0.4

0.6

0.8

1.0

r = 0.0

r = 0.1

r = 0.2

Voltage

Power density

Powerdensity,

[W/cm

2]


Voltage,

V[volts]

0.0 0.1 0.2 0.3 0.4 0.5 0.6

Fig. 5. Polarization curves for different AOG recirculation, S/C3, SOFC temperature 850 C, reformer temperature 550 C, no CO2 adsorption, 0.85 fuel utilization factor, and

1 m2 for active area.

Powerdensity,

[W/cm

2]


0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.80.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.0

0.2

0.4

0.6

0.8

1.0

no CO2 Adsorption15% CO

2Adsorption

30% CO2

Adsorption

Voltage

Power density

Voltage,

V[volts]

Fig. 6. Polarization curves for different CO2 capture, S/C3, SOFC temperature 850 C, reformer temperature 550 C, no recirculation, 0.85 fuel utilization factor, and 1 m2 for

active area.



9/13

sults in a smaller limiting current density. For example, based on

Eq. (12), concentration overpotential is a strong function of the

limiting current density and inversely proportional to it. According

to Eqs. (16) and (17), a higher SOFC temperature yields a larger

limiting current density that results in lower concentration

overpotentials.

Fig. 5 demonstrates how the terminal voltage and power den-

sity change with the AOG recirculation ratio. It is recalled that

the recirculation of AOG puts more water vapor back to the SOFC

anode inlet and its effect is very similar to an increase of the S/C.

As the recirculation ratio is increased, terminal voltage, peak power

output, fuel consumption rate for a constant voltage, current den-

sity at peak power, and the maximum limiting current density all

decrease, that is very similar to the results shown in Fig. 3 where

the steam to carbon ratio is increased. It may be deduced that

increasing the recirculation ratio represents increasing the steam

to carbon ratio at the SOFC inlet due to high water vapor content

in recirculation stream.

The last system parameter examined is the percentage of car-

bon dioxide in the reformate captured by the adsorbent and the re-

sults are illustrated in Fig. 6. As shown in the process flow diagram,

Fig. 1, the carbon dioxide adsorption process is located between

0.20

0.25

0.30

0.35

0.40

0.45

0.50

TSOFC

= 750oC

TSOFC

= 800oC

TSOFC

= 850oC

TSOFC

= 900oC

TSOFC

= 950oC

S/C = 2

S/C = 3

S/C = 4

0.3 0.4 0.5 0.6 0.7 0.8

Peakpowerden

sity[W/cm

2]

Current density @ peak power, ipeak [A/cm2]

Fig. 7. The peak power density for different steam to carbon ratios as well as SOFCtemperature for reformer temperature 550 C, no recirculation, no CO2 adsorption,

0.85 fuel utilization factor, and 1 m2 for active area.

0.3 0.4 0.5 0.6

0.20

0.25

0.30

0.35

0.40

TSOFC

= 750oC

TSOFC

= 800oC

TSOFC

= 850oC

TSOFC

= 900oC

TSOFC

= 950oC

Peakpowerdensity[W/cm

2]


r = 0

r = 0.1

r = 0.2

Fig. 8. The peak power density for different recirculation ratios as well as SOFC

temperature for reformer temperature 550 C, S/C 3, no CO2 adsorption, 0.85 fuelutilization factor, and 1 m2 for active area.

0.4 0.5 0.6 0.7 0.80.20

0.25

0.30

0.35

0.40

0.45

0.50

0.55

TSOFC

= 750oC

TSOFC

= 800oC

TSOFC

= 850oC

TSOFC

= 900oC

TSOFC

= 950oC

Peakpowerdensity[W/cm

2]


no CO2

removal

15% CO2

removal

30% CO2

removal

Fig. 9. The peak power density for different CO2 removal percentage as well asSOFCtemperature for reformer temperature 550 C, S/C3, no recirculation, 0.85fuel

utilization factor, and 1 m2 for active area.

