Thermal Reduction of Iron Oxide under Reduced Pressure and

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Jun 10, 2015 - reduction of Fe3+ to Fe2+ to evolve oxygen gas: ... O (g). 684 kJ at. 1450 C. 3 4. 2. (4). In order to avoid sintering the iron ...... Enthalpy (kJ kmol.

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Thermal Reduction of Iron Oxide under Reduced Pressure and Implications on Thermal Conversion Efficiency for Solar Thermochemical Fuel Production Abhishek K. Singh,† Nicholas J. AuYeung,*,‡ Kelvin Randhir,† Rishi Mishra,† Kyle M. Allen,†,∥ Jörg Petrasch,§ and James F. Klausner† †

Department of Mechanical and Aerospace Engineering, University of Florida, Gainesville, Florida 32611-6250, United States School of Chemical, Biological, and Environmental Engineering, Oregon State University, Corvallis, Oregon 97331, United States § FH Vorarlberg, Hochschulstraße 1, 6850 Dornbirn, Austria ‡

ABSTRACT: Successful implementation of solar thermochemical metal oxide water splitting cycles is dependent upon the ability to reach low partial pressures of oxygen during the thermal reduction step. Low partial pressures of oxygen are required to provide a thermodynamic driving potential for the thermal reduction reaction and avoidance of recombination. Achieving low partial pressures of oxygen (e.g., < 10−2 bar) may require a nontrivial energy input to the solar-to-fuel conversion process, negatively impacting the solar to fuel energy conversion efficiency. Three different strategies to reduce the partial pressure associated with oxygen generated during an iron oxide thermal reduction process were investigated using an open system thermodynamic analysis. These strategies include vacuum pumping, purging with an inert gas, and purging with steam. If the difficult to achieve solid-phase heat recuperation is neglected, open-system thermodynamic simulations show that vacuum pumping will have over twice the overall cycle energetic and exergetic efficiencies than those of inert purging; assuming oxygen separation is required every cycle in the case of inert or steam purging. To demonstrate the concept of vacuum pumping, thermal reduction of an iron−zirconia bed in a tubular reactor was performed at low pressures of approximately 10−4 bar at a temperature of 1450 °C. The maximum extent of reduction (14.2 ± 1.7 mol %) was reached after approximately 1 h of reduction at 1450 °C, while the predicted theoretical extent of reduction ranges from 16.5 mol % at 10−2 bar to 76.9 mol % at 10−4 bar. In the present analysis, reaction kinetics are not considered, and its application is limited to the thermodynamically driven processes.



INTRODUCTION AND THEORETICAL BASIS Solar thermochemical metal oxide cycles have garnered substantial interest recently due to the potential for high efficiency conversion of solar thermal energy into clean renewable fuels such as hydrogen or syngas. During such cycles, a metal oxide is first reduced to a lower oxidation state to evolve oxygen in a highly endothermic high temperature reaction: MOox ↔ MOred + O2

Δh > 0

carbothermal reduction of alumina to aluminum at temperatures achievable by solar thermal concentration.14,15 Darken et al.16 conducted an experimental study to determine the relation between temperature, gas composition (oxygen content), and composition of the iron oxide phases. They used the experimental data to calculate thermodynamic quantities such as activity and partial molal heat of iron and oxygen, heat of formation, and fusion of iron oxide of any composition possible at oxygen pressure up to 1 atm. They performed experiments for the temperature range up to 1640 °C and the equivalent pO2 in the gas phase from 10−10 to 1 atm. Using the experimental data and the calculated thermodynamic quantities, phase diagrams of the iron−oxygen system were constructed. Salmon et al.17 also carried out a high temperature thermodynamic study on the iron oxide system at various subatmospheric pressures. The theoretically developed activity expressions were derived and then validated with the experimental data for the iron oxide system. Using thermodynamic principles, expressions were derived relating standard heat of reaction and standard heat of entropy for the reaction to the activity expressions and oxygen dissociation pressure. These expressions can be used to calculate the oxygen dissociation

(1)

The second step then involves addition of a gaseous oxidizing agent in the form of a relatively abundant feedstock such as H2O or CO2, which is then split into H2 or CO, leaving a metal oxide in a higher oxidation state: MOred + H 2O ↔ MOox + H 2

Δh < 0

(2)

MOred + CO2 ↔ MOox + CO

Δh < 0

(3)

Heavily investigated candidates for such cycling include mixedmetal ferrites,1−6 cobalt ferrite via atomic layer deposition on alumina,7 and nonstoichiometric ceria.8−12 The extent of the reduction reaction can be enhanced by decreasing the total system pressure via vacuum pumping. Charvin et al. demonstrated the reduction of iron oxides under simulated sunlight at reduced pressure in a quartz vessel, achieving complete conversion of Fe2O3 to FeO at a temperature of 1700 °C and 0.1 bar over a duration of 2 min.13 Total pressure reduction via vacuum pumping has also enabled © 2015 American Chemical Society

Received: Revised: Accepted: Published: 6793

November 28, 2014 June 3, 2015 June 10, 2015 June 10, 2015 DOI: 10.1021/ie504402x Ind. Eng. Chem. Res. 2015, 54, 6793−6803

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Industrial & Engineering Chemistry Research

though largely inert above 1400 °C, would require an additional heat input for vaporization. Taking these energetic considerations into account is essential to identifying the most appropriate pressure reduction method for driving solar thermochemical reactions. In addition to the sensible heating requirement, there is an energy penalty associated with gas separation. The theoretical minimum amount of energy required to separate a binary mixture of gases can be described by

pressure of any oxygen content and temperature of the iron oxide system. In this investigation, the reaction of interest is the thermal reduction of Fe3+ to Fe2+ to evolve oxygen gas: Fe3O4 (s) ↔ 3FeO(s) + at T = 1450 °C

1 O2 (g) 2

Δh = +684 kJ (4)

In order to avoid sintering the iron oxide compounds, the iron is dissolved in a relatively homogeneous solid solution of 8 mol % yttria-stabilized zirconia (YSZ). The redox process is therefore a transition between Fe2+ and Fe3+ dissolved in YSZ. A greater extent of thermal reduction can be realized if the process is carried out under a reduced partial pressure of oxygen, as dictated by Le Chatelier’s principle obeying the equilibrium equation. Reaching a greater extent of reduction results in greater oxidation kinetics and fuel production in the subsequent step. Three options for reducing the partial pressure of oxygen, pO2, include (1) active pumping of oxygen; (2) purging with an inert gas; and (3) purging with steam. Although reaching a greater extent of reduction would intuitively seem desirable, the energy costs of pumping oxygen must be evaluated in order to assess the impact on the solar-to-fuel energy conversion efficiency. In addition, each of these three options has an associated economic cost that must be considered for large scale adoption. Siegel et al.18 recently discussed the implications of pumping on the energy conversion efficiency for solar thermochemical processes through the introduction of energy required for the pump work, Epump =

