Thermodynamic modeling and experimental validation

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Production of hydrogen via an Iron/Iron oxide looping cycle: Thermodynamic modeling and experimental validation A. Singh, F. Al-Raqom, J. Klausner, J. Petrasch* Department of Mechanical and Aerospace Engineering, University of Florida, Gainesville, FL 32611, USA

article info

abstract

Article history:

An incremental thermodynamic equilibrium model has been developed for the chemical

Received 8 November 2011

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

Received in revised form

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

18 January 2012

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

Accepted 19 January 2012

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

Available online 23 February 2012

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

Keywords:

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

Looping cycle

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

Hydrogen

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

C operating temperature, the measured hydrogen yield

Iron oxide

dicted conversion. At a 960

Thermochemical

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

Syngas

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

Coal

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 variations in steam flow rate. This observation suggests that the reaction kinetics are significantly slower at lower temperature, and the model assumption is not satisfied. Copyright ª 2012, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rights reserved.

1.

Introduction

The solar thermochemical production of fuels such as hydrogen or Syngas using metal/metal oxide looping processes [1e3] is considered an interesting route to carbonneutral fuels. The concept shows promise in helping to satisfy a growing global energy demand, reducing oil price volatility, and mitigating anthropogenic climate change. Solar driven looping processes use water and CO2 as the sole feedstocks and concentrated solar radiation as the sole energy source. Looping processes using natural gas [4e7] or coal-

derived Syngas [8,9] as the reducing agent constitute an important stepping-stone toward carbon-neutral hydrogen production. This study uses iron/iron oxide redox pairs as the 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 and 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.: þ1 352 392 9129. E-mail address: [email protected] (J. Petrasch). 0360-3199/$ e see front matter Copyright ª 2012, Hydrogen Energy Publications, LLC. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2012.01.074


Nomenclature cp G g0i h _ m M mFe P

1

1

Specific heat capacity, kJ kmol K Gibbs free energy, kJ Reference Gibbs function of species i evaluated, kJ kmol1 Enthalpy, kJ kmol1 Mass flow rate, kg s1 Molar mass, kg kmol1 Mass of iron, kg Total system pressure, N m2

the 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 twostep 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 experimentally 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-ironoxide 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 is lowered below 800 C, and (iii) the oxygen partial pressure during 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 experimental studies. However, they could not experimentally verify the increased hydrogen yield at lower water splitting temperatures of approximately 800 C. They concluded that kinetics play an important role in the oxidation 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 equilibrium for the steam oxidation of pure iron to magnetite (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.

ni mFe,init Pref PID R sLPM T t yi,eq yi

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Number of moles for a species i, kmol Initial mass of iron, kg Reference pressure, N m2 Proportional-integral-derivative Universal gas constant, kJ kmol1 K1 Standard liters per minute Temperature, K time, s Mole fraction at equilibrium Species i mole fraction

In the current study, an incremental thermodynamic equilibrium model is employed to predict the maximum attainable reaction yields. The model approximates a wellmixed 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 step [2]: Fe þ 4=3 H2 O/1=3Fe3 O4 þ 4=3H2 ; Dh ¼ 31:75 kJ=mol at 960 C;

(1)

followed by the reduction step: 1=3Fe3 O4 þ 2=3CO þ 2=3H2 /Fe þ 2=3CO2 þ 2=3H2 O; Dh ¼ þ1:25 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 process producing 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:

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Fig. 1 e Equilibrium reactor diagram.

dyi;gas ¼ n_ yi;gas;in yi;gas;eq ngas ; nsolid ; T; P dt yi;solid ¼ yi;solid;eq ngas ; 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;eq ngas ; nsolid ; T; p ¼ arg min G ngas ; nsolid ; T; P ngas ;nsolid

Fig. 2 e Conceptual looping plant layout. (4)

where ngas ¼ n1;gas ; n2;gas ; .; nn;gas nsolid ¼ n1;solid ; n2;solid ; .; nm;solid ni;gas ni;solid yi;gas ¼ Pn ; yi;solid ¼ Pn i¼1 ni;gas i¼1 ni;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. G¼

Xn i¼1

Gi;gas þ

Xm i¼1

Gi;solid

Gi;gas ¼ ni g0i þ ni RT ln yi P=Pref ; Gi;solid ¼ ni g0i þ ni RT ln yi

(6) (7)

The number of moles of all species, ni,gas and ni,solid is constrained 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 and mass 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 C and 960 C, at an operating pressure of 1 bar [9].