750 800 850 900 9500.1

0.2

0.3

0.4

0.5

0.6

S/C = 2

S/C = 3

S/C = 4

S/C = 5

PeakPowerde

nsity,

[W/cm

2]

SOFC Operating Temperature, TSOFC [oC]

Fig. 10. Peak power density with different S/C ratio for reformer temperature

550 C, no recirculation, 15% CO2 adsorption, 0.85 fuel utilization factor, and 1 m2for active area.



10/13

the SOFC and the reformer unit. It is assumed that a certain fraction

of the CO2 in reformate is removed before entering the SOFC anode

and this CO2 capture causes a molar fraction increase for the other

chemical species in the reformate. The effect of the CO 2 capture is

opposite to that of recirculation of AOG in hydrogen molar fraction

point of view. Recirculation of AOG dilutes hydrogen concentration

at the SOFC anode inlet, while CO2 capture enhances the hydrogen

concentration at the SOFC inlet. Considering Eq. (13), a higher

hydrogen molar fraction for the SOFC causes a higher limiting

current density that in turn results in a lower concentration over-

potential. As shown in Fig. 6, a higher limiting current density

could be achieved under a higher CO2 capture percentage. In sum-

mary, CO2 removal yields an opposite trend to those of AOG recir-

culation and higher reformer inlet steam to carbon ratio.

3.2. Peak power density

In engineering applications, the peak power density is of the

most concern. Therefore, here we focus on this main performance

indicator. The peak power density vs. the corresponding current

density is plotted as a function of different S/Cratios, AOG recircu-

lation ratios, CO2 capture percentages as well as SOFC operating

temperatures in Figs. 79. For each figure, two parameters are var-

ied that results in one set of curves with different symbols. For

example, in Fig. 7 the S/C ratio and the SOFC temperature are var-

ied with all other parameters and conditions kept identical to the

baseline case. However, three curves represent different S/Cratios,

while the symbols denote the SOFC temperatures. It may be ob-

served from the figure that a relatively linear dependence between

the peak power density and the corresponding current density. It is

clear that the maximum SOFC performance is associated with the

highest SOFC operating temperature and the lowest S/C ratio. In

Fig. 8, the AOG recirculation ratio and the SOFC temperature are

varied with all other parameters and conditions kept identical to

the baseline case. Similar to Fig. 7, a linear relationship is observed.

The peak power density and the corresponding current density

move towards the upper right corner as the recirculation ratio de-

creases and/or SOFC operating temperature increases. However,

the variation of AOG recirculation ratio on the peak power density

and the corresponding current density is not as significant compar-

ing with that of the steam to carbon ratio. The reason that an in-

crease in the AOG recirculation ratio would cause the peak

power density to decrease instead of increasing is mainly due to

the fact that AOG typically contains $90% water vapor and only

750 800 850 900 9500.3

0.4

0.5

0.6

0.7

0.8

0.9

S/C = 2

S/C = 3

S/C = 4

S/C = 5Currentdensity@

peakpow

er,ip

eak

[A/cm

2]


Fig. 11. Current density @ peak power with different S/C ratio for reformer

temperature 550 C, no recirculation, 15% CO2 adsorption, 0.85 fuel utilization

factor, and 1 m2 for active area.

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.80.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50

0

10

20

30

40

50

60

70

80

TSOFC

= 750 [oC]

TSOFC

= 850 [oC]

TSOFC

= 950 [oC]

System efficiency

Power density

Powerdensity,

[W/cm

2]


SystemEfficiency,LHV

[%]

Fig. 12. System efficiency based on lower heating value of the fuel and power density for different SOFC temperatures, S/C 3, pre-reformer temperature 550 C, norecirculation, 15% CO2 adsorption, 0.85 fuel utilization factor, and 1 m

2 for active area.



11/13

$5% hydrogen. AOG recirculation essentially creates higher steam

to carbon ratio that causes the peak power density to decrease that

is demonstrated in Fig. 10. In Fig. 9, the CO2 removal percentage

and the SOFC temperature are varied with all other parameters

and conditions kept identical to the baseline case. Consistent with

Figs. 7 and 8, a linear trend is shown. The peak power density and

the corresponding current density approach the upper right corner

as the CO2 removal percentage is increased and/or SOFC operating

temperature gets higher.

Figs. 10 and 11 demonstrate the steam to carbon ratio and SOFC

temperature effects on peak power density and current density at

peak power density for typical conditions with no AOG recircula-

tion and 15% CO2 capture. Since effects of changing AOG recircula-

tion and CO2 capture percentage have been evaluated in Figs. 8 and

9, any AOG recirculation and CO2 capture percentage would serve

the purpose. 15% and zero CO2 capture would be as good as any

other selection.

