Ich

Article history:

Received 8 November 2011

Received in revised form

18 January 2012

The solar thermochemical production of fuels such as

driven looping processes use water and CO2 as the sole feed-

stocks and concentrated solar radiation as the sole energy

source. Looping processes using natural gas [4e7] or coal-

production. This study uses iron/iron oxide redox pairs as the

iron oxides during the first reaction step. Coal-derived Syngas

is then passed through the oxides, reducing them back to iron

during the second reaction step. Since the gaseous products of

* Corresponding author. University of Florida, Department of Mechanical and Aerospace Engineering, 330 MAE-B, Gainesville, FL 32611,USA. Tel.: 1 352 392 9129.

Available online at www.sciencedirect.com

w.

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 0E-mail address: joerg.petrasch@fhv.at (J. Petrasch).hydrogen or Syngas using metal/metal oxide looping

processes [1e3] is considered an interesting route to carbon-

neutral fuels. The concept shows promise in helping to

satisfy a growing global energy demand, reducing oil price

volatility, and mitigating anthropogenic climate change. Solar

reactive material [10]. This process is capable of producing

significantly higher purity hydrogen than conventional coal

gasification and subsequent water gas shift [11,12]. Another

advantage is that the process avoids gas-phase separation.

Metallic iron is oxidized by steam, producing hydrogen and1. Introduction derived Syngas [8,9] as the reducing agent constitute animportant stepping-stone toward carbon-neutral hydrogenAccepted 19 January 2012

Available online 23 February 2012

Keywords:

Looping cycle

Hydrogen

Iron oxide

Thermochemical

Syngas

Coal0360-3199/$ e see front matter Copyright doi:10.1016/j.ijhydene.2012.01.074An incremental thermodynamic equilibrium model has been developed for the chemical

reactions driving a clean, hydrogen producing iron/iron oxide looping cycle. The model

approximates a well-mixed reactor with continuous reactant gas flow through a stationary

solid matrix, where the gas residence time is long compared to time constants associated

with chemical kinetics and species transport. The model, which computes the theoretical

limit for steam-to-hydrogen conversion, has been experimentally validated for the

oxidation reaction using an externally heated, 21 mm inner diameter, tubular fluidized bed

reactor. Experiments were carried out at 660 and 960 C with steam flow rates ranging from

0.9 to 3.5 g/min. For small flow rates, i.e., for long residence times, the experimentally

observed cumulative steam-to-hydrogen conversion approaches the theoretically pre-

dicted conversion. At a 960 C operating temperature, the measured hydrogen yield

approaches the theoretical limit (experimental yields are always within 50% of the theo-

retical limit), and the yield is insensitive to variations in the steam flow rate. In contrast,

the measured hydrogen yield deviates significantly from the theoretical limit at a 660 C

operating temperature, and strong variations in hydrogen yield are observed with varia-

tions in steam flow rate. This observation suggests that the reaction kinetics are signifi-

cantly slower at lower temperature, and the model assumption is not satisfied.

Copyright 2012, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rightsreserved.a r t i c l e i n f o a b s t r a c tDepartment of Mechanical and Aerospace Engineering, University of Florida, Gainesville, FL 32611, USAProduction of hydrogen via anThermodynamic modeling and

A. Singh, F. Al-Raqom, J. Klausner, J. Petras

journal homepage: ww2012, Hydrogen Energy Pron/Iron oxide looping cycle:experimental validation

*

elsevier .com/locate/heublications, LLC. Published by Elsevier Ltd. All rights reserved.

step [2]:

Fe 4=3 H2O/1=3Fe3O4 4=3H2;

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 0 7443the oxidation reaction consist of hydrogen and steam only,

the process can generate highly pure hydrogen through steam

condensation. Furthermore the reduction step yields highly

concentrated CO2 suitable for sequestration.