RTpump ηelec

ln

p0 pO

2

1 Esep,min = − RT[(1 − y) ln(1 − y) + y ln y] y

(6)

where R is the universal gas constant and y is the mole fraction of the gas in the binary mixture.22 It should be noted that this is the minimum energy required, and should the process involve electrical or mechanical work, a corresponding efficiency must be taken into account, as the actual amount of energy required may be orders of magnitude greater. Furthermore, this equation does not reveal anything about the proper method of separation or if it is even feasible via commonly used temperature, pressure, chemical, or membrane-based techniques.23 In an effort to determine the thermal implications of reducing the partial pressure of oxygen of an iron oxide based redox material, we have conducted an energy and exergy analysis of various partial pressure reduction strategies. In addition, a set of iron oxide vacuum reduction experiments investigating the extent of reaction has been completed.



THERMODYNAMIC ANALYSIS An open system, quasi-steady state, thermodynamic equilibrium model developed by Singh et al.24 was used to assess the thermodynamic consequences of reduced pressure for the three cases including vacuum pumping, inert gas purge, and steam purge. Assuming constant temperature and pressure and ideal gas behavior, the species balance for a gaseous/solid substrate is expressed as

(5)

where Epump is the total energy required by the pump per mole of O2; R is the universal gas constant; Tpump is the operating temperature of the pump; ηelec is the conversion efficiency of primary electrical energy to pump work; p0 is atmospheric pressure; and pO2 is the partial pressure of oxygen in the system. Purging a reactor with an inert gas is a common method of reducing the partial pressure of oxygen in laboratory scale experimentation.10,18,19 For large scale production, this would likely be costly both from an energetic and from an economic perspective unless the inert gas is purified and recycled to remove oxygen. Even if the gas can be recycled, Lapp et al. showed that the increase in ceria nonstoichiometry due to decreasing pO2 (via increase in inert gas flow rate) resulted in a lower efficiency, primarily due to sensible heating of the inert gas.20 Inert gas is therefore only a feasible solution if the gas can be purified for reuse and heat can be effectively exchanged between the exiting and entering streams. Ermanoski et al. have given a theoretical justification for vacuum pumping rather than inert purging during reduction of ceria. One of the main advantages is the relatively inconsequential thermal mass of the oxygen stream in the vacuum case relative to that of the inert gas used in the purging case. The latter would require a gas−gas heat exchanger in order to be thermally efficient.21 To avoid purification, there are alternative purging mediums such as air or steam that could be used in place of inert gases to avoid separation. However, a system purged with air (pO2 ∼ 0.21) would have a similar sensible heating requirement to that of a system purged with inert, as well as introduce oxygen which would be counterproductive. A system purged with steam,

dyi ,gas dt

= n(̇ yi ,gas,in − yi ,gas,eq (ngas, n solid , T , P))

yi ,solid = yi ,solid,eq (ngas, n solid , T , p)

(7)

where ngas = (n1,gas , n2,gas , ..., nn ,gas) n solid = (n1,solid , n2,solid , ..., nm ,solid) yi ,gas =

ni ,gas n ∑i = 1 ni ,gas

, yi ,solid =

ni ,solid n ∑i = 1 ni ,solid

(8)

Species considered in the thermodynamic equilibrium model for the iron oxide thermal reduction process are O2, H2O, H2, N2, FeO, and Fe3O4. Formation of Fe and Fe2O3 is negligible at the considered operating conditions,25 thereby these species are not considered for the thermodynamic equilibrium analysis. Equilibrium compositions for the open system reduction process are calculated using Gibbs free energy minimization yi ,gas/solid,eq (ngas, n solid , T , p) = arg min G(ngas, n solid , T , P) n gas, nsolid

(9)

where arg min is used to find the constrained minimization of Gibbs free energy. The Gibbs free energy is calculated assuming 6794

DOI: 10.1021/ie504402x Ind. Eng. Chem. Res. 2015, 54, 6793−6803

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Industrial & Engineering Chemistry Research two separate phases in close contact, namely, a mixture of ideal gases and a perfectly mixed incompressible solid, n

G=

m

∑ Gi ,gas + ∑ Gi ,solid i=1

i=1

(10)

Gi ,gas = nigi̅ 0 + n iRT ln(yP /Pref ) i Gi ,solid = nigi̅ 0 + niRT ln yi

(11)

where g0i̅ is reference gibbs function of species i evaluated (kJ kmol−1), yi is the species i mole fraction, P is total system pressure (N m−2), and Pref is a reference pressure (N m−2). 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.26 Figure 1 shows the thermodynamic limit of magnetite reduction at a given operating O2 partial pressure for thermal reduction of

Figure 2. Solid molar composition for varying O2 partial pressure at 1450 °C operating temperature.

is performed. In this analysis, thermal energy involved in solid material heating and cooling is not considered, as it is equal for all three cases. Simplified energy flow diagrams are also developed for each case. Inert Gas. In Figure 3, the energy flow diagram for the case in which an inert gas is used to reduce the partial pressure of O2

Figure 1. O2 yield for varying O2 partial pressure at 1450 °C operating temperature.

Figure 3. Energy flow diagram for the low partial pressure system using inert purge gas.

1 mol of magnetite (Fe3O4). The solar thermochemical reactor is assumed to be isothermal at 1450 °C for the reduction process. ∏O2 shows the ratio of the cumulative oxygen yield discharging the reactor to the stoichiometric amount of oxygen that can be produced from complete conversion from Fe3O4 to FeO. nO ,yield ∏= 1 2 n O2 (12) 2 Fe3O4 ,initial

inside the thermochemical reactor during the reduction step is shown. The output gaseous stream of the thermochemical reactor contains a mixture of inert gas and O2. Inert gas is separated using a high temperature separation membrane working at 950 °C, which was chosen to be a reasonable estimate of operating temperature based on current membrane technology reported in the literature.27−30 The separated inert gas is again fed to the thermochemical reactor. The high temperature membrane separation process considered here is well studied but not yet commercially available. In the current study, the inert gas separation process is assumed to provide pure inert gas after separation, although in practice, some O2 remains in the separated inert gas. A further step of purifying the retentate would be needed which likely involves passing the inert gas through an oxygen sorbent, incurring further energy penalty and cost. Sensible heat is supplied to the separated inert gas to boost the temperature to that at which the reactor is operating. The energy required for the complete process is given by

A pressure range of 10−4 to 10−2 bar is used to compile an energy balance on the system for each case. Equal cumulative oxygen production at a given O2 partial pressure and temperature for all three cases (vacuum pumping, inert gas purge, and steam purge) is considered for this analysis. The amount of inert gas or steam needed to obtain a given partial pressure of O2 is calculated using the open system, thermodynamic equilibrium model. The limit of magnetite reduction predicted by the open system thermodynamic analysis ranges from 16.5 mol % at 10−2 bar to 76.9 mol % at 10−4 bar. Figure 2 shows the variation of the corresponding solid composition of the iron oxide inside the reactor. Formation of magnetite increases with increasing partial pressure of oxygen. It is emphasized tha, for this analysis and all analyses to follow, thermodynamic equilibrium is always assumed.