Both the peak power density and current density at peak power

are increasing with increasing SOFC temperature as well as with

decreasing steam to carbon ratio. Both also share the same trend

that the SOFC temperature dependency is more significant at lower

steam to carbon ratios.

3.3. System thermal efficiency

As mentioned above in the energy balance section, the un-recy-

cled AOG plays an important role in the overall energy budget for

the external reformer since the un-recycled AOG is combusted in

the after-burner and the combustion process provides heat for

the strongly endothermic steam reforming reaction in the refor-

mer. In general, the heat supplied by the combustion of the un-

recycled AOG is not enough for the need of the reformer. Based

on the molar balance and energy balance under chemical equilib-

rium and given system operating conditions, this heat deficit can

be determined and is made up by consuming additional fuel in

the after-burner (second term inside the parentheses in Eq. (17)).

Therefore, this additional fuel required for the steam reforming

reaction should be included when defining the system thermal effi-

ciency as follows:

gLHV WSOFC;elect

Ff Ff;ABLHVf17

Fig. 12 demonstrates how the system efficiency varies as a func-

tion of the SOFC temperature as well as the current density. From

Fig. 11, it is observed that the thermal efficiency generally starts

out at a very high value (>70%) near the open circuit voltage and

then monotonically decreases to around 55% where the maximum

power densityoccurs before dropping steeply to zero at thelimiting

750 800 850 900 95040

42

44

46

48

50

52

54

56

58

60

S/C = 2

S/C = 3

S/C = 4

S/C = 5

SystemEfficiency@P

eakPower,LHV[%]


Fig. 13. System efficiency based on lower heating value of the fuel at peak power

density with different S/C ratio for pre-reformer temperature 550 C, no recircu-

lation, 15% CO2 adsorption, 0.85 fuel utilization factor, and 1 m2 for active area.

TSOFC

= 850oC

S/C = 4

no recirculation

noCO2

capture

Fuel utilization factor: 0.85TReformer

= 550oC

TReformer

= 700oC

TReformer

= 550oC

TReformer

= 700oC


Powerdens

ity,[W/cm2]

Voltage,

V[Volts]

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0.50

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

Fig. 14. Polarization curves for S/C= 4, SOFC temperature 850 C, reformer temperature 700 C, no recirculation, no CO2 adsorption, 0.85 fuel utilization factor, and 1 m2 foractive area.



12/13

current density condition. From an engineering application point of

view, the proposed combined fuel reforming and SOFC system with

a range of efficiency between 50% and 70% is considered very attrac-

tive. It is noted that only for the case of 750 C SOFC temperature

under vanishingly small current densities, the efficiency does not

follow the monotonic decreasing trend. As a matter of fact, it shows

a peak efficiency of 76.6% at 0.0167 A/cm2 that is due to additional

fuel consumption (Ff,AB

> 0 in Eq. (17)) at the after-burner required

for low SOFC temperature and/or low current density.

Fig. 13 demonstrates the steam to carbon ratio and SOFC tem-

perature effects on the system efficiency at the peak power density

for typical system conditions with no AOG recirculation and 15%

CO2 capture. The system efficiency under the peak power condition

generally increases sharply with the SOFC temperature up to

850 C and then levels off and moves up slowly to reach the max-

imum at 900 C. After that it decreases slightly as the SOFC temper-

ature increases further to 950 C. Generally, a lower steam to

carbon ratio yields a lower system efficiency but the dependence

of the system efficiency on the S/C ratio is not strong except for

S/C= 2. It is interesting to note that in Figs. 10 and 11 both the peak

power and the current at peak power decrease with increasing S/C

ratio while the system efficiency at peak power shows an opposite

trend with the S/C ratio. This is mainly due to the fact that in the

definition of the system efficiency (Eq. (17)), the numerator and

the denominator are determined by the power density and current

density, respectively, but the dependence on the S/C ratio is rela-

tively stronger for the current density than the power density as

shown in Figs. 10 and 11.