The suggested two-step process uses the same reactor for

both the reduction and the oxidation reaction. The solid

reactants remain in the reactor and streams of steam and

Syngas are alternatingly fed to the reactors [9]. In contrast, the

three-step steam-iron process [6,11e20] employs two separate

reactors for hydrogen production and iron oxide reduction

[21], hence necessitating the transport of hot solids between

reactors.

To evaluate the theoretical potential of the suggested two-

step process, an open system incremental thermodynamic

equilibrium model is developed for both the hydrogen

production (oxidation) step and the regeneration (reduction)

step. The hydrogen production step is also carried out exper-

imentally to study the validity of the thermodynamic model

and to determine the conditions for its applicability. Roeb

et al. [22] conducted a thermodynamic analysis for two-step

water splitting with mixed iron oxides including nickel-iron-

oxide and zinc-iron-oxide to evaluate the maximum

hydrogen production potential of coating materials using

FactSage software [23]. Their analysis showed that maximum

hydrogen yield is realized when (i) the reduction temperature

is raised to 1300 C, (ii) the water splitting temperature islowered below 800 C, and (iii) the oxygen partial pressureduring reduction is minimized. This is consistent with similar

findings by Singh et al. [9]. Roeb et al. have also validated the

effect of reduction temperature and oxygen partial pressure in

Nomenclature

cp Specific heat capacity, kJ kmol1 K1

G Gibbs free energy, kJ

g0i Reference Gibbs function of species i evaluated,

kJ kmol1

h Enthalpy, kJ kmol1

_m Mass flow rate, kg s1

M Molar mass, kg kmol1

mFe Mass of iron, kg

P Total system pressure, N m2experimental studies. However, they could not experimen-

tally verify the increased hydrogen yield at lower water

splitting temperatures of approximately 800 C. Theyconcluded that kinetics play an important role in the oxida-

tion step. Svoboda et al., have carried out a thermodynamic

study of the potentials and limitations of iron based chemical

looping processes for the production of high purity hydrogen.

They studied the FeeFe3O4 system for cyclic hydrogen

production in the temperature range of 400e800 K [8]. In their

analysis, they have evaluated the hydrogen yield at equilib-

rium for the steam oxidation of pure iron tomagnetite (Fe3O4).

In accordance with Singh et al. and Roeb et al., [9,22] their

theoretical results showed that lower oxidation temperatures

are favorable for attaining higher hydrogen yields. They have

also indicated that at lower temperatures, the reaction is

limited by kinetics.Dh 31:75 kJ=mol at 960 C; (1)

followed by the reduction step:

1=3Fe3O4 2=3CO 2=3H2/Fe 2=3CO2 2=3H2O;Dh 1:25 kJ=mol at 960 C: (2)

High purity hydrogen and magnetite are produced during

the first step. During the second step, magnetite is reduced

back to iron using Syngas as the reducing agent. Coking and

iron carbide formation may occur during reduction. These

products may react with steam in the oxidation processIn the current study, an incremental thermodynamic

equilibrium model is employed to predict the maximum

attainable reaction yields. The model approximates a well-

mixed reactor with continuous reactant gas flow through

a stationary solid matrix where the gas residence time is long

compared to time constants associated with chemical

kinetics. The model is validated experimentally for the

oxidation case using an externally heated tubular fluidized

bed reactor. The current study is limited to the oxidation

reaction of the looping cycle.

2. Thermodynamic analysis

The ideal two-step iron based looping process for the

production of hydrogen consists of the hydrogen production

ni Number of moles for a species i, kmol

mFe,init Initial mass of iron, kg

Pref Reference pressure, N m2

PID Proportional-integral-derivative

R Universal gas constant, kJ kmol1 K1

sLPM Standard liters per minute

T Temperature, K

t time, s

yi,eq Mole fraction at equilibrium

yi Species i mole fractionproducing CO, CO2, and CH4 along with hydrogen. A detailed

analysis of the by-products of the reduction reaction has

been carried out in [9]. In the ideal process hydrogen is

completely consumed in the reduction reaction. However, in

real processes a large fraction of the hydrogen will not react.