Ereq = nO2Esep,min + n N2Δh N2, TN ,in → 1450oC 2

− εnO2ΔhO2, TN +O ,out → 950oC 2



2

− εn N2Δh N2, TN +O ,out → 950oC − εnO2ΔhO2,950oC → 25oC 2

ENERGY BALANCE An energy balance for each of the three cases of pressure reduction (inert gas purge, steam purge, and vacuum pumping)

2

(13)

where ε is the heat exchanger effectiveness. Equation 6 is used to calculate Esep,min. 6795

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Industrial & Engineering Chemistry Research Steam as a Purge Gas. Steam at temperatures exceeding 1400 °C is not a strong oxidizing agent and has been suggested as a possible purge gas.31 Steam purging also offers a simple separation process via condensation of water. Figure 4 shows

Figure 6. Required energy for the inert gas purge and steam purge cases.

temperature. In the case of steam purge gas, heat of vaporization (40 kJ/mol) and the energy required to boost the steam temperature to the reactor operating temperature amounts to a large energy input for the process. The energy requirement for the steam purge case is approximately 1 order of magnitude higher than that of the inert gas purge case, which suggests that steam purging is unlikely to be a preferred method of reducing the partial pressure of oxygen during the thermochemical reduction step within solar reactors. On the basis of this study, steam as a purge gas case is not considered in further calculations. Vacuum Pumping. Figure 7 shows the process flow diagram for the case in which a vacuum pump is used to generate

Figure 4. Ratio of reactor output steam to the reactor input steam during thermal reduction at various partial pressures of oxygen. During the reduction process steam is assumed to be in contact with magnetite.

the variation of the ratio of the reactor output steam to the reactor input steam based on reactor operating pressure. Thermodynamic predictions show that only a small fraction of the steam reacts at 1450 °C. Figure 5 shows the process flow diagram for the steam purge case. The high thermal energy content of the output gaseous

Figure 7. Energy flow diagram for the vacuum pumping system. Figure 5. Energy flow diagram of the steam purge system.

subatmospheric pressures inside the thermochemical reactor. The required energy to carry out the complete process is given by

stream of the reactor is used to evaporate water as well as superheat steam. Additional thermal energy is required to boost the steam temperature to that of the reactor. The required energy for the complete process is given by

Ereq = nO2Epump − εnO2ΔhO2,1450oC → 100oC − εnO2ΔhO2,100oC → 25oC

where nO2 is calculated using the open system, thermodynamic equilibrium model and eq 5 is used to calculate Epump. Bulfin et al. suggested that a vacuum pump efficiency decreases with decrease in the working pressure of the vacuum pump.32 Various electric to pump work efficiencies, ηelec = 0.001, 0.01, 0.1, and 0.2,18 are assumed for the investigation. The vacuum pump is assumed to be operating at a gas temperature of 100 °C.18 The recovered heat from the exhaust stream of the reduction process can be utilized for heating up the working fluid (CO2 or H2O) for the oxidation process. Figure 8 shows the energy required to carry out Fe3O4 reduction using vacuum pumping compared with an inert gas purge. The energy required for inert gas separation using a high temperature separation membrane was calculated at separation temperatures of 950 °C, 1150 °C, and 1350 °C. Equation 13 is used to calculate the energy requirement for the inert gas purge case. Esep,min is calculated using eq 6. Heat exchanger effectiveness (ε) is taken to be 0.9 for all the cases. Energy required for the vacuum pumping case is highly dependent on the electric to pump work efficiency (ηelec)

Ereq = nO2Esep,min + nsteamΔhTSteam,in → 1450oC − εnsteamΔhsteam, TH O+O ,out → 25oC 2

2

− εnO2ΔhO2, TH O+O ,out → 25oC 2

2

(15)

(14)

Figure 6 shows the required energy for the inert gas purge and steam purge cases. Heat exchanger effectiveness (ε) is taken to be 0.9 for both cases. In the case of inert gas purge, a separation membrane is assumed to be operating at 950 °C which considerably reduces the total energy requirement in comparison to the lower temperature separation for the current process. Equation 13 is used to calculate the energy requirement for the inert gas purge case. Esep,min is calculated using eq 6. According to eq 6, a decrease in separation temperature decreases the energy required to separate the inert gas or steam, but the sensible heating energy requirement increases. The increase in sensible heating of gas is much higher than the work for the inert separation which led to reduction of total energy requirement of the complete process at higher separation 6796

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Industrial & Engineering Chemistry Research of a vacuum pump. For ηelec = 0.2,the energy required for the vacuum pumping case is typically an order of magnitude lower than that using an inert gas purge. This large difference is due to the high amount of sensible heating required for an inert gas. At lower pO2 the difference in required energy is much higher because of the large amount of input gas (inert gas or steam) needed to obtain the lower partial pressures of O2. Figure 8 also shows a comparison of the required energy for the

inert gas to the reactor temperature. At higher gas separation temperature and lower vacuum pump efficiencies (ηelec), the inert purge case requires equivalent or lower energy input as compared to the vacuum pump case. It should be noted that this analysis has not included additional energy inputs that might be required of an actual membrane separation process, such as those associated with the creation and/or continuous purification of a second purge gas on the sweep side of the membrane.