It is noted that all the above results are based on that the refor-

mer is operating at 550 C and some questions arise on whether a

thermal dynamic equilibrium state can be reach at this tempera-

ture. For a kilowatt-scale autothermal reformer for the production

of hydrogen from heavy hydrocarbons, Liu et al. [9] reported that

reformate composition obtained from the experiment is well

matching with thermodynamic calculation at 590 C. In order to

show that the reformer temperature would not affect the major

trends and outcome, we made an additional calculation for the re-former temperature at 700 C. The other system conditions are

kept the same as the 550 C case as follows: SOFC temperature

850 C, S/C= 4, no recirculation, no CO2 adsorption, 0.85 fuel utili-

zation factor, and 1 m2 for active area.

The results are given in Figs. 14 and 15. From the figures pro-

vided, we can see that the fuel cell performance (voltage and

power density) are almost identical between the 550 C and

700 C cases. While, the system efficiency is lower by about 15%

for the 700 C reformer temperature that is basically due to the to-

tal amount of heat available from the AOG for reformer operation is

limited by the SOFC.

As mentioned above, the AOG recycle percentage is a function of

the S/C ratio. Fig. 16 provides this relationship as a function of the

SOFC operating temperature. As expected the percentage increases

with increasing S/Cratio, but the relationship is independent of the

SOFC operating temperature. Fig 16 also indicates that for the prac-tical range ofS/Cratios up to 5, there is no need to add water from

an external source.

4. Conclusions

A parametric study for a combined fuel reforming and SOFC sys-

tem with AOG is presented for thermal management applications.

The model and the methodology are verified by a favorable com-

parison with an independent experimental work from the open

literature.

Based on a parametric study, we found that in general the SOFC

terminal voltage decreases with increasing current density and

after reaching the maximum power density it sharply drops to zero

at the limiting current density condition. Whereas the SOFC powerdensity increases with increasing current density until the maxi-

mum is reached and after that it decreases steeply to zero at the

limiting current density point. In general, the open-circuit voltage,

the terminal voltage, the peak power density, and the limiting cur-

rent density all increase with increasing S/C ratio, increasing AOG

recirculation ratio, decreasing CO2 adsorption percentage, and

decreasing SOFC temperature, respectively. It is also found that

the AOG recycle ratio increases with increasing S/C ratio and for

most applications, the addition of AOG recycle eliminates the need

for external water consumption.

For most of the operating conditions investigated, the system

efficiency starts at around 70% and then monotonically decreases

to an average of 50% at the peak power density before dropping

down to zero at the limiting current density point. From an engi-neering application point of view, the proposed combined fuel

TReformer

= 700oC

TReformer

= 550oC


SystemEfficienc

y,LHV[%]

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4 0.45 0.50

10

20

30

40

50

60

70

80

Fig. 15. System efficiency curves for S/C= 4, SOFC temperature 850 C, reformer

temperature 700 C, no recirculation, no CO2 adsorption, 0.85 fuel utilization factor,and 1 m2 for active area.


RecyclePercent@P

eakPower,[%]

750 800 850 900 950

66

68

70

72

74

76

78

80

82

84

S/C = 2

S/C = 3

S/C = 4

S/C = 5

Fig. 16. AOG recycle percentage based on peak power density with different S/C

ratio for a pre-reformer temperature of 550 C, no recirculation, 15% CO2 adsorp-tion, 0.85 fuel utilization factor, and 1 m2 for active area.



13/13

reforming and SOFC system with a range of efficiency between 50%

and 70% is considered very attractive.

Our results point out that for the SOFC at 900 C, the steam to

carbon ratio at 5 and no AOG recirculation, the system efficiency

peaks. Also the peak power density increases with decreasing

CO2 capture. Therefore, it is suggested that the optimal system is

the one where the SOFC operates around 900 C with S/C ratio

higher than 3, maximum CO2

capture, and minimum AOG

recirculation.

Acknowledgements

This research was supported by the NASA Hydrogen Research

for Spaceport and Space Based Applications at the University of

Florida (Grant number NAG3-2930). The support by the Andrew

H. Hines Jr./Progress Energy Endowment Fund is also acknowl-

edged. The first author was partially supported by the Korea Sci-

ence and Engineering Foundation Grant funded by the Korea

government (Ministry of Science and Technology) (2005-215-

D00037).

References

[1] EG&G. Fuel cell handbook. 7th ed. West Virginia: U.S. Department of Energy;

2004.

