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Demonstration of a solar reactor for carbon dioxide splitting via the isothermal ceria redox cycle and practical implications Brandon Jay Hathaway, Rohini Bala Chandran, Adam C. Gladen, Thomas R. Chase, and Jane H. Davidson Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.6b01265 • Publication Date (Web): 11 Jul 2016 Downloaded from http://pubs.acs.org on July 12, 2016
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Demonstration of a solar reactor for carbon dioxide splitting via the isothermal ceria redox cycle and practical implications Brandon J. Hathaway, Rohini Bala Chandran, Adam C. Gladen, Thomas R. Chase, Jane H. Davidson1 Department of Mechanical Engineering University of Minnesota Minneapolis, MN 55455 Abstract The performance of a 4.4 kW solar receiver/reactor to split carbon dioxide via the isothermal cerium dioxide thermochemical redox cycle is characterized during steady state operation in a high flux solar simulator. The solar reactor is the first to implement the isothermal redox cycle. Design innovations for continuous fuel production and gas-phase heat recuperation distinguish it from prior art. During steady-periodic operation at 1750 K, 360 mL min-1 of CO is produced continuously over 45 redox cycles and up to 95% of the sensible heat of the process gases is recovered. The solar-to-fuel efficiency is 1.64% without consideration of the energy costs of producing nitrogen used as a sweep gas during reduction. With inclusion of the solar energy required to produce N2 via cryogenic separation, the efficiency is 0.72%. Based on the thermodynamic limitations of the cycle and the limited opportunity for increasing reactor efficiency beyond 2%, we conclude the isothermal approach to split CO2 or water via a thermochemical metal oxide redox cycle is not attractive for future development. Future research should leverage the demonstrated advances in reactor design that permit continuous fuel production and recovery of the sensible heat of process gases for alternative cycles such as hybrid isothermal reforming/redox cycles or two-temperature metal redox cycles capable of solid-phase heat recovery.
Keywords: solar, thermochemical, fuel, carbon dioxide, cerium dioxide, ceria, redox, isothermal
1
Corresponding author contact 612-626-9850,
[email protected] 1 ACS Paragon Plus Environment
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Introduction Developing sustainable solar-derived hydrocarbon fuels is imperative to address the global energy demands of the future and to reduce the anthropogenic emissions of greenhouse gases. The demand for liquid hydrocarbon fuels continues to rise, particularly in the emerging economies of China, India, and the Middle East. Global consumption of petroleum and other liquid fuels was 87 MMbbl per day in 2010 and is projected to reach 119 MMbbl per day in 2040.1 Meeting a fraction of this demand with solar-derived fuels presents enormous benefits but poses formidable scientific, technologic, and economic challenges. Numerous solutions to capture and store solar energy as a fuel, including solar photolysis, electrolyzers driven by photovoltaic cells or with integrated photoelectrochemical devices,2–4 artificial photosynthesis,5,6 and solar thermochemical redox cycles to split water and carbon dioxide,7–9 are being pursued worldwide. These nascent technologies have not yet reached commercial viability and the question of which approaches will be most economical and practical at large-scale in the future remains a topic of continued research and development. In the present work, we demonstrate operation of the first complete solar reactor to implement the isothermal ceria-based solar thermochemical cycle. Recently, the isothermal metal-oxide redox cycle was proposed for ceria10,11 and hercynite12 as an alternative to twotemperature thermochemical redox cycles. The cycle exploits the difference in oxygen chemical potential at high temperature between an inert sweep gas (or sub-atmospheric environment) during reduction and steam or carbon dioxide during oxidation. For ceria, the redox cycles for splitting water and carbon dioxide are given by reactions (R1) and (R2a) and (R2b), respectively. CeO2−𝛿ox → CeO2−𝛿rd +
Δ𝛿 O 2 2
(R1)
CeO2−𝛿rd + Δ𝛿 H2 O → CeO2−𝛿ox + Δ𝛿 H2
(R2a)
CeO2−𝛿rd + Δ𝛿 CO2 → CeO2−𝛿ox + Δ𝛿 CO
(R2b)
Fuel is produced in proportion to the difference in the concentration of oxygen vacancies between the reduced and oxidized states of ceria, ∆𝛿 = 𝛿rd − 𝛿ox . The motivations for implementing reactions (R1) and (R2) at the same temperature, as opposed to the two-temperature thermochemical metal oxide redox cycle, are lower thermal stresses in reactor components, faster oxidation kinetics,13,14 and elimination of energy losses associated with cyclic heating and cooling of the metal oxide. Together these advantages are 2 ACS Paragon Plus Environment