ENERGY AND EXERGY EFFICIENCY ANALYSIS On the basis of energy flow considerations, an energy and exergy efficiency analysis is performed to compare vacuum pumping with inert gas purging. An input of heat (Q) is assumed to have an exergy of Q(1 − T0/T), where T0 is taken to be 298 K. Work inputs are assumed to be reversible. The exergetic efficiency (ηexergy) of a system is the ratio of the exergy exiting a system to the exergy entering a system, ηexergy =

ηenergy =

inert gas purge case when the separation membrane working temperature is varied. As shown, the energy requirement of the system decreases with increasing temperature due to the reduced amount of sensible heat input needed to boost the

nO2ψO + n H2ψH 2

(

(Q reaction + Q N + Q solid,heat − Q solid,cool) 1 −

Here ψi is the specific molar exergy of species i (kJ mol−1), Qsolid,heat is the sensible heat for heating the material from oxidation temperature to reduction temperature, and Qsolid,cool is the amount of heat lost by the reduced material while cooling down to the oxidation temperature from the reduction temperature. It is assumed that 100% of Qsolid,cool can be recovered from the system. Qreaction is the heat required for the endothermic reduction (2nO2Δh and Δh is the heat required to carry out the endothermic reduction per mole of magnetite based on the stoichiometric eq 4), and QN2 is the heat required to raise the temperature of separated nitrogen to reduction temperature (QN2 = mN2CpN2(Treduction − TN2,in), and mN2 is the total mass of nitrogen which passes through the system in one cycle). CpN2 is the average specific heat capacity of nitrogen for a temperature range, and TN2,in is the temperature of inert gas at the heat exchanger exit as shown in Figure 3 (where TN2,in = ε(Treduction − Tsep) + Tsep and ε is the effectiveness of the heat exchanger). Heat recovery for steam production during the oxidation step has not been considered for this analysis and is assumed to be available from waste heat coming off the reactor. The material composition considered is 80 wt % YSZ/20 wt % Fe3O4. The solid phase heat requirements, Qsolid,heat and Qsolid,cool, are calculated using the following equations,

∫T

Treduction

oxidation

(16)

nFuelHHVFuel Energyin

(17)

Inert Gas. When inert purging is used during thermal reduction, the exergetic efficiency is given by

2

Q solid,heat =

Exergyin

The energy-to-fuel conversion efficiency (ηenergy) of a system is the ratio of the sum of all forms of useful energy that comes out of a system to the sum of all forms of energy that goes into the system,

Figure 8. Required energy for magnetite reduction using vacuum pumping compared with an inert gas purge at separation process temperatures of 950, 1150, and 1350 °C.

ηexergy =

Exergyout

2

T0 Treduction

Q solid,cool =

)+n

∫T

O2Esep,min

Toxidation

reduction

+ n H2OψH O

(18)

2

[(mFe3O4 − 2nO2M Fe3O4)C p

Fe3O4

+ 6nO2MFeOC p ] dT + 6nO2MFeOL FeO FeO

+ mZrO2C p

ZrO2

(20)

where m denotes mass, M denotes the molecular weight, and the subscripts denote the respective materials. LFeO is the latent heat of FeO when it has a phase change at 1377 °C. It is important to note that the heat capacity term pertaining to FeO was neglected in the calculation of Qsolid,heat. This is based on the assumption that the oxidation step results in almost complete conversion of the reactive material to Fe3O4 and only a negligible amount of FeO is left in the system. Temperature dependent specific heat capacities are were taken from the HSC 7 database. Latent heat of phase change of FeO is added to the heat lost calculation. Qsolid,cool is assumed to be completely recovered from the system. The solar to fuel energy conversion efficiency for the inert gas purge case is computed as ηenergy = n H2HHV H 2

(mFe3O4 C p

(T ) + mZrO2C p

Fe3O4

(T )) dT

Q reaction + Q N + Q solid,heat − Q solid,cool + nO2Esep,min

ZrO2

2

(19)

(21) 6797

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Industrial & Engineering Chemistry Research Here HHVH2 is the higher heating value of hydrogen, 286 kJ mol−1. In this case, energy conversion and exergetic efficiencies are close to each other so they are shown in the same figure. Figure 9 shows the variation of solar to fuel energy

As expected, the energy conversion and exergetic efficiencies increase with increasing gas phase heat exchanger effectiveness. For any given gas phase heat exchanger effectiveness, a maximum energy conversion and exergetic efficiency is reached at highest considered partial pressure of oxygen at which the system is operating during thermal reduction. The maximum energy conversion and exergetic efficiency (39%) is obtained when the heat exchanger effectiveness is 0.9, partial pressure of oxygen is 10−2, and it is assumed all solid phase heat can be recuperated. If heat recuperation of the solid-phase species is not possible, efficiencies suffer dramatically under all conditions. It is observed that the energy conversion and exergetic efficiencies do not significantly differ and thus are plotted using the same line. This is due to the high temperature operation, and as a result, the term nH2OψH2O is negligible compared to the other terms in the denominator of eq 18. Also the magnitude of nO2ψ O2 + nH2ψH2 is about the same as nH2HHVH2. As a result, the computed solar to fuel energy efficiency is nearly the same as the exergetic efficiency. Vacuum Pumping. For the case of vacuum pumping the exergetic and solar to fuel energy conversion efficiencies are respectively given as

Figure 9. Solar-to-fuel energy conversion efficiency and exergetic efficiency for varying system pressure and gas phase heat exchanger effectiveness using an inert gas purge.

and exergetic efficiencies for the case where there is a hypothetical 100% solid heat recovery and the case where there is no solid heat recovery. ηexergy =

nO2ψO + n H2ψH 2

(

(Q reaction + Q solid,heat − Q solid,cool) 1 −

Treaction

)+n

O2Epump

+ n H2OψH O 2

(22)

Partial pressures below 10−4 bar are not feasible since below this pressure an economically prohibitive pumping scheme would be required. As expected, lowering the system pressure yields higher solar-to-fuel energy and exergetic efficiencies for the system if solid heat recovery is not considered. However, if solid heat recovery is considered, both energy and exergetic efficiency show a slight upward trend with increasing pressure, which is primarily due to the increasing pump work required to reach lower pressures. With 100% solid heat recovery, the maximum achievable efficiency using vacuum pumping during thermal reduction approaches 45% for the ηelec = 0.2 case. At lower vacuum pump efficiencies, the difference between energy and exergy efficiency values for the no solid heat recovery case and 100% solid heat recovery case is negligible due to the predominance of the vacuum pump energy requirement in both the cases. As is the case with an inert gas purge, the vacuum pumping energy conversion and exergetic efficiencies do not significantly differ. When considering energy and exergetic efficiencies, this analysis suggests that at lower operating pressures and higher pump efficiencies vacuum pumping is superior to inert gas purging for thermal reduction of iron oxide. At higher partial pressures, the situation becomes more complex. If a hypothetical 100% solid heat recovery and high gas−gas heat exchanger effectiveness are both possible, inert gas purging shows superior efficiencies. However, if solid heat recovery is not possible, vacuum purging is more beneficial than inert purging in this higher pressure regime. As shown in Figures 1 and 2, working at lower partial pressure is beneficial for the thermal reduction reaction to proceed. This analysis assumes that the inert gas would have to be purified of oxygen in each

and ηenergy =

2

T0

n H2HHVH2 Q reaction + Q solid,heat − Q solid,cool + nO2Epump (23)