The hydrogen and CO2 in the off-gases of the reduction step

may be separated via conventional techniques, such as

pressure swing absorption (PSA) [25] leading to lower purity

hydrogen.

An open system equilibrium model (Fig. 1) for a single

looping reactor is implemented. Small amounts of steam are

added to the system and the ensuing equilibrium reactant gas

mixture is removed from the system. Solid material remains

within the system. Assuming constant temperature and

pressure and ideal gas behavior, the species balance for

a gaseous component follows:

inside an aluminum chamber is used to generate vapor.

Stainless steel wool and a stainless steel screen are inserted in

the aluminum chamber to separate out water droplets and to

ensure dry steam discharges the steam generator. The steam

generator is thermally insulated with fiber glass insulation.

The rate of steam generation is controlled with a pulse-width

modulated signal (PMS) and solid-state relay at a frequency of

2 Hz. A 120 VAC power source provides power to the steam

generator. The steam is superheated to about 200 C bypassing it through an Omega Engineering, 1.37 cm outer

diameter (0.2500 NPT) 200 W in-line gas heater [30] that ismounted vertically and is capable of heating gas from an inlet

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 07444dyi;gasdt

_nyi;gas;in yi;gas;eq

ngas;nsolid;T;P

yi;solid yi;solid;eqngas;nsolid;T;p

(3)

Species considered in the thermodynamic reactor models

are H2, H2O, CH4, C, CO, CO2, Fe, FeO, Fe3O4, Fe2O3, FeCO3, O2,

and Fe3C. Equilibrium compositions for the open system

oxidation/reduction process are calculated using Gibbs free

energy minimization,

yi;gas=solid;eqngas;nsolid;T;p

arg minngas ;nsolid

Gngas;nsolid;T;P

(4)

where

ngas n1;gas;n2;gas;.;nn;gas

nsolid n1;solid;n2;solid;.;nm;solid

yi;gas ni;gasPni1ni;gas

; yi;solid ni;solidPni1ni;solid

(5)

The Gibbs free energy is calculated assuming two separate

phases in close contact, namely a mixture of ideal gases, and

a perfectly mixed incompressible solid.

G Xn

i1Gi;gas Xm

i1Gi;solid (6)

Gi;gas nig0i niRT lnyiP=Pref

;Gi;solid nig0i niRT ln yi (7)

The number of moles of all species, ni,gas and ni,solid is con-

strained such that the elemental balance of the total system is

satisfied. Reference values for enthalpy, entropy, and the

temperature dependent specific heat, cp, have been obtained

from the HSC 7.0 database [24].

A steady state model, coded in Matlab [26], has been

developed for the conceptual looping plant layout shown in

Fig. 2; the model features open system chemical equilibrium

analysis, heat andmass balance on the reactors, and heat and

mass balance on the heat exchangers. The model is used to

predict reactor yields and identify the amount of reactant

gases necessary to achieve satisfactory conversion. Analyses

have been carried out for the temperature range between 27 Cand 960 C, at an operating pressure of 1 bar [9].

Fig. 1 e Equilibrium reactor diagram.2.1. Experimental facility

A bench scale experimental facility featuring a 21 mm inner

diameter tubular fluidized bed reactor for the iron/iron oxide

hydrogen production looping process has been fabricated. A

pictorial view of the hydrogen production experimental

facility is shown in Fig. 3 and a corresponding flow diagram is

shown in Fig. 4. The facility includes a 21 mm inner diameter,

0.6 m long fused quartz tube. Fused quartz is a non-crystalline

form of silica with a melting point of 1665 C [27]. To preventpowder carry-over, a 20 mm pore size stainless steel frit is

inserted at the top of the tube as depicted in reactor diagram

(Fig. 5). The powder is placed on a distributor made of

a Cotronics ceramic blanket thermal insulation material that

can withstand a temperature up to 1650 C [28]. The tube endsare sealed with stainless steel fittings using silicon O-rings

that canwithstand temperatures up to 300 C. The quartz tubereactor extends through an MTI electric furnace. The furnace

has a continuous operational range of 100e1000 C and canoperate at 1100 C for a short time span (less than 2 h). Thefurnace has a heating rate of 20 C /min. It is equipped witha PID controller and features 30 programmable segments (/1 C accuracy) [29]. The length of the furnace heating zone is300 mmwith a constant temperature zone length of 80 mm. A