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exploited to simplify the design and operation of a solar thermochemical reactor. On the other hand, based on chemical thermodynamics, the isothermal cycle is less favorable than the twotemperature cycle, as pointed out in both the proposing literature10,11 and subsequent analyses of the cycle.14–17 An isothermal metal oxide redox cycle has the same thermodynamic limitations as direct thermolysis of CO2 or H2O, but with the benefit of separating the product gases into two streams.15 The thermal efficiency of the solar-to-fuel conversion process is used by the research community as a surrogate to economics to project the commercial viability of thermochemical metal redox cycles because reactors to implement these cycles are still early in the research and development phase.18,19 Efficiency correlates inversely with the size and cost of the solar concentrating field. Various definitions of efficiency have been proposed, but a full accounting of the thermal energy and work input to cycle should be considered. With this full accounting, efficiency is the ratio of the higher heating value (HHV) of the produced fuel to the solar energy plus work required to produce it: 𝜂th =
𝑛̇ fuel HHVfuel . 𝑄̇solar + 𝑊̇ /𝜂s→e
(3)
The input energy includes the direct solar thermal input to the process, 𝑄̇solar , and the solar thermal energy required to provide the parasitic work, 𝑊̇ /𝜂s→e , to produce and pump process gases and to separate the fuel from the product stream. Prior to the present work, solar reactors for the ceria-based thermochemical redox cycle have carried out the two-temperature cycle. The highest reported efficiency is 1.72% for a 3.8 kW reactor operated in batch mode with reduction on-sun at a maximum operating temperature of 1847 K, and CO production after cooling to ~960 K with no solar input. 20 The assumption made in calculating efficiency is that during off-sun oxidation, the available solar energy is used for another reactor or a different process. The reactor did not have a mechanism for heat recovery. Recovery of the sensible heat of the ceria as it is cycled has the potential to raise the efficiency.16 The only two-temperature reactor to demonstrate solid-phase heat recovery suffered a number of mechanical issues and reached a maximum average efficiency of 0.66%.21 The practical difficulty of implementing solid-phase heat recovery was one of the original motivations for pursuing the design simplifications offered by a reactor which operates isothermally. 3 ACS Paragon Plus Environment
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Prior projections of the efficiency of the isothermal ceria redox cycle were based on thermodynamic analyses of hypothetical reactors with either the assumption of chemical equilibrium 11,15–17 or using measured reaction rates from bench-top systems.14 The projected efficiencies vary from 0.2% to 18% due to the wide range of assumptions for the flow rates of inert sweep gas (or use of vacuum17,22) and oxidizer, the rate of thermal loss to the ambient, the effectiveness of recovery of the sensible energy of the process gases, and the method of accounting for work input. However, the earliest and highest predictions of efficiency for an isothermal cycle projected by Bader et al. 11 (co-authored by the corresponding author of the present work) were unrealistically high due to the use of an unsound counter flow equilibrium model. Both Krenzke and Davidson16 and Brendelberger et al.23 point out that the counter flow model is flawed due to the assumption of chemical equilibrium at both the inlet and outlet of the bed. Other projections of efficiency above 2% neglected the work required for gas separation or pumping and/or also applied the counter flow model for oxidation, or reduction. The most accurate projections of efficiency for the reactor presented in the present study are based on a fixed bed equilibrium model developed by Venstrom et al.,24 the equivalent model formulations presented by Bulfin et al.17 and Davenport et al.,25 and a detailed computational model of transport and kinetics in the reactor by Bala Chandran and Davidson.26 For operating conditions very similar to those used in the present study, the projected efficiency is 0.9%. In this study we present the design and demonstrate the performance of a 4.4 kW prototype solar reactor to carry out the isothermal ceria redox cycle. The prototype reactor features design innovations for continuous fuel production and highly effective gas-phase heat recuperation that distinguish it from the two-temperature ceria-based prototype solar reactors demonstrated to-date.20,21 The reactor was operated in a high flux solar simulator with gas flow rates and cycling times identified in prior work to maximize efficiency.14,27 The data are interpreted with the aid of an energy balance on the reactor to determine if modifications to the reactor are warranted and to recommend future research direction for solar fuels via the isothermal thermochemical approach. Reactor Design and Characterization The nominal 4 kWth prototype reactor is illustrated in Figure 1. Figure 2 shows images of the reactor in the solar simulator viewed (a) from the aperture and (b) from the rear where the gases enter and exit the reactor. Functionally the reactor comprises two fully integrated sections: 4 ACS Paragon Plus Environment
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Figure 1. Isometric cross-section view of the prototype isothermal ceria redox reactor indicating key component locations. Coordinate origins for the z and θ directions as indicated. The inset provides a detail view of the portion of a reactive element within the reactor cavity.