As in the previous case, energy conversion and exergetic efficiencies are nearly identical and are shown in the same figure. Energy and exergetic efficiencies for the case of 100% solid heat recuperation between redox steps and the case of no heat recovery for vacuum pumping are shown in Figure 10 at

Figure 10. Solar-to-fuel energy conversion/exegetic efficiency for varying system pressure using vacuum pumping.

various system pressures and for various heat to pump work efficiencies ηelec. 6798

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Figure 11. Experimental apparatus used for thermal reduction and water splitting studies. A series of valves were used to direct the process flow toward either the mass spectrometer during oxidation or the vacuum pump during reduction.

sample (∼99%) was not in contact with either the alumina tube or the insulation frit, it was assumed that any reactivity with these materials and the sample was negligible. Dark iron based stains on the inside of the tube were noticed after the experimental runs but did not penetrate the fully dense alumina material significantly deeper than the immediate surface. Shorter, smaller diameter alumina tubes were placed above and below the sample to keep the sample in the same vertical position in the heated zone throughout rapid pressure swings. The experimental apparatus is shown in Figure 11. A high temperature furnace capable of sustaining temperatures up to 1700 °C (Sentrotech) was used. Gas flows were precisely controlled by volumetric mass flow controllers (Alicat) and entered the reactor via the same inlet as steam. Steam was delivered via a syringe pump flowing into heated stainless steel tubing, where temperature was kept at approximately 180 °C using proportionally controlled resistive heating tape. The steam generator was tested up to 3 g/min successfully. During oxidation, helium and argon were sent through the reactor at 50 ± 0.5 sccm each. Liquid water was injected via a syringe pump (New Era NE-450IL) at 0.2 ± 0.002 mL/min. Excess steam was kept hot in the section immediately following the reactor and then condensed using chilled water in a vertically oriented coaxial heat exchanger, allowing for condensate collection in a trap that was emptied periodically. A Welch 1400 vacuum pump was used to pump oxygen out of the chamber. Pressure gauges (Instrutech CVM-201) were installed both upstream and downstream of the bed. During vacuum operation, the steam generator and reactor outlet were closed, while the pathway to the vacuum pump was open. All physical (temperature and pressure) data were acquired by interfacing instrumentation with National Instruments DAQ board and Labview software.

cycle and the material undergoing reduction is Fe3O4. It is important to note that each metal oxide will have different thermodynamics, and in certain cases it may be more beneficial to use an inert gas purge as opposed to vacuum pumping. This may in fact be the case if the inert gas can be used for multiple cycles without purification and/or cooling/reheating, as it is possible that an inert gas might accumulate oxygen for a number of cycles before scrubbing is required. Nonetheless, storage or transport of the sheer volume of gas required will make heat retention quite difficult, in which case further losses are expected. As a final caveat, the analysis used to compute the efficiencies shown in Figures 9 and 10 does not consider heat loss from the reactor. Thus, even lower efficiencies are expected for a practical solar thermochemical reactor system.



THERMAL REDUCTION EXPERIMENTS To validate the concept of vacuum reduction and corroborate the theoretical thermodynamic analysis, several experiments were conducted on both iron oxide and cobalt ferrite. A representative material of 20 wt % Fe3O4 in 8 mol % yttria stabilized zirconia (8-YSZ) was prepared as described previously by Allen et al.33 Briefly, the material was prepared via coprecipitation of Fe, Y, and ZrO from their respective nitrate precursors using ammonium hydroxide. The resulting slurry was then filtered, and the resulting solids were dried and calcined. The sintered metal oxide ceramic was then crushed to roughly 100 μm and mixed with graphite particles of two distinct nominal size ranges of −20 + 100 and +300 mesh (Alfa Aesar). The mass ratio of metal oxide ceramic to 20 + 100 and +300 graphite is 6.25:2:1, respectively. The particle mixture was then placed in a 25 mm O.D./19 mm I.D. alumina tube with aluminosilicate insulation (Zircar Zirconia Buster M-35) rated to 1700 °C on the top and bottom. Since the majority of the 6799

DOI: 10.1021/ie504402x Ind. Eng. Chem. Res. 2015, 54, 6793−6803

Article

Industrial & Engineering Chemistry Research Measurement of the product gas production rates was done using a residual gas analyzer−mass spectrometer (Hiden HPR-20). Two inert gases (argon and helium) were used to dilute the product gases to act as an internal calibration standard. Volumetric flow rate ratios between the gas of interest and helium were calibrated to partial pressure signal ratios as measured by the instrument. The helium to argon signal ratio was used as an additional check to ensure constant gas flow conditions throughout the experiment. In actual experiments, the measured partial pressure ratio of product gas and helium was used to find the ratio of flow rates between that of the two gases, and the actual flow rate of the product gas was found by virtue of knowing the helium flow rate. During injection of steam into a nonreactive system, the baseline for hydrogen increase was determined. When oxidation of a reactive system was performed for an extended duration, the hydrogen level eventually reached the same base level. This baseline level of hydrogen production (due to water splitting inside the mass spectrometer itself) was then subtracted from all measurements of hydrogen productivity in order to avoid overestimating the production rates. Although mass spectrometers can achieve fast sampling rates, on the order of several seconds, the identification of gases can be problematic due to similar fragmentation patterns and molecular weights. Furthermore, apparent changes in gas composition can also be attributed to total pressure changes rather than composition changes. To avoid such problems, a gas chromatograph equipped with a thermal conductivity detector (TCD) and molecular sieve column was used in initial trials to validate the mass spectrometer measurements.

Figure 13. Influence of thermal reduction dwell time at 1450 °C on subsequent H2 production.

temperature to 1450 °C and ramping down without dwelling resulted in approximately half (1.5 ± 0.2 cm3 gtotal−1) of the maximum H2 production (3.0 ± 0.4 cm3 gtotal−1) in the subsequent oxidation step. The maximum H2 yield reached after 60 and 90 min reveals that approximately 14.2 ± 1.7 mol % of the Fe3O4 was reduced, while the predicted theoretical extent of reduction ranges from 16.5 mol % at 10−2 bar to 76.9 mol % at 10−4 bar. Considering the pressures achieved experimentally (see Figure 14), the equilibrium values are certainly far short of



RESULTS AND DISCUSSION Thermal reduction of an 8.84 ± 0.01 g bed of 20 wt % magnetite in 8-YSZ was performed under reduced pressure for various durations. The extent of reduction was evaluated based on the yield of H2 in the subsequent oxidation step. A typical hydrogen production rate curve is shown in Figure 12. Figure 14. Exemplary pressure profiles during thermal reduction as measured upstream and downstream of the reactive bed.