K-type thermocouple is placed near the center of the furnace.

A steam generator consisting of four 200 W cartridge heaters

Fig. 2 e Conceptual looping plant layout.temperature of 121 C up to 540 C with a maximum gas

volumetric flow rate of 0.227 m3/min (8 CFM). Two water

cooled condensers are used. One condenser is used to deter-

mine the steam mass flow rate based on the volume of

condensate collected in a separate steady state measurement

prior to the experiment. The other condenser is used to

remove excess water from the hydrogen/steam mixture flow

discharging the reactor. The condensed water is accumulated

in a water trap and the weight of the water accumulated is

used to determine the amount of unreacted steam. The

volume of the produced hydrogen is determined by visual

Fig. 3 e Pictorial view of hydrogen production experimental facility.

Fig. 4 e Flow diagram of hydrogen production experimental facility.

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 0 7445

powder are used. The powder is mixed with 99.5% pure silica

in a 2:1 silica to iron volume ratio to retard sintering. The silica

Table 1 eWD experiment operating conditions.

Experiment # Steam mass flowrate (g/min)

Bed Ta (C)

1 3.5 0.2 956 72 1.9 0.1 962 73 0.9 0.1 950 74 3.5 0.2 641 55 0.9 0.1 681 56 1.9 0.1 693 5

a The temperature error for k-type thermocouples is estimated to

be 0.75% of the measured temperature [34].

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 07446inspection of water displacement in an inverted graduated

cylinder at normal conditions (NTP, 20 C and 101 kPa). Theinverted graduated glass cylinders with 2000 ml capacity are

immersed in a water bath. Stainless steel sheathed E-type and

J-type thermocouples are used to monitor and record the gas

temperatures entering and exiting the tube reactor as well as

the temperatures of the fittings. A K-type thermocouple is

used to monitor the bed temperature. A National Instruments

data acquisition board, NI USAB-6225 [31], is used to collect

the thermocouple and flow meter voltage signals. A Labview

virtual instrument is used to observe, control and collect the

experimental data.

2.2. Description of experiments

Experiments are carried out to evaluate the water dissociation

(WD) step in the Iron/Iron oxide looping cycle. Reactor bed

temperatures of 660 and 960 C and steam mass flow rates of0.9, 1.9, and 3.5 g/min are considered. Table 1 lists the oper-

ating conditions for the six WD experiments. The total dura-

tion of the WD experiments ranges between 35 and 50 min.

High purity Ancor MH-100 porous iron powder with 99.56%

purity manufactured by Hoeganaes Corporation is used [32].

The iron is a porous powder with an average apparent density

of 2.5 g/cm3, a material density of 7.87 g/cm3 and a melting

point of 1536 C. Results of the iron powder sieve analysis areshown in Fig. 6. In each experiment approximately 25 g of iron

Fig. 5 e Schematic depiction of electrical furnace and

tubular reactor.(SIL-CO-SIL 63, U.S. Silica) sieve analysis is also illustrated in

Fig. 6 [33]. The mixed iron/silica bed is placed on the distrib-

utor in the quartz tube, which is then sealed with stainless

steel fittings. The quartz tube extends outside the electrical

furnace. The bottom portion of the quartz tube is insulated

with a ceramic blanket that is held in placewith stainless steel

bands to prevent steam condensation. A nitrogen flow is

passed through the reactor with a volumetric flow rate of 2

sLPM to heat the system to at least 150 C and to purge the airin the system, thus preventing oxidation of the iron powder.