(a.)
(b.) Figure 2. Photographs of the isothermal reactor. (a) Front view, showing the cavity aperture with a manually operated door used during experiments to reduce heat loss between experiments. (b) Rear view, showing gas connections to the heat exchangers and the high-flux solar simulator lamp array in the background. 5 ACS Paragon Plus Environment
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a solar receiver/reactor cavity and the heat recovery section. The geometry of the receiver/reactor, including the cavity, fixed bed reactive elements, and heat exchanger, was developed with the aid of multi-physics computational models of the radiative exchange, chemical kinetics, and transport processes in the reactor26,28 and evaluation of a prototype heat exchanger.29 Selection of the materials of construction considered the chemical compatibility, temperature limits, mechanical properties, and thermophysical properties of available ceramic materials. Concentrated sunlight enters the receiver/reactor cavity through a converging conical frustum and 48 mm diameter circular aperture. This inlet section matches the optical characteristics of the University of Minnesota high-flux solar simulator.30 Within the cavity, solar radiation is distributed to six assemblies of co-axial dense alumina tubes, referred to as reactive elements. The annulus of each tube assembly is filled with 5-mm (length and diameter) cylindrical porous ceria particles. Total mass of ceria is 3.2 kg. The ceria particles were formed of 6 µm diameter and 62 µm long fibers. The effective porosity is 78% and specific surface area prior to cycling is 0.114 m2 g-1. The overall bed void fraction is 45%. The morphology of the particles ensures rapid inter-particle diffusion of oxygen and intra-particle transport of heat and mass.14 The tube assemblies extend beyond the solar cavity to the heat recovery section. In this section, the inner tube and the annulus are filled with alumina reticulated porous ceramic (RPC) to enhance radiative and conductive heat transfer.29,31 The RPC in the annulus has a fluid accessible porosity of 85-90% and 10 PPI (mean pore diameter of 2.5 mm). The RPC in the inner tube is 85-90% porous and 5 PPI (mean pore diameter of 5.1 mm). Prior experiments demonstrated 90% gas-phase heat recovery for gas flow rates slightly higher than those of the present study.27,29,31 The reactor is insulated with layers of porous ceramic and glass fiber insulation. Additional details of the architecture and materials of construction of the reactor are reported in a prior publication.27 The redox cycle is implemented by alternating the flow of sweep gas and CO2 through each reactive element/heat exchanger assembly. During reduction, nitrogen (< 10 ppm oxygen) is introduced at 48 L min-1 to the center tube of each reactive element. The gas flows through the inner tube, and then reverses direction and flows over the ceria particles in the annulus. Heat is 6 ACS Paragon Plus Environment
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recovered from the product gases as they leave the reactive ceria bed. After 100 s, the gas flow is switched automatically to pure (99.999%) CO2 at the same flow rate of 48 L min-1. Fuel is produced continuously by operating half of the reactive elements in reduction and the other half in oxidation. The flow rates and duration of oxidation and reduction were selected in an optimization procedure based first on measured rate data in an IR imaging furnace,14 followed by experiments with two reactive elements in the prototype reactor.27 The high flux simulator was controlled to provide 4.4 kW with an average concentration of 2420 kW m-2 at the aperture. Incident power was measured before and after the test using a water-cooled black-body calorimeter (±9% of reading). Temperature was measured at z = 0.022, 0.125, 0.33 m along the outer surface of each reactive element by alumina-sheathed Type B (Pt30% Rh / Pt-6% Rh) thermocouple probes (± 9 K absolute accuracy, ± 0.5 K precision). The surface temperature of the cavity was measured at angular positions of = 30°, 330° and axial position z = 0.28 m. The temperature of the outer wall of the heat recovery section of one reactive element was monitored at z = 0.593, 0.796, 0.896, 1.015, 1.423, and 1.651 m. Gas temperatures were measured at the inlet and outlet of the reactive element/heat exchanger assemblies with Type-K (Ni-Cr / Ni-Al) thermocouples (± 3 K). The absolute pressure (± 0.9 kPa) at the inlet as well as the differential pressure (± 0.3 kPa) from inlet to outlet were monitored. Temperature and pressure were recorded at one second intervals. Gas composition was measured using Raman laser gas spectroscopy (± 0.025% volume concentration). The composition of the gas stream at the exit of each reactive element was measured at one second intervals for a complete cycle every 4.5 cycles. The combined gas streams were also monitored. Gas production rates were calculated using eq. (4) 𝑉̇𝑖 = (𝑋𝑖 ⋅