[2] Singhal SC, Kendall K. High temperature solid oxide fuel cells. Oxford, United

Kingdom: Elsevier Advanced Technology Publishing; 2004.

[3] Lindstrm BJ, Karlsson AJ, Ekdunge P, De Verdier L, Hggendal B, Dawody J,

et al. Diesel fuel reformer for automotive fuel cell applications. Int J Hydrogen

Energy 2009;34:336781.

[4] Chrenko D, Couli J, LecoqS, Pra MC, Hissel D. Staticand dynamic modelingof

a diesel fuel processing unit for polymer electrolyte fuel cell supply. Int J

Hydrogen Energy 2009;34:132435.

[5] Emonts B, Bgild J, Hansen S, Jrgensen L, Hhlein B, Peters R. Compact

methanol reformer test for fuel-cell powered light-duty vehicles. J Power

Sources 1998;71:28893.

[6] Rostrup-Nielsen JR. Conversion of hydrocarbons and alcohols for fuel cells.

Phys Chem Chem Phys 2001;3:2838.

[7] Papadias D, Lee SHD, Chmielewski DJ. Autothermal reforming of gasoline for

fuel cell applications: a transient reactor model. Ind Eng Chem Res

2006;45:584158.

[8] Kang I, Kang Y, Yoon S, Bae G, Bae J. The operating characteristics of solid fuel

cells driven by diesel autothermal reformate. Int J Hydrogen Energy

2008;33:6298307.

[9] Liu DJ, Kaun TD, Liao HK, Ahmed S. Characterization of kilowatt-scale

autothermal reformer for production of hydrogen from heavy hydrocarbons.

Int J Hydrogen Energy 2004;29:103546.[10] Shekhawat D, Berry DA, Gardner TH, Haynes DJ, Spivey JJ. Effects of fuel cell

anode recycle on catalytic fuel reforming. J Power Sources 2007;168:

47783.

[11] Shi L, Bayless DJ. Analysis of jet fuel reforming for solid oxide fuel cell

applications in auxiliary power units. Int J Hydrogen Energy 2008;33:106775.

[12] Cutillo A, Specchia S, Antonini M, Saracco G, Specchia V. Diesel fuel processor

for PEM fuel cells: two possible alternatives (ATR versus SR). J Power Sources

2006;154:37985.

[13] Ahmed S, Krumpelt M. Hydrogen from hydrocarbon fuels for fuel cells. Int J

Hydrogen Energy 2001;26:291301.

[14] Sasaki K, Teraoka Y. Equilibria in fuel cell gases. J Electrochem Soc

2003;150(7):A8858.

[15] Colpan CO, Dincer I, Hamdullahpur F. A review on macro-level modeling of

planar solid oxide fuel cells. Int J Energy Res 2008;32:33655.

[16] Colpan CO, Dincer I, Hamdullahpur F. Thermodynamic modeling of direct

internal reforming solid oxide fuel cells operating with syngas. Int J Hydrogen

Energy 2007;32:78795.

[17] Bessette NF, Wepfer WJ, Winnick J. A mathematical model of a solid oxide fuel

cell. J Electrochem Soc 1995;142:3792800.

[18] Chan SH, Low CF, Ding OL. Energy and exergy analysis of simple solid-oxide

fuel-cell power systems. J Power Sources 2002;103:188200.

[19] Kim JW, Virkar AV, Fung K-Z, Mehta K, Singhal SC. Polarization effects in

intermediate temperature, anode-supported solid oxide fuel cells. J

Electrochem Soc 1999;146:6978.

[20] Kays WM, London AL. Compact heat exchangers. New York: McGraw-Hill;

1984.

[21] Smith JM, Van Ness HC, Abbott MM. Introduction to chemical engineering

thermodynamics. 7th ed. New York: McGraw-Hill; 2005.

[22] Ding Y, Alpay E. Adsorption-enhanced steam-methane reforming. Chem Eng

Sci 2000;55:392940.

[23] Sasaki K, Watanabe K, Shiosake K, Susuki K, Teraoka Y. Multi-fuel capability of

solid oxide fuel cells. J Electroceram 2004;13:66975.


Design and Optimization of a Combined Fuel Reforming and Solid Oxide Fuel

Documents

Transcript of Design and Optimization of a Combined Fuel Reforming and Solid Oxide Fuel