In industrial practice, no nitrogen will be used, discharge gas

will be recirculates through the reactors. Using a three-way

valve, the steam is either directed to a condenser, which

empties into a graduated cylinder, or the steam is directed to

the reaction chamber. The mass flow rate of steam is

controlled via the heat input to the boiler. The exact steam

mass flow rate is determined by measuring the rate of

condensate when steam is directed to the condenser prior to

the actual experiment. The electrical furnace temperature is

set for the desired reaction temperature and held at the

temperature for the duration of the experiment. Once the

stainless steel fitting temperatures reach 150 C and the steamflow rate reaches steady state, the nitrogen is shut off, and the

steam is directed into the gas heater section, where it is

superheated and then directed to the reactor. Hydrogen and

excess steam leave the reactor and pass through a condenser

upon initiation of the oxidation reaction. The condensed

water is collected in a sealed cylinder (water trap). After theFig. 6 e Iron and silica powder size distributions by weight.

removal of all excess water, pure hydrogen is directed into an

inverted water-filled, graduated cylinder. The accumulated

amount of hydrogen is determined by visual observation of

the water displaced from the graduated cylinders.

2.3. Error analysis

An error analysis is used to assess the measurement uncer-

tainty. The steam mass flow rate is determined from the rate

ation for the steam mass flow rate error is listed in Table 2.

PH2 _mH2$t

43MH2MFe

mFe;initial

(9)

Table 2 e Uncertainty in steam mass flow ratemeasurements.

Steam mass flow rate (g/min) Uncertainty (g/min)

3.5 0.2

1.9 0.1

0.9 0.1

relative error.

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 0 7447T (C) _mH2O(kg/min)

t(min)

Uncertainty(ml)

H2volume(ml)

Relativeerror (%)

660 3.5 0e15 10 200 5.0

15etfinal 5 200 2.5The volumetric hydrogen yield is determined bymeasuring

the displaced water volume in an inverted graduated cylinder.

The uncertainty associated with the measurement involves

a visual inspection of the water meniscus. The water

meniscus reading is affected by the disruption of hydrogen

bubbles rising through the inverted cylinder. These disrup-

tions are more frequent at higher rates of reaction. The

measurement uncertainty and relative error (error in

meniscus reading) are estimated and listed in Table 3.

3. Results and discussion

The measured hydrogen yields using the fluidized bed of iron

particles at different steam flow rates are compared to the

Table 3 eHydrogen yieldmeasurement uncertainty andof the steam condensate accumulation. The measurements

are repeated, and the standard deviation (s) is used as

a statistical measure of the absolute error. The measurement

uncertainty is taken as s. For each operating condition, thesteam mass flow rate is measured twice. The standard devi-1.9 0e15 5 200 2.5

15etfinal 1 200 0.5

0.9 0e15 5 200 2.5

15etfinal 1 200 0.5

960 3.5 0e6 20 200 10.0

6e20 10 200 5.0

20etfinal 5 200 2.5

1.9 0e20 10 200 5.0

20etfinal 5 200 2.5

0.9 0e20 5 200 2.5

20etfinal 1 200 0.5theoretical open system incremental equilibrium yield at bed

temperatures of 660 and 960 C. Figs. 7 and 8 show thehydrogen yield as a function of the cumulative steam fed to

the reactor for the 660 and 960 C respective bed temperatures.The abscissa shows the ratio of the cumulative steam mass

flowing into the reactor to the stoichiometric steam mass

necessary for complete conversion of Fe to Fe3O4,

PH2O _mH2O$t

43MH2OMFe

mFe;initial

(8)

The ordinate shows the ratio of the cumulative hydrogen

mass discharging the reactor to the stoichiometric mass of

hydrogen that can be produced from complete conversion

from Fe to Fe3O4,

Fig. 7 e The Open system hydrogen production at 660 C

for flow rates of 0.9, 1.9, and 3.5 g/min.Fig. 8 e The Open system hydrogen production at 960 Cfor flow rates of 0.9, 1.9, and 3.5 g/min.