𝑉̇in 1−𝑋O2
)
(4)
The volumetric flow rate of species 𝑖, given by 𝑉̇𝑖 , is determined from the measured volume fraction of species 𝑖, 𝑋𝑖 , the volume fraction of oxygen, 𝑋O2 , and the known inlet flow of either carbon dioxide or nitrogen 𝑉̇in . Cycle-average production rates of O2 and CO were obtained by numerically integrating the release rates over each half-cycle and dividing by the total cycle duration.
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Data Analysis The reactor efficiency defined by eq. (3) is the primary performance metric. The numerator is the product of the time-averaged total fuel production rate and the HHV of CO (283 kJ mol-1). The denominator is the measured radiative power input plus the calculated solar thermal equivalent of the work input. The work for producing nitrogen is based on the measured flow rate of sweep gas and the commercially relevant value of 12 kJ (mol-N2)-1 to provide sweep gas with less than 10 ppm oxygen via a cryogenic separation.32 Nitrogen produced cryogenically would be delivered at pressure sufficient to drive flow through the reactor. The work requirement for the separation of fuel from the oxidizer is predicted using the measured average concentration of CO at the reactor outlet and assuming a commercial separation process with an efficiency of 10% with respect to the ideal separation work, requiring 0.58 kJ (mol-CO2)-1.33 The mechanical pumping power is calculated using the measured pressure drop across the reactor and assuming an 80% efficient pump, requiring 0.44 kJ (mol-CO2)-1. A solar-to-electric efficiency of 25%, representative of a solar Stirling engine,34 is applied to convert the work to a solar thermal equivalent. A second performance metric is the effectiveness of gas phase heat recovery, defined in eq. (5) for either sweep gas (sg) or oxidizer (ox) flows. 𝜀{sg,ox} =
ℎ̅{sg,ox} (𝑇bed,in ) − ℎ̅{sg,ox} (𝑇amb ) ℎ̅{sg,ox} (𝑇bed,out ) − ℎ̅{sg,ox} (𝑇amb )
(5)
The effectiveness, 𝜀, is the change in enthalpy, ℎ̅, of incoming gas when heated from ambient temperature to the temperature at the inlet of cavity receiver/reactor section of the reactive element divided by the maximum possible change in enthalpy of the gas leaving the ceria bed if cooled to ambient temperature. The effectiveness of the heat exchanger is evaluated with the aid of the previously published computational model of the reactor.27 The two-dimensional, axisymmetric numerical model applies the transient, axial profiles of temperatures measured along the outer tube wall of the reactive element and the heat exchanger as time-dependent boundary conditions to determine temperature distributions along the flow path of the gases in the reactive element and the heat exchanger. The predicted gas temperatures entering the annulus of the heat exchanger from the ceria bed and exiting the inner tube of the heat exchanger prior to the bed are used to establish Tbed,out and Tbed,in in eq. (5). The temperature of the gas at the inlet of
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the reactor (Tamb) is the measured temperature of N2 or CO2, for reduction or oxidation, respectively. Complete details of the model of the packed bed of ceria and the alumina RPC in the heat exchanger are provided in previous publications.27,31 To interpret the contributions of fuel production, sensible heating, and thermal losses to reactor efficiency, a cycle-averaged energy balance on the reactor is applied. 𝑄̇solar = 𝑄̇rad + 𝑄̇loss + 𝑄̇chem + 𝑄̇sg + 𝑄̇ox
(6)
The solar input power equals the sum of thermal power lost to reflection and thermal emission, 𝑄̇rad , power lost via conduction through the insulation and natural convection within the open cavity, 𝑄̇loss , power consumed by the chemical reaction, 𝑄̇chem , and the sensible heating for the sweep gas and oxidizer gas flows, 𝑄̇sg and 𝑄̇ox respectively. The radiative loss term is calculated using eq. (4) based on the measured concentration, 𝐶 = 2420, the spatially averaged measured reactive element wall temperature, 𝑇w = 1750 K, and 𝜀app = 0.99, calculated using a conventional analysis of radiative exchange between gray, diffuse surfaces. 𝑄̇rad = 𝑄̇solar
𝜎𝑇w4 𝜀 𝐶 𝐼 app
(7)