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 prevent

powder 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 ends are sealed with stainless steel fittings using silicon O-rings that can withstand temperatures up to 300 C. The quartz tube reactor extends through an MTI electric furnace. The furnace has a continuous operational range of 100e1000 C and can operate at 1100 C for a short time span (less than 2 h). The furnace has a heating rate of 20 C /min. It is equipped with a PID controller and features 30 programmable segments (þ/ 1 C accuracy) [29]. The length of the furnace heating zone is 300 mm with 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 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 by passing it through an Omega Engineering, 1.37 cm outer diameter (0.2500 NPT) 200 W in-line gas heater [30] that is mounted vertically and is capable of heating gas from an inlet 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 determine 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.

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Table 1 e WD experiment operating conditions. Experiment # 1 2 3 4 5 6

Steam mass flow rate (g/min) 3.5 1.9 0.9 3.5 0.9 1.9

0.2 0.1 0.1 0.2 0.1 0.1

Bed Ta ( C) 956 7 962 7 950 7 641 5 681 5 693 5

a The temperature error for k-type thermocouples is estimated to be 0.75% of the measured temperature [34].

Fig. 5 e Schematic depiction of electrical furnace and tubular reactor.

inspection of water displacement in an inverted graduated cylinder at normal conditions (NTP, 20 C and 101 kPa). The inverted 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.

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 (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 distributor 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 place with 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 air in 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 steam flow 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 the

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 of 0.9, 1.9, and 3.5 g/min are considered. Table 1 lists the operating conditions for the six WD experiments. The total duration 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 are shown in Fig. 6. In each experiment approximately 25 g of iron

Fig. 6 e Iron and silica powder size distributions by weight.


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Table 2 e Uncertainty in steam mass flow rate measurements. Steam mass flow rate (g/min)

Uncertainty (g/min)

3.5 1.9 0.9

0.2 0.1 0.1

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 uncertainty. The steam mass flow rate is determined from the rate of 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, the steam mass flow rate is measured twice. The standard deviation for the steam mass flow rate error is listed in Table 2. The volumetric hydrogen yield is determined by measuring 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 disruptions 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

Fig. 7 e The Open system hydrogen production at 660 C for flow rates of 0.9, 1.9, and 3.5 g/min.

theoretical open system incremental equilibrium yield at bed temperatures of 660 and 960 C. Figs. 7 and 8 show the hydrogen 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, PH2 O ¼

_ H2 O $t m 4 MH2 O mFe;initial 3 MFe

(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, PH2 ¼

_ H2 $t m 4 MH2 mFe;initial 3 MFe

(9)

Table 3 e Hydrogen yield measurement uncertainty and relative error. T ( C)

660

_ H2 O m (kg/min)

t (min)

3.5

0e15 15etfinal 0e15 15etfinal 0e15 15etfinal 0e6 6e20 20etfinal 0e20 20etfinal 0e20 20etfinal

1.9 0.9 960

3.5

1.9 0.9

Uncertainty H2 Relative (ml) volume error (%) (ml) 10 5 5 1 5 1 20 10 5 10 5 5 1

200 200 200 200 200 200 200 200 200 200 200 200 200

5.0 2.5 2.5 0.5 2.5 0.5 10.0 5.0 2.5 5.0 2.5 2.5 0.5

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

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Fig. 9 e The open system solid molar composition for the hydrogen production step at 660 C.

Fig. 11 e Predicted total mass of the solid phase for hydrogen production step at 660 C.

Fig. 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 at high cumulative steam throughput. 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 associated 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 conversion occurs with small cumulative amounts of steam. This allows for high theoretical energy efficiencies, since little excess steam needs to be produced. However, even at high temperatures, experiments do not match the steep initial rise

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 bed temperatures. At higher temperature (960 C) relatively more Fe3O4 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 iron mass as a function of the cumulative mass of steam into the reactor for the 660 and 960 C respective bed temperatures. 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 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. 10 e The open system solid molar composition for the hydrogen production step at 960 C.

Fig. 12 e Predicted total mass of the solid phase for hydrogen production step at 960 C.


4.

Conclusions

Hydrogen production via the iron/iron oxide looping cycle has been studied theoretically and experimentally. An incremental thermodynamic equilibrium open system model has been developed. The model has been used to predict hydrogen yields, and solid compositions for the oxidation step and offgas 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 steam flow 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 in the reactor increases, the experimentally observed cumulative steam-to-hydrogen conversion approaches the theoretically predicted values. The steep initial rise of the theoretical yield shows the potential for efficient conversion of steam-tohydrogen. However, particularly at low temperatures and during the initial reaction phase, experimental yields remain significantly below the theoretical limit. Increasing the residence 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 diffusion-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) minimize 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 Department of Energy under Award No. DE-FE0001321 is gratefully acknowledged.

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