High efficiency solar chemical-looping methane reforming with ceria in a fixed-bed reactor

Energy 169 (2019) 597e612

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High efficiency solar chemical-looping methane reforming with ceria in a fixed-bed reactor Jesse R. Fosheim, Brandon J. Hathaway, Jane H. Davidson* Department of Mechanical Engineering, University of Minnesota, 111 Church St. S.E., Minneapolis, MN, 55455, United States

a r t i c l e i n f o

a b s t r a c t

Article history: Received 18 May 2018 Received in revised form 6 December 2018 Accepted 7 December 2018 Available online 8 December 2018

High efficiency solar chemical-looping methane reforming is demonstrated in a prototype reactor operated in a high-flux solar simulator. The reactor includes six tube assemblies, which each comprise a fixed-bed of ceria particles and a gas-phase heat recuperator. The cycle was accomplished by alternating the flow to one tube assembly between CH4 and CO2. In the initial series of experiments, temperature, CH4 concentration, reduction flow rate, and cycle duration were varied to minimize carbon accumulation and maximize efficiency. In the second set of tests, the reactor was operated at optimized conditions for ten cycles at 1228 and 1274 K. Higher temperature favors better performance. At 1274 K, CH4 conversion is 0.36, H2 selectivity is 0.90, CO selectivity is 0.82, CO2 conversion is 0.69, and the energetic upgrade factor is 1.10. Heat recovery effectiveness is over 95%. Solar-to-fuel efficiency is 7% and the thermal efficiency is 25%. Projected solar-to-fuel and thermal efficiencies are 31 and 67% for the full-scale reactor and 56 and 85% for a commercial reactor with lower thermal losses. The demonstrated efficiencies are the highest reported to-date for this process. The projected scaled-up efficiencies suggest solar chemicallooping methane reforming could be a competitive approach for production of solar fuels. © 2018 Elsevier Ltd. All rights reserved.

Keywords: Solar thermochemical Chemical-looping Reforming Redox cycle Metal oxide Ceria

1. Introduction Solar chemical-looping methane reforming (sCL-MR) is a promising approach for producing mixtures of H2 and CO, referred to as synthesis gas or syngas, using concentrated sunlight. Chemical-looping methane reforming proceeds through two reaction steps. In the endothermic reduction step (R1), CH4 is partially oxidized to H2 and CO by oxygen released from a metal oxide oxygen carrier (MeOx).

CH4 þ

1

MeOxdox /

Dd

1

MeOxdrd þ 2H2 þ CO

Dd

(R1)

In the subsequent exothermic oxidation step (R2), H2O and/or CO2 are split by the reduced oxygen carrier, regenerating the metal oxide and producing a separate stream of H2 and/or CO.

H2 O þ

1

Dd

MeOxdrd /

1

Dd

MeOxdox þ H2

(R2a)

* Corresponding author. E-mail addresses: [email protected] (J.R. Fosheim), [email protected] (B.J. Hathaway), [email protected] (J.H. Davidson). https://doi.org/10.1016/j.energy.2018.12.037 0360-5442/© 2018 Elsevier Ltd. All rights reserved.

CO2 þ

1

MeOxdrd /

Dd

1

MeOxdox þ CO

Dd

(R2b)

The terminology of reduction and oxidation is with respect to the oxygen carrier and is the convention consistent with the literature on solar metal oxide redox cycles [1e5]. The nonstoichiometry, d, is the concentration of oxygen vacancies in the oxygen carrier. The amount of oxygen transferred during each reaction step is given by the swing in nonstoichiometry between reduction and oxidation, Dd ¼ drd - dox. The net products of (R1) and (R2) are the same as conventional methane reforming, but with the benefit of being separated into two valuable fuel streams. The reduction step (R1) can provide syngas with a H2/CO ratio of 2, which is favorable for the synthesis of liquid hydrocarbon fuels [6,7]. The H2 and/or CO from the oxidation step (R2) can augment the H2/CO ratio of the syngas from (R1) or provide a separate fuel stream (e.g. high-purity H2 for electricity generation in fuel cells or the synthesis of chemical commodities such as ammonia). Driving the process with solar energy rather than combustion of fossil fuels can reduce CO2 emissions by 41% and upgrade the energy content of the feedstock by a factor of 1.28. Utilizing readily available biogenic sources of CH4 would make the process carbon-neutral. In a comprehensive review of sCL-MR, Krenzke et al. predict