The chemistry term is evaluated based on the measured fuel production rate and the net enthalpy change of carbon dioxide dissociation at the spatially averaged temperature of the ceria bed (~279 kJ mol-CO-1). 𝑄̇chem = 𝑛̇ CO ⋅ Δℎ̅rxn (𝑇̅r )
(8)
The cycle-averaged sensible heating terms are evaluated as a function of the heat exchanger temperatures, gas flow rates, and duration of reduction and oxidation. 𝜏{rd,ox} 𝑄̇{sg,ox} = 𝑛̇ {sg,ox} ( ) [ℎ̅{sg,ox} (𝑇bed,out ) − ℎ̅{sg,ox} (𝑇bed,in )] 𝜏rd + 𝜏ox
(9)
The gas phase heat recovery effectiveness defined in eq. (5) provides the relation between the bed outlet temperature, bed inlet temperature, and the ambient temperature. The thermal loss due to conduction across the insulation and convection from the open cavity is the value required to close the energy balance.
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Results An important attribute of the reactor is mechanical integrity. The reactor was operated for more than one hundred hours prior to the reported experiment. There was no loss of pressure or cracking of the reactive elements even with cavity temperatures as high as 1800 K, indicating their ability to withstand thermal stress without fracture. Figure 3 summarizes the key operational data (temperatures and fuel production) over 45 cycles (150 minutes). During steady-state operation, temperatures within the reactor, including
Figure 3. Steady-state performance of the 4.4 KW reactor for CO2 splitting via the isothermal ceria-based redox cycle: (a) total fuel production, (b) temporal and spatial-averaged surface temperatures in the solar receiver/cavity, and (c) gas temperatures at the inlet and outlets of the gas-phase heat exchangers (HX). 10 ACS Paragon Plus Environment
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the cavity, reactive elements, and insulation, are stable from cycle to cycle within ±0.2 K with a maximum rate of change of 1 K hr-1. With a 4.4 kW solar input, the spatially and temporally averaged temperature of the surfaces of the reactive elements is 1750 K±15 K (reported as one standard deviation) (Figure 3b). Over the course of a cycle, there are variations of up to 1 K due to differences in the thermal capacity of N2 (33 J mol-1 K-1) and CO2 (55 J mol-1 K-1). The volume averaged temperature of the ceria particles is 1740 K, based on a detailed computational model of the reactor at the present operating conditions.27,26 A major advance in reactor technology is continuous versus batch fuel production during on-sun operation. Average fuel production over 45 cycles is 360 ± 11 mL min-1, equivalent to 70 W based on the high heating value of CO (Figure 3a). Due to slight differences in the mass of ceria and temperature among the reactive elements there is a temporal variation of 45 mL min-1. The temporal variation could be ameliorated (as indicated by the dashed line in Figure 3a) by modifying the cycling order of the reactive elements. The dashed line was calculated from single-element fuel production measurements. The total fuel flow rate for different element cycling sequences was calculated from these data and the dashed curve represents the production that would be possible if the cycling sequence of the reactive elements was altered to minimize the temporal variation in total fuel flow rate. The fuel is provided at an average concentration of 0.25%, necessitating the fuel separation work term included in the efficiency definition, while the used sweep gas contains on average 0.12% oxygen by volume. The cycle averaged molar specific fuel production is 0.075 µmol s-1 g-(CeO2)-1 in agreement with the predicted performance for isothermal redox cycling based on bench scale measurements14 and exceeding the 0.02 µmol s-1 g -1 initially demonstrated for isothermal redox cycling of hercynite at 1623 K at un-optimized cycling conditions. A second advance is the integration of gas-phase heat recovery within the reactor and the high degree of heat recovery achieved. During reduction, 95% of the sensible heat of the N2 sweep gas is recovered. During oxidation, 93% of the sensible heat of the oxidizing gas is recovered. Gases enter the heat exchangers at ~298 K and exit at ~410 K (Figure 3c) indicating a large amount of thermal energy is maintained within the reactor and cycled between incoming and outgoing gases during operation. The 4 to 8 K cyclic variation in outlet temperature and heat recovery effectiveness is due to differences in the thermal capacity of CO2 and N2. The recovered 11 ACS Paragon Plus Environment