those predicted. A likely reason for this discrepancy is inactivity of a substantial portion due to sintering. The energy to fuel conversion efficiency is not computed for these experiments since the substantial heat loss from the small scale reactor renders such a computation meaningless. An exemplary pressure profile is shown in Figure 14. As expected, pressure upstream of the bed was consistently greater than that downstream of the bed; the pressure increase in the upstream pressure profile, which coincides with O2 production during the dwell period at 1450 °C, is of particular interest. It is indicative of the fact that the pressure field within the reactor is not homogeneous, and thus local thermodynamic equilibrium is influenced by the local pressure field. Further experimentation using reduced pressure was performed on a 10 wt % CoFe2O4−8 mol % YSZ mixture, which has been the subject of many other investigations.3,33 Using the same apparatus as was used in the iron oxide experiments, thermal reduction using both inert purging and vacuum pumping was carried out on an 18.00 ± 0.01 g bed. Again, temperature was ramped at 10 °C/min and held at the chosen reduction temperature between 1350 to 1500 °C for 12 min.

Figure 12. Typical volumetric production rate of H2 versus time for oxidation (1200 °C) following thermal reduction (90 min at 1450 °C under vacuum). Volumetric production rate is normalized to the total mass of material (including YSZ).

Operation of the experimental facility involved ramping the temperature from the oxidation temperature (1200 °C) to the target reduction temperature of 1450 °C at 10 °C per minute under vacuum and dwelling for a specified amount of time. The heating and cooling ramps required a combined total of 50 min, which in most cases was longer than the dwell time at 1450 °C. As can be seen in Figure 13, the case of simply ramping the 6800

DOI: 10.1021/ie504402x Ind. Eng. Chem. Res. 2015, 54, 6793−6803

Article

Industrial & Engineering Chemistry Research

thermal efficiency. Similarly, bed depth negatively affects the ultimate pressure of the vacuum system. Any real process implementation must take these issues into account.

In this case, vacuum levels downstream and upstream of the bed reached nominal values of 5 × 10−4 bar and 2 × 10−3 bar, respectively, which it should be noted is greater than that achieved in the iron oxide experiments (see Figure 12). Reasons for this discrepancy may include the larger bed size (18.00 g vs 8.84 g in the iron oxide case). In the case of inert purging, the majority of the reduction time was spent at a pO2 of about 10−3 with the peak pO2 ranging from 5 to 7 × 10−3 bar. It is therefore interesting to note that the inert purging case performed better than the vacuum case under similar pO2 levels, with the exception of the 1500 °C case (see Figure 15). The



PRACTICAL CONSIDERATIONS There are many practical concerns that must be considered when performing thermal reduction under reduced pressure using vacuum pumping. In several instances, minor leaks caused atmospheric air to rush into the reactor volume. The result was an unsuccessful subsequent oxidation due to oxygen having already oxidized the small quantity of thermally reduced metal oxide material. It is important to note that inert gas was used to sweep the reactor volume between reduction and oxidation steps. Though this procedure was followed in order to reach a steady partial pressure profile in the mass spectrometer before starting oxidation, this inert sweep also acts to remove remaining oxygen from the system. In a practical system where the use of inert gas is minimized, perhaps steam (relatively inert at high temperatures) could be used to purge out any oxygen from the reactor or downstream lines. By virtue of the low density of the gas, vacuum pumping inhibits recombination of H2 and residual O2. However, it also poses a significant risk of air infiltration and rapid oxidation of the reduced metal oxide upon opening the system postreduction to atmospheric or elevated pressure. During vacuum pumping, a large pressure discrepancy was noticed between the upstream and downstream positions. This is likely caused by a decrease in hydraulic conductance through the tortuous porous bed. In a larger reactor with greater packing depth, the actual achievable vacuum pressure is likely to be greater (i.e., worse) than that reached in these laboratory experiments. Since surface area in a porous material is commonly linked to smaller and higher density pores, achieving a high hydraulic conductance through a high surface area material remains a challenge. Finally, it is inevitable that pump oils will be degraded rapidly with exposure to water, oxygen, and other reactive gases. To avoid frequent filter or oil changes of mechanical pumps, a dry pump would be a better solution to deal with the wet, oxidizing gas environment. Actual implementation of vacuum pumping in a solar fuel production plant will involve many design considerations. Since processes operating close to 1500 °C will likely utilize dish concentration, vacuum pumping will also have to be done in a relatively distributed fashion. Conductance losses will limit the effective pressure reached and the fuel production efficiency and ultimately impact the selling price. Therefore, the question of how many reactors per pump becomes a function of both pump price and conductance loss for a given piping scenario. Finally, although the overall energy efficiency of a typical medium range vacuum pump of 0.2 was used here as has been used in other references,18 the actual efficiency can vary with the pressure regime. Although decreasing the pressure increases the thermodynamic driving force for thermal reduction, there is certainly a point at which operating at lower pressures is counterproductive. Pumping below the 10−3 Torr level typically requires specialized pumps (e.g., turbomolecular or cryogenic) backed by a roughing pump. By increasing the number of pumps, the parasitic energy expenditures will increase, and when factoring in the increasing equipment costs, it becomes an unattractive option compared to pumping at more reasonable pressures.

Figure 15. Comparison of hydrogen yield during subsequent oxidation after thermal reduction under both inert purge and vacuum pumping for an 18.00 g bed of 10 wt % CoFe2O4−8 mol % YSZ.

total reduction time (ramp up + 12 min dwell + ramp down to 1200 °C) increased with the dwell temperature which could be a reason why the vacuum trial outperformed the inert purge case at 1500 °C. To further compare inert purging versus vacuum, an additional experiment with a smaller bed of 10 wt % CoFe2O4 in 8 mol % YSZ (4.8 ± 0.1 g) was carried out at 1400 °C with a longer dwell time (45 min). With the lower bed mass, and thus a shorter bed, a lower pressure was achieved during pumping. As seen in Figure 16, the subsequent H2 yield

Figure 16. Hydrogen production during subsequent oxidation following thermal reduction at 1400 °C under both inert purging and vacuum pumping for a 4.8 g bed of 10 wt % CoFe2O4−8 mol % YSZ.

was comparable between the two regimes. It is worth restating that the pO2 temporal profiles are constant in the case of vacuum but dynamic in the case of inert purging, which could affect the results. Nevertheless, the partial pressures should be the same order of magnitude in either case. Longer dwell times used in the vacuum case would negatively impact the overall



CONCLUSION The extent of high temperature thermal reduction of metal oxides can be greatly enhanced via reducing the partial pressure 6801