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solar-to-fuel efficiency can reach 54% at 1273 K with a mean flux concentration of 1000 suns (1 sun ¼ 1 kW/m2) if chemical equilibrium is achieved [8]. If equilibrium products were obtained with oxygen released from the oxygen carrier providing a CH4/O2 ratio of 2, syngas with a H2/CO ratio of 2 would comprise more than 99% of the reduction products above 1223 K. Experimental studies of numerous oxygen carrier materials demonstrate that equilibrium products are rarely obtained [8]. Two challenges for reaching high efficiency are achieving high CH4 conversion and high H2 and CO selectivities during the reduction step (R1) and high oxidizer conversion in the oxidation step (R2). There is a tradeoff in the selection of oxygen carrier and operating conditions between attaining high CH4 conversion and attaining high oxidizer conversion and syngas selectivities [8,9]. Ceria is a promising oxygen carrier due to its stable cubic fluorite structure [10e12], rapid oxygen diffusion kinetics [13], resistance to carbon deposition [14e16], and favorable thermodynamics for realizing high syngas selectivities and oxidizer conversion [8]. Based on conversions and selectivities measured in benchtop experiments conducted in a bed of ceria particles heated in an electric furnace, Krenzke et al. project that solar-to-fuel efficiency of a solar reactor operating under representative conditions would be 27% [9]. The potential for efficient production of solar fuel via sCL-MR supports development of prototype solar reactors as a step toward assessing the scalability and commercial viability of the approach. However, prior to the present work, reported efficiencies are lower than possible due to high thermal losses from the prototype reactors, high sensible heating requirements, and unoptimized operating conditions. Warren et al. demonstrated a single cycle at 1393 K in one of fourteen ceria fixed-beds in a solar reactor [17]. The active tube held 1130 g of 355e1000 mm ceria particles with a bed porosity of 69%. Reduction was performed for 120 min with 10% CH4 in Ar at a total flow rate of 2.97 mmol s1 g1 CeO2 . Oxidation was carried out for 120 min with 100% CO2 at flow rates of 0.30e2.97 mmol s1 g1 CeO2 . The CH4 conversion was 0.52 and H2 and CO selectivities were 0.83 and 0.59, respectively. The reduction products had a H2/CO ratio of 2.8. The low selectivity towards CO compared to H2 was due to carbon accumulation when the ceria bed was reduced to nonstoichiometries greater than 0.25. The average CO2 conversion was 0.21 due to the long oxidation duration. The reported solar-to-fuel efficiency was 0.92%. Warren et al. project that the solar-to-fuel efficiency would reach 7.51% if the process were performed at the full capacity of the reactor. Welte et al. investigated the reduction of ceria with diluted CH4 over a range of operating conditions in a solar particle-transport reactor [18]. Ceria particles (40 mm mean diameter) were fed (0.1e0.6 g s1) downward through a vertically oriented alumina tube enclosed in a solar cavity-receiver. The reactor was operated with both co-current and counter-current flow of CH4 (2.5e10% CH4 in Ar with total flow rates of 0.67e2.69 mmol s1) with residence times of <1 s. The nominal temperature of the reactive tube was 1423e1623 K. (The authors state that the temperatures of the reactive flows are likely lower than the reactive tube.) Thermal efficiency and upgrade factor were projected assuming that the ceria particles would be completely reoxidized by CO2 with no additional energy input. Projected performance is therefore predicated on the assumption that the heat release of the exothermic oxidation reaction is sufficient to provide sensible heating for the CO2 and make up for any other heat losses from the oxidation reactor. A peak thermal efficiency of 12% was projected for operation at 1576 K with co-current flows of 0.13 g s1 of ceria and 2.02 mmol s1 of 10% CH4 in Ar (CH4/CeO2 molar ratio of 0.27). The projected energetic upgrade factor was 1.14. The projected efficiencies are lower than prior thermodynamic predictions, in part,

due to high sensible heating requirements incurred by feeding the ceria particles and heavily diluted CH4 into the reactor at near ambient temperature. Most recently, Chuayboon et al. conducted sCL-MR with reticulated porous ceria foam (18.3705 g, 91.8% porous, <1 m2 g1 surface area) that was directly irradiated in a solar concentrator [19]. The nominal temperature of the ceria foam was varied from 1223 to 1323 K. Reduction was performed with 33e67% CH4 in Ar at total flow rates from 11.0 to 20.9 mmol s1 g1 CeO2 . Oxidation was carried