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heat represents 12 kW that otherwise would have to be provided by solar input. Losses from the heat exchanger to the ambient were ~0.2 kW due to conduction while unrecovered heat was lost at a rate of ~0.5 kW from gases exiting the heat exchanger above ambient temperature. For CO2 splitting, the average solar-to-fuel efficiency without consideration of the energy cost of producing nitrogen is 1.64%. The efficiency is 0.72% accounting for all work input as defined for eq. (3). The difference in these values illustrates the importance of evaluating the overall process when considering and comparing approaches for solar fuel production. The measured efficiency, including work input is slightly less than the 0.9% value predicted by a detailed computational model of the prototype reactor and estimates based on the fixed bed quasi-equilibrium model.26 The measured fuel production rate is 25% less than the value predicted. This difference could be due to a number of factors, most likely differences in modeled and actual temperature distribution. The order of magnitude agreement supports the prior projections that oxygen diffusion within the ceria lattice and surface kinetics are so rapid that the ceria particles are in thermodynamic equilibrium with the surrounding gas atmosphere.24 In concurrence with the cited prior studies, we conclude that the sweep and oxidizer gases are used efficiently in the present reactor and overall fuel production rate is governed by the thermodynamic driving force for isothermal operation.
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3 2 1
0
Figure 4. Contributions to the calculated efficiency of the prototype reactor. The thermal input to the reactor is broken down into the components of the reactor energy balance. Similarly, the components of the solar-equivalent-work terms are delineated. 12 ACS Paragon Plus Environment
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A breakdown of how the solar and work input are used is shown in the bar chart of Figure 4. Of the 4.4 kW solar input, 69 W is required for the endothermic redox reaction. The remainder is lost due to radiant reflection and emission from the solar cavity (1050 W), sensible heating of the sweep (550 W) and oxidizer (730 W) gases, plus 1970 W lost via conduction across the insulation and convection from the windowless cavity. Opportunities for improved efficiency are limited. Despite high losses relative to the solar input, there is little to be gained in efficiency with a change in reactor design for isothermal operation because of the inherent high thermodynamic threshold. While relatively high conductive losses are inherent to laboratory-scale devices, the efficiency would rise to only 0.9% even if the external surfaces of the reactor are assumed adiabatic. The degree of radiative emission relative to solar input is representative of a similar design at any scale. Moreover, while maintaining the same total input power and cavity surface temperature, only modest improvements can be anticipated if the solar concentration were increased to 3000 kW m-2, representing a high-end value for commercial solar optics.11 Even with 95% heat recovery effectiveness, sensible heating represents nearly 30% of the solar input. Gains in heat recovery effectiveness are unlikely without increasing the size of the heat exchangers beyond what is cost effective. Production of nitrogen represents the largest requirement for work input at 5 kW. Work input to separate CO from the product gas stream (24 W) and pumping (18 W) is relatively small. Thus considering the significant requirements for heating plus production of nitrogen, vacuum operation during reduction represents an opportunity for modest gains in efficiency.15,17 Cyclic application of a vacuum during reduction in the present reactor will require a vacuum pump that maintains highly efficient operation over a wide range of pressure differentials. If reduction under 1 Pa vacuum were implemented with an assumed 10% solar-to-pumping efficiency in place of the sweep gas, the projected solar-to-fuel efficiency of the reactor would reach 1.6%. An alternate approach is to modify the operation of the present reactor with the goal of reducing the amount of sweep gas without penalizing fuel production. The analysis of Brendelberger et al.,23 which uses a physically realizable counter flow model, indicates that a counter flow arrangement, which may be approximated by alternating the direction of flows between half-cycles, might possibly lead to improved performance. However, initial tests at the bench scale (unpublished) did not provide convincing evidence of improved fuel production.