DOI: 10.1021/ie504402x Ind. Eng. Chem. Res. 2015, 54, 6793−6803

Industrial & Engineering Chemistry Research



of oxygen. The open system energy balance, including energy flow diagrams, for the iron oxide reduction process using a vacuum pump, an inert gas, and steam as a purge gas has been performed. Equal amounts of cumulative oxygen production in each case at an operating pressure and temperature was considered as a basis of comparison. The least energy intensive option for reducing the partial pressure of O2 during the reduction step is by vacuum pumping at lower operating pressures. The high sensible heating requirement for the inert gas purge and high steam generation energy requirement for the steam purge makes both of these options energetically expensive. Both an energy and exergy analysis indicates that vacuum pumping yields efficiencies greater than those for inert gas purging when pumping efficiencies are greater than 0.1. The difference is primarily due to the high sensible heating requirement for inert gas. The work required for separating inert gas from oxygen, at least on a theoretical basis, is relatively small. The inert gas purge disadvantages could be mitigated if the inert gas (or steam) could be used for multiple cycles without requiring additional heating. However, such operation has yet to be demonstrated. An efficient (i.e., heat to pump efficiency greater than 0.1) vacuum pumping system for reducing the partial pressure of O2 will likely yield the highest solar to fuel energy conversion efficiency. If oxygen-inert gas separation is carried out at a temperature of 1350 °C with a gas-phase heat recovery effectiveness of 0.9 and 100% solid-phase heat recovery between redox steps, inert purging shows better energy and exergetic efficiency at high partial pressures of oxygen than the vacuum case; however, these conditions would be very difficult to achieve in a real system at an affordable cost. On the basis of thermodynamic considerations, an experimental investigation of iron oxide thermal reduction using a vacuum pump for thermal reduction was carried out. The thermal reduction of Fe3+ dissolved in an 8 mol %−YSZ matrix was carried out under reduced pressure at temperatures up to 1450 °C. The maximum extent of reduction (14.2 ± 1.7 mol %) was reached after approximately 1 h of reduction at 1450 °C. The thermodynamic limit of the magnetite reduction at 1450 °C temperature was obtained using an open system analysis is 16.5 mol % at 10−2 bar and 76.9 mol % at 10−4 bar. Further experimentation using a 10 wt % CoFe2O4 in 8 mol % YSZ material showed comparable performance between inert purging and vacuum pumping. In future work a systems-level approach for each new reactive material is needed to determine whether or not vacuum pumping is preferred.



Article

ACKNOWLEDGMENTS

Funding for this research was provided by the Department of Energy Advanced Research Projects Agency-Energy (ARPA-E) Award DE-AR000184. The authors would like to acknowledge Kuwait Institute for Scientific Research for their generous donation of the high temperature furnace used in this investigation, as well as Paul Annandale at Ebara Vacuum for his helpful discussions



NOMENCLATURE cp Specific heat capacity (kJ kmol−1 K−1) Esep,min Theoretical minimum amount of energy required to separate a binary mixture of gases (kJ/kmol O2) Ereq Required energy to carry out the complete iron oxide reduction process (kJ/kmol Fe3O4) Epump Total energy required by the pump per mole of O2 (kJ kmol−1) G Gibbs free energy (kJ) g0i̅ Reference Gibbs function of species i evaluated (kJ kmol−1) h Enthalpy (kJ kmol−1) Kp Equilibrium constant mi Mass of a species i (kg) M Molar mass (kg kmol−1) ni Number of moles for a species i (kmol) P Total system pressure (N m−2) Pref Reference pressure (N m−2) pO 2 Partial pressure of oxygen PN2 Partial pressure of nitrogen P0 Atmospheric pressure (N m−2) Qsolid,heat Sensible heat for heating the material from oxidation temperature to reduction temperature (kJ) Qsolid,cool heat lost from reactive material while cooling from reduction temperature to oxidation temperature (kJ) Qreaction Heat required for the endothermic reduction reaction (kJ) QN2 Heat required to raise the temperature of separated nitrogen to reduction temperature (kJ) R Universal gas constant (kJ kmol−1 K−1) SLPM Standard liters per minute t time (s) Tout, exchanger Temperature of the inert at heat exchanger exit (K) Toxidation Temperature at which oxidation of the reduced material is carried out (K) Treduction Thermal reduction temperature (K) Tsep Separation temperature of inert gas and oxygen (K) yi,eq Mole fraction at equilibrium yi Species i mole fraction ε Heat exchanger effectiveness ψi Specific molar exergy of species i (kJ mol−1) ηexergy Exergetic efficiency ηenergy Energy to fuel conversion efficiency ηelec Electric to pump work efficiency

AUTHOR INFORMATION

Corresponding Author

*Nicholas J. AuYeung. E-mail: [email protected] Phone: (541) 737-3766. Present Address





Kyle M. Allen. Sandia National Laboratories, P.O. Box 969 MS 9161, Livermore, California 94551-0969, United States. Author Contributions

REFERENCES

(1) Inoue, M.; Hasegawa, N.; Uehara, R.; Gokon, N.; Kaneko, H.; Tamaura, Y. Solar hydrogen generation with H2O/ZnO/MnFe2O4 system. Solar Energy 2004, 76, 309−315. (2) Fresno, F.; Yoshida, T.; Gokon, N.; Fernandez-Saavedra, R.; Kodama, T. Comparative study of the activity of nickel ferrites for solar hydrogen production by two-step thermochemical cycles. Int. J. Hydrogen Energy 2010, 35, 8503−8510.

The manuscript was written through contributions of all authors. All authors have given approval to the final version of the manuscript. Notes