Fig. 9 e The open system solid molar composition for the

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 07448Fig. 7 shows that hydrogen yield increases with decreasing

flow rate (increasing residence time) and moves toward the

thermodynamic limit with increasing cumulative mass of

steam entering the reactor. The influence of slow reaction

kinetics at lower temperatures is clearly discernible. In Fig. 8,

the hydrogen production rate is observed to be relatively

insensitive to the steam flow rate because reaction kinetics are

enhanced at higher temperatures. The thermodynamic limit is

approached,particularly athigh cumulative steamthroughput.

The cumulative steam throughput is quite important since

there is an energy cost for water to steam conversion.

The sharp bends in the theoretical yield curves are asso-

ciated with completion of the oxidation of metallic iron and

the completion of oxidation of FeO respectively (see also Figs.

9 and 10). Most of the theoretical steam-to-hydrogen conver-

sion occurs with small cumulative amounts of steam. This

allows for high theoretical energy efficiencies, since little

hydrogen production step at 660 C.excess steam needs to be produced. However, even at high

temperatures, experiments do not match the steep initial rise

Fig. 10 e The open system solid molar composition for the

hydrogen production step at 960 C.in cumulative H2 production. The experimental curves also do

not exhibit the two sharp bends. This is due to non uniform

mixing of the solid phase.

Figs. 9 and 10 show the variation of the theoretical solid

phase composition as a function of the cumulative amount of

steam employed for the 660 and 960 C respective bedtemperatures. At higher temperature (960 C) relatively moreFe3O4 and Fe2O3 are formed. A small amount of elemental Fe

persist at the 960 C bed temperature.Figs. 11 and 12 show the total solid phase mass normalized

by the initial ironmass as a function of the cumulativemass of

steam into the reactor for the 660 and 960 C respective bedtemperatures. Both the theoretical limit of solid phase mass

and that inferred from experimental hydrogen production

data via a gas-phase mass balance are shown. In both figures,

the large symbols at the end of the experimental curves denote

the final mass determined via weighing at the end of the

Fig. 11 e Predicted total mass of the solid phase for

hydrogen production step at 660 C.experiment. The discrepancy is attributed to the breakdown of

Fe-particles swept away during the experiment as well as

incomplete extraction of the solid phase after the experiment.

Fig. 12 e Predicted total mass of the solid phase for

hydrogen production step at 960 C.

the reactor increases, the experimentally observed cumula-

tive steam-to-hydrogen conversion approaches the theoreti-

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 0 7449cally predicted values. The steep initial rise of the theoretical

yield shows the potential for efficient conversion of steam-to-

hydrogen. However, particularly at low temperatures and

during the initial reaction phase, experimental yields remain

significantly below the theoretical limit. Increasing the resi-

dence time partially alleviates these issues. At higher

temperatures reduction of the flow rate (i.e., increasing the

residence time) has only a marginal effect on conversion,

indicating very slow effective kinetics beyond a certain

Hydrogen yield. This is consistent with ongoing kinetic

modeling in which two distinct kinetic regimes, (i) a shrinking

sphere regime, and (ii) a diffusion-limited regime, has been

identified. Based on this study it is concluded that the diffu-

sion-limited regime proves a severe obstacle to efficient

reactor operation and should be avoided. A combination of

measures is suggested to overcome these obstacles: (i) mini-

mize the particle size as far as possible without unacceptable

mass losses to maximize the surface to volume ratio, (ii)

increase the gas-phase residence time, e.g., via recirculation,

and (iii) only partially reduce and oxidize the iron based

reactants to avoid the diffusion-limited regime.