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The fact that no improvement was observed is attributed to the small physical size of the fixed bed in the bench-scale setup and the resultant spatial uniformity of nonstoichiometry in the bed. Discussion Siegel et al. suggest a solar thermochemical fuel process must achieve an annual average solar-to-fuel efficiency of 20% to compete with photovoltaics plus alkaline electrolysis.19 Depending on the efficiency of solar optics and fuel separation processes, this value translates to a reactor-level solar to fuel efficiency on the order of 35%. Achieving this target is not possible for the isothermal ceria cycle represented by reactions (R1) and (R2). Based on the thermodynamic limitations of that cycle and the fact that the present reactor is close to the best one can anticipate for isothermal operation, it is unlikely that the solar-to-fuel efficiency of the isothermal metal redox cycle will ever exceed 2%. The kinetics of ceria cycling are sufficiently fast so that fuel production is not kinetically limited within the bounds of temperatures and gas flow rates expected in practice.24,25 Without kinetic limitations, the thermodynamics of water or carbon dioxide splitting under isothermal cycling are independent of the choice of metal oxide.15,25 Despite the low efficiency demonstrated for isothermal ceria-based CO2 splitting, outcomes of the study are attractive for future development of reactors. We recommend that future efforts to develop solar thermochemical reactor incorporate the design simplifications offered by the reactor presented here, namely no moving high temperature components, and integrated gas-phase heat recovery, but move beyond the cycle represented by reactions (R1/R2). Retention of the simplification of isothermal operation may be possible in a hybrid cycle that couples the isothermal metal redox CO2 or water splitting cycle with the partial oxidation of methane to co-produce synthesis gas and hydrogen.35 A thermodynamic analysis of the hybrid cycle suggests solar-to-fuel efficiency could reach 40%, which is competitive with solar electrolysis for H2 production.35 There are significant differences in the enthalpies of the chemical reactions of the proposed isothermal methane hybrid cycle and the CO2 splitting redox cycle as implemented in the present study. Thus modification of the reactor and reassessment of operating conditions and heat transfer are needed prior to implementation of the hybrid cycle. Two-temperature metal-oxide cycles (which require a swing in both oxygen partial pressure and temperature between oxidation and reduction) also offer favorable thermodynamics, and if solid-
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and gas-phase heat recovery is implemented successfully in reactors, it is theoretically possible to achieve the suggested target efficiency. Conclusion The solar thermochemical reactor demonstrated in this study converts concentrated solar energy into fuel by splitting carbon dioxide (or water) via the cerium dioxide redox cycle. The reactor is unique compared to prior art in that the redox cycle is conducted isothermally. Advances in the state of the art of solar thermochemical reactors are continuous, versus batch, fuel production and integration of a gas-phase heat recovery system capable of recovering as high as 95% of the sensible heat of process gases at 1750 K. The measured efficiency is 0.72%, according to the definition of eq. (3) which includes the solar thermal equivalent of work input. Based on both measured reactor performance and realistic projections of efficiency with a physically sound model of the chemical transport and kinetics within the reactor, we conclude that opportunities to reach higher efficiency for the isothermal ceria redox cycle are limited in scope and in expected gains. The isothermal cycle has inherent thermodynamic constraints that combined with the practical limitations of providing large amounts of sweep gas or implementing vacuum operation during reduction limit the efficiency to about 2%. Even so, future efforts to develop solar thermochemical reactors should seek to incorporate design simplifications offered by the current isothermal ceria redox reactor, namely no moving high temperature components and integrated gas-phase heat recovery, but focus on two-temperature cycles or hybrid isothermal cycles with lower thermodynamic thresholds. The two-temperature ceria redox cycle (which requires a swing in oxygen partial pressure as well as temperature) and the hybrid cycle with partial oxidation of methane combined with a metal oxide redox cycle (which can be operated isothermally at 1273 K) can theoretically reach or exceed the target efficiency of 35%. In the case of the two-temperature ceria redox cycle, implementation of gas and solid heat recovery is imperative. Acknowledgement The authors acknowledge the support of Jasper Adamek-Bowers, Jesse Fosheim, Peter Krenzke, Stephen Sedler, and Daniel Thomas during experiments in the solar simulator.