The authors declare no competing financial interest. 6802

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(23) Kerry, F. G. Industrial gas handbook gas separation and purification; CRC Press: 2006. (24) Singh, A.; Al-Raqom, F.; Klausner, J.; Petrasch, J. Production of hydrogen via an Iron/Iron oxide looping cycle: Thermodynamic modeling and experimental validation. Int. J. Hydrogen Energy 2012, 37, 7442−7450. (25) Nowotny, J.; Rekas, M. Defect structures and thermodynamic properties of the wustite phase [Fe1‑yO]. J. Am. Ceram. Soc. 1989, 72, 1221−1228. (26) Roine, A. HSC Chemistry 7; Outotec, 2009. (27) Tong, J.; Yang, W.; Zhu, B.; Cai, R. Investigation of ideal zirconium-doped perovskite-type ceramic membrane materials for oxygen separation. J. Membr. Sci. 2002, 203 (1−2), 175−189. (28) Van Noyen, J.; Middelkoop, V.; Buysse, C.; Kovalevsky, A.; Snijkers, F.; Buekenhoudt, A.; Mullens, S.; Luyten, J.; Kretzschmar, J.; Lenaerts, S. Fabrication of perovskite capillary membranes for high temperature gas separation. Catal. Today 2012, 193, 172−178. (29) Wang, Z.; Kathiraser, Y.; Kawi, S. High performance oxygen permeable membranes with Nb-doped BaBi0.05Co0.95O3-delta perovskite oxides. J. Membr. Sci. 2013, 431, 180−186. (30) Xue, J.; Liao, Q.; Wei, Y.; Li, Z.; Wang, H. A CO2-tolerance oxygen permeable 60Ce(0.9)Gd(0.1) O(2‑delta)-40Ba(0.5)Sr(0.5)Co(0.8)Fe(0.2)O(3‑delta) dual phase membrane. J. Membr. Sci. 2013, 443, 124− 130. (31) Diver, R. B.; Siegel, N. P.; Moss, T. A.; Miller, J. E.; Evans, L.; Hogan, R. E.; Allendorf, M. D.; Stuecker, J. N.; James, D. L. Innovative Solar Thermochemical Water Splitting; Sandia National Laboratories: 2008. (32) Bulfin, B.; Call, F.; Lange, M.; Loebben, O.; Sattler, C.; Pitz-Pall, R.; Shvets, I. V. Thermodynamics of CeO2 thermochemical fuel production. Energy Fuels 2015, 29, 1001−1009. (33) Allen, K. M.; Coker, E. N.; AuYeung, N. J.; Klausner, J. F. Cobalt ferrite in YSZ for use as reactive material in solar thermochemical water and carbon dioxide splitting. J. Met. 2013, 65, 1670−1681.

(3) Scheffe, J. R.; McDaniel, A. H.; Allendorf, M. D.; Weimer, A. W. Kinetics and mechanism of solar-thermochemical H2 production by oxidation of a cobalt ferrite-zirconia composite. Energy Environ. Sci. 2013, 6, 963−973. (4) Allendorf, M. D.; Scheffe, J.; McDaniel, A. H.; Evans, G. H.; Weimer, A. W. Kinetics mechanism development for solar CO2 and H2O splitting using mixed-metal ferrites; Abstracts of Papers of the American Chemical Society; American Chemical Society: Washington, DC, 2009; p 238. (5) Neises, M.; Roeb, M.; Schmuecker, M.; Sattler, C.; Pitz-Paal, R. Kinetic investigations of the hydrogen production step of a thermochemical cycle using mixed iron oxides coated on ceramic substrates. Int. J. Energy Res. 2010, 34, 651−661. (6) Roeb, M.; Gathmann, N.; Neises, M.; Sattler, C.; Pitz-Paal, R. Thermodynamic analysis of two-step solar water splitting with mixed iron oxides. Int. J. Energy Res. 2009, 33, 893−902. (7) Arifin, D.; Aston, V. J.; Liang, X.; McDaniel, A. H.; Weimer, A. W. CoFe2O4 on a porous Al2O3 nanostructure for solar thermochemical CO2 splitting. Energy Environ. Sci. 2012, 5, 9438−9443. (8) Le Gal, A.; Abanades, S.; Flamant, G. CO2 and H2O Splitting for Thermochemical Production of Solar Fuels Using Nonstoichiometric Ceria and Ceria/Zirconia Solid Solutions. Energy Fuels 2011, 25, 4836−4845. (9) Chueh, W. C.; Falter, C.; Abbott, M.; Scipio, D.; Furler, P.; Haile, S. M.; Steinfeld, A. High-Flux Solar-Driven Thermochemical Dissociation of CO2 and H2O Using Nonstoichiometric Ceria. Science 2010, 330, 1797−1801. (10) Furler, P.; Scheffe, J.; Gorbar, M.; Moes, L.; Vogt, U.; Steinfeld, A. Solar Thermochemical CO2 Splitting Utilizing a Reticulated Porous Ceria Redox System. Energy Fuels 2012, 26. (11) Furler, P.; Scheffe, J. R.; Steinfeld, A. Syngas production by simultaneous splitting of H2O and CO2 via ceria redox reactions in a high-temperature solar reactor. Energy Environ. Sci. 2012, 5, 6098− 6103. (12) Scheffe, J. R.; Steinfeld, A. Thermodynamic Analysis of CeriumBased Oxides for Solar Thermochemical Fuel Production. Energy Fuels 2012, 26, 1928−1936. (13) Charvin, P.; Abanades, S.; Flamant, G.; Lemort, F. Two-step water splitting thermochemical cycle based on iron oxide redox pair for solar hydrogen production. Energy 2007, 32, 1124−1133. (14) Kruesi, M.; Galvez, M. E.; Halmann, M.; Steinfeld, A. Solar Aluminum Production by Vacuum Carbothermal Reduction of Alumina-Thermodynamic and Experimental Analyses. Metall. Mater. Trans. B 2011, 42, 254−260. (15) Halmann, M.; Frei, A.; Steinfeld, A. Carbothermal reduction of alumina: Thermochemical equilibrium calculations and experimental investigation. Energy 2007, 32, 2420−2427. (16) Darken, L. S.; Gurry, R. W. The System Iron-Oxygen. 11. Equilibrium and Thermodynamics of Liquid Oxide and Other Phases. J. Am. Chem. Soc. 1946, 68, 798. (17) Salmon, O. N. High temperature thermodynamics of the iron oxide system. J. Phys. Chem. 1961, 65, 550−556. (18) Siegel, N. P.; Miller, J. E.; Ermanoski, I.; Diver, R. B.; Stechel, E. B. Factors Affecting the Efficiency of Solar Driven Metal Oxide Thermochemical Cycles. Ind. Eng. Chem. Res. 2013, 52 (9), 3276− 3286. (19) Schunk, L. O.; Haeberling, P.; Wepf, S.; Wuillemin, D.; Meier, A.; Steinfeld, A. A receiver-reactor for the solar thermal dissociation of zinc oxide. J. Solar Energy Eng.-Trans. ASME 2008, 130 (2), 021009. (20) Lapp, J.; Davidson, J. H.; Lipinski, W. Efficiency of two-step solar thermochemical non-stoichiometric redox cycles with heat recovery. Energy 2012, 37, 591−600. (21) Ermanoski, I.; Siegel, N. P.; Stechel, E. B. A New Reactor Concept for Efficient Solar-Thermochemical Fuel Production. J. Sol. Energy Eng.-Trans. ASME 2013, 135, 031002. (22) Hänchen, M.; Stiel, A.; Jovanovic, Z. R.; Steinfeld, A. Thermally Driven Copper Oxide Redox Cycle for the Separation of Oxygen from Gases. Ind. Eng. Chem. Res. 2012, 51, 7013−7021. 6803

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