Acknowledgments

Financial support for this study by the United States Depart-

ment of Energy under Award No. DE-FE0001321 is gratefully

acknowledged.

r e f e r e n c e s

[1] Charvin P, Abanades S, Flamant G, Lemort F. Two-step watersplitting thermochemical cycle based on iron oxide redoxpair for solar hydrogen production. Energy 2007;32:1124e33.

[2] Nakamura T. Hydrogen production from water utilizing solarheat at high temperatures. Solar Energy 1976;19:467e75.

[3] Steinfeld A, Sanders S, Plumbo R. Design aspects of solarthermochmical engineering e a case study: two-step watersplitting cycle using the Fe3O4/FeO redox system. SolarEnergy 1999;65:43e53.

[4] Figueiredo JL, Orfao JJM, Cunha AF. The use of iron oxide asan oxygen carrier in chemical-looping combustion of4. Conclusions

Hydrogen production via the iron/iron oxide looping cycle has

been studied theoretically and experimentally. An incre-

mental thermodynamic equilibrium open system model has

been developed. Themodel has been used to predict hydrogen

yields, and solid compositions for the oxidation step and off-

gas as well as solid composition for the reduction step.

Theoretical predictions have been experimentally validated

for the hydrogen production step at 660 and 960 C for steamflow rates between 0.9 and 3.5 g/min. As the steam flow rate to

the reactor decreases, i.e., as the residence time of steam inmethane with inherent separation of CO2. InternationalJournal of Hydrogen Energy 2001;35:9795e800.[5] Mattisson T, Lyngfelt A, Cho P. The use of iron oxide as anoxygen carrier in chemical-looping combustion of methanewith inherent separation of CO2. Fuel 2001;80:1953e62.

[6] Dougherty W, Kartha S, Rajan C, Lazarus M, Bailie A, RunkleB, and Fencl A. Greenhouse gas reduction benefits and costsof a large-scale transition to hydrogen in the USA. EnergyPolicy 2009;37; 1: 56e67.

[7] Ishida MI., Jin H. Chemical-looping combustion powergeneration plant system. US Patent 1995; 447,024.

[8] Svoboda K, Slowinski G, Rogut J, Baxter D. Thermodynamicpossibilities and constraints for pure hydrogen productionby iron based chemical looping process at lowertemperatures. Energy Conversion and Management 2007;48:3063e73.

[9] Singh A, Al-Raqom F, Klausner J, Petrasch J. Hydrogenproduction via the iron/iron oxide looping cycle. Proceedingsof ASME 2011 5th International Conference on EnergySustainability & 9th Fuel Cell Science, Engineering andTechnology Conference ESFuelCell; 2011.

[10] Gregory DP, Pangborn JB. Hydrogen energytechnologydupdate 1976. International Journal of HydrogenEnergy; 1976:331e40.

[11] Simbeck DR, Chang E. Hydrogen supply: cost estimate forhydrogen pathways e scoping analysis. NREL/SR-540-32525.Washington, DC: USDOE; 2002.

[12] Shoko E, McLellan B, Dicks AL, Diniz da Costa. Hydrogenfrom coal: production and utilisation technologies.International Journal of Coal Geology 2006;65(3e4):213e22.

[13] Hacker V, Fankhauser R, Faleschini G, Fuchs H, Friedrich K,Muhr M, et al. Hydrogen production by steam iron process.Journal of Power Sources 2000;86:531.

[14] Hacker V. A novel process for stationary hydrogenproduction: the reformer sponge iron cycle (RESC). Journal ofPower Sources 2003;118:311e4.

[15] Hacker V, Faleschini G, Fuchs H, Fankhauser R, Simader G,Ghaemi M, et al. Usage of biomass gas for fuel cells by the SIRprocess. Journal of Power Sources 1998;71:226e30.

[16] Velazquez-Vargas LG, Gupta P, Li F, Fan LS. Hydrogenproduction via redox reaction of syngas with metal oxidecomposite particle. 22nd Annual International PittsburghCoal Conference 2005.