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Funding The financial support by the U.S. Department of Energy’s Advanced Research Projects Agency—Energy (DOE ARPA-E, award no. DE-AR0000182) to the University of Minnesota, the University of Minnesota Initiative for Renewable Energy and the Environment (IREE, grant no. RM-0001-12) is gratefully acknowledged. The information, data, or work presented herein was funded in part by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof. Nomenclature C HHV ℎ̅ I K M 𝑚 𝑛̇ PPI p 𝑄̇ RPC 𝑟 T t 𝑉̇ 𝑊̇ X z
Flux concentration ratio [-] Higher heating value of fuel [J mol-1] Molar specific enthalpy [J mol-1] Mass spectrometer signal [A], or standard insolation, 1000 [W m-2] Equilibrium constant Molar mass [g mol-1] Mass [kg] Molar flow rate [mol s-1] Pores per inch [inch-1] (English units are customary) Pressure [Pa] Heat transfer rate [W] Reticulated porous ceramic Rate of oxygen release or uptake [mol s-1] Temperature [K] Time [s] Volumetric flow rate [m3 s-1 @ 100 kPa, 298 K] Mechanical power [W] Volume or molar fraction [-] Axial coordinate [m] 16 ACS Paragon Plus Environment
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Greek 𝜓̇ 𝛿 Δ𝛿 𝜀 𝜂 𝜎 𝜏 𝜙 Subscripts
Rate of thermal dissociation of CO2 [mol s-1] Ceria nonstoichiometery [-] Cycle nonstoichiometry swing [-] Heat exchanger effectiveness [-], or emissivity [-] Efficiency [-] Stefan-Boltzmann constant [W m-2 K-4] Half-cycle time [s] Porosity or void fraction [-]
amb app bed chem fuel i in loss out ox pump r rad rd rxn s s→E sep sg solar th
Ambient conditions Apparent emissivity of cavity receiver Pertaining to the ceria bed Heat transfer due to chemical reactions of oxidation or reduction Pertaining to the useful fuel species (CO) General index Value at inlet Heat losses due to convection and conduction Value at outlet Pertaining to the oxidation half-cycle or oxidizer flow Pumping power Spatially averaged value across the reactive bed Radiative heat transfer Pertaining to the reduction half-cycle Enthalpy change due to chemical reaction Solid region Solar to electric conversion efficiency Energetic cost of gas separation Pertaining to sweep gas flow Incident solar power Thermal power or efficiency value
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GmbH & Co. KGaA: Weinheim, Germany, 2008; pp 9–109. (33) House, K. Z.; Baclig, A. C.; Ranjan, M.; van Nierop, E. A.; Wilcox, J.; Herzog, H. J. Proc. Natl. Acad. Sci. U. S. A. 2011, 108 (51), 20428–20433. (34) Mancini, T.; Heller, P.; Butler, B.; Osborn, B.; Schiel, W.; Goldberg, V.; Buck, R.; Diver, R.; Andraka, C.; Moreno, J. J. Sol. Energy Eng. 2003, 125 (2), 135–151. (35) Krenzke, P. T.; Davidson, J. H. Energy & Fuels 2014, 28 (6), 4088–4095.
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Energy & Fuels
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Figure 2. Photographs of the isothermal reactor. (a) Front view, showing the cavity aperture with a manually operated door used during experiments to reduce heat loss between experiments. (b) Rear view, showing gas connections to the heat exchangers and the high-flux solar simulator lamp array in the background. Figure 2 711x237mm (96 x 96 DPI)
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Qfuel,HHV [W]
(a.)
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20
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380 360
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Yield Rate [mL min -1]
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Power [kW]
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