[17] Velazquez-Vargas LG, Gupta P, Li F, Fan LS. Hydrogenproduction from coal derived syngas using novel metal oxideparticles. 23rd Annual International Pittsburgh CoalConference 2006.

[18] Velazquez-Vargas LG, Li F, Gupta P, Fan LS. Reduction ofmetal oxide particles with syngas for hydrogen production.AIChE Annual Meeting, Conference Proceedings 2006.

[19] Gupta P, Velazquez-Vargas LG, Thomas T, Fan LS. Hydrogenproduction from combustion looping (solids-coal).Proceedings of the 29th International Technical Conferenceon Coal Utilization & Fuel Systems 2005; 1: 313e318.

[20] Gupta P, Velazquez-Vargas LG, Thomas T, Fan LS. Chemicallooping combustion of coal to produce hydrogen.Proceedings of the 30th International Technical Conferenceon Coal Utilization & Fuel Systems 2005;1:349e52.

[21] Gnanapragasam NV, Reddy BV, Rosen MA. Hydrogenproduction from coal using coal direct chemical looping andSyngas chemical looping combustion systems: assessmentof system operation and resource requirement. InternationalJournal of Hydrogen Energy 2009;34:2606e15.

[22] Roeb M, Gathmann N, Neises M, Sattler C, Pitz-Paal R.Thermodynamic analysis of two-step solar water splittingwith mixed iron oxide. International Journal of EnergyResearch 2009;33:893e902.

[23] Bale CW, Chartrand P, Degterov SA, Eriksson G, Hack K,Mahfoud RB, et al. FactSage thermochemical software and

databases 2002, vol. 26. Calphad; 2002. 189.

[24] Roine A, HSC Chemistry 7, Outotec, ISBN 978-952-92-6242-7.

[25] Sharma SD. Fuels hydrogen production gas cleaning:pressure swing absorption. Encyclopedia of ElectrochemicalPower Sources; 2009:335e49.

[26] MATLAB version 7.10.0.499 (R2010a).[27] Cs Vass, Smausz T, Hopp B. Wet etching of fused silica:

a multiplex study. Journal of Physics D: Applied Physics 2004;37(17).

[28] Cotronics Corp. 3000 F ceramic blanket. http://www.cotronics.com/vo/cotr/pdf/370.pdf. [accessed 02.10.11].

[29] MTI Corporation. Multi position tube furnace 1100 Cproduct manual. http://www.mtixtl.com/gsl1100x2tubefurnacewith2quartztubevacuumflangeand30segmentstemperaturecontroller-110v.aspx. [accessed02.10.11].

[30] Omega Engineering Inc. T type process air heater manual.http://www.omega.com/Heaters/pdf/AHP_SERIES.pdf.[accessed 02.10.11].

[31] National Instruments DAQ NI USB-6225 manual. http://www.ni.com/pdf/products/us/cat_usbmseries_625x.pdf. [accessed02.10.11].

[32] Hoeganaes corp. Ancor MH-100 specification sheet.http://www.hoeganaes.com/Product%20Datasheets/DataSheets%20Jan2001/ANCOR%20MH-100-1.pdf.[accessed 02.10.11].

[33] US silica. SIL-CO-SIL63. Silica material specification sheet.http://www.ussilica.com/uploads/files/product-data-sheets/industry/building-products/SILCOSIL63-Ottawa.pdf.[accessed 02.10.11].

[34] Omega Engineering. The temperature handbook, vol. 29.Stamford, CT: Omega Engineering Inc; 1995.

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 7 ( 2 0 1 2 ) 7 4 4 2e7 4 5 07450

Production of hydrogen via an Iron/Iron oxide looping cycle: Thermodynamic modeling and experimental validation1. Introduction2. Thermodynamic analysis2.1. Experimental facility2.2. Description of experiments2.3. Error analysis

3. Results and discussion4. ConclusionsAcknowledgmentsReferences