out with 55% H2O in Ar at 16.3 mmol s1 g1 CeO2 . The duration of each

reaction step was controlled to achieve complete reduction and oxidation of the ceria foam. The energetic upgrade factors were 0.97e1.10 and thermal efficiencies were 2.73e5.22%. Sensible heating requirements were high due to the need to cool the windowed aperture. Long reduction and oxidation durations yielded low average syngas product rates. Methane cracking during reduction resulted in the formation of up to 1.38 mmol g1 CeO2 of solid carbon and yielded syngas with H2/CO ratios up to 3.5. A portion of the carbon particles were swept from the reactor. The efficiency was low because of the combined effects of high sensible heating requirements, long reduction and oxidation durations, and the loss of carbon particles. In the present study, we report operation of a prototype fixedbed solar reactor with ceria at thermal steady state in a solar simulator for more than 20 chemical-looping methane reforming cycles. The experimental study was conducted in two phases. In the first phase, incremental changes were made in the operating conditions to eliminate carbon accumulation and maximize efficiency. In the second phase, the reactor was operated with a CH4 concentration of 75% at 1228 K and 1274 K for ten cycles at thermal steady state. An energy balance is performed on the reactor to project efficiency for scaled-up operation. The novelties of the present work are i) operation of the reduction step with higher CH4 concentrations than prior work, ii) implementation of gas-phase heat recuperation, and iii) demonstration of repeatable operation of sCL-MR over many cycles with negligible carbon accumulation and higher efficiency than has been reported. High CH4 concentrations are important to achieve high syngas production rates and reduce downstream separation requirements [20]. Implementation of gas-phase heat recuperation is important for achieving high efficiency. Demonstration of repeatable operation of sCL-MR is crucial for demonstrating the commercial viability of the process. 2. Experimental facility and methods 2.1. Facility The prototype reactor is illustrated in Fig. 1. The reactor was originally designed and characterized for continuous on-sun production of CO and H2 via a pressure swing, near-isothermal, thermochemical ceria redox cycle [21e23]. Here we provide an overview of the design of the reactor and describe modifications made for operation of sCL-MR. A more detailed description of the reactor geometry and construction is available in prior publications [21,22]. The 305 mm diameter, 347 mm long cylindrical receiver/reactor cavity has a 42 mm diameter windowless circular aperture within a converging conical frustum designed to match the radiative flux distribution of the University of Minnesota high-flux solar simulator [24,25]. Within the cavity, the radiative input from the solar simulator is distributed to six assemblies of co-axial alumina tubes (see Fig. 1(c) for details). The annulus of each tube assembly contains a fixed-bed of 5 mm (length and diameter) cylindrical fibrous ceria particles. The porosity and specific surface area of the particles

J.R. Fosheim et al. / Energy 169 (2019) 597e612

599

Fig. 1. Section views of the prototype solar reactor: (a) side view of the reactor cut from midplane AA; (b) front-facing view of reactor cut from plane BB; (c) detail view of cross section of the reactive element and integrated gas-phase heat recuperator.

are 78% and 0.114 m2 g1, respectively [22]. Each tube assembly extends beyond the receiver/reactor cavity to integrate with a gasphase heat recuperator [26,27]. In the heat recuperator, the center and annular regions of the tube assemblies contain alumina reticulated porous ceramic (RPC). The RPC is 85e90% porous with a pore density of 5 PPI (pores per inch) in the inner tube and 10 PPI in the annulus. Gases enter the reactor through the inner tube of the heat recuperator, reverse direction at the closed end of the outer tube located at the front of the cavity, then flow through the annulus over the fixed-bed of ceria particles and through the heat recuperator to the outlet. The reactor is insulated with layers of porous ceramic and glass fiber insulation within a 1.5 mm thick stainless steel housing. A schematic of the experimental facility is shown in Fig. 2. In the present study, sCL-MR was performed in the top-center tube assembly, hereafter referred to as the reactive element. The reactive element was designed with no connecting joints to permit hermetic operation under partial vacuum. The proprietary assembly method involves constructing the tubular assembly with the ceria particles in situ. The mass of ceria in the reactive element is 336 g and the interparticle void fraction of the bed is 55%. The cycle was accomplished by alternating the inlet gas flow between CH4 (reduction R1) and CO2 (oxidation R2) on a timed cycle. During operation with CH4 concentration less than 100%, N2 was added as the diluent. Gas flows to the reactive element were controlled with mass flow controllers (±1% of the full-scale output). The reactive element was held at 10 kPa gauge by a pressure controller (±0.15 kPa accuracy) immediately upstream of a dry diaphragm

vacuum pump capable of flows up to 40 mmol s1 on the outlet gas line. (The pressure control system was tuned to minimize oscillations, but transient fluctuations are observed at the start/end of each cycle as the controller responds to changing gas flows.) The five remaining tube assemblies, referred to as passive elements, were supplied with 6.7 mmol s1 of N2 at atmospheric pressure to prevent mechanical failure due to thermal stress. Surface temperatures of the tube assemblies and wall of the receiver cavity were measured using Type B (Pt-30% Rh/Pt-6% Rh) thermocouple probes (±6 K accuracy). Thermocouple probes were placed in contact with the outer surface of the tube assemblies at axial positions z ¼ 0.022, 0.127, 0.227, 0.330 m. Surface temperatures of the cavity were measured at axial position z ¼ 0.28 m and angular positions q ¼ 30 and 330 . Gas temperatures at the inlet and outlet of the tube assemblies and surface temperatures along the heat recuperator section of the reactive element were measured using Type K (NieCr/NieAl) thermocouple probes (±2.2 K accuracy). Absolute pressure was measured at the inlet of each tube assembly (±0.9 kPa accuracy). The pressure drop across each tube assembly was measured with a differential pressure sensor (±0.3 kPa accuracy). The outlet gas stream from the reactive element passed through a water-cooled heat exchanger to reduce the gas temperature to below ambient conditions (~295 K) and collect any condensate before gas analysis. The outlet mole fractions of CH4, H2, CO, CO2, and N2 were measured with a Raman laser gas analyzer (RLGA) (±0.025 mol% absolute accuracy). The RLGA was calibrated immediately prior to testing with analytical grade gas mixtures. A constant reference flow of 10.0 mmol s1 g1 CeO2 of N2

600

J.R. Fosheim et al. / Energy 169 (2019) 597e612

Fig. 2. Schematic of the experimental facility. Gas flow is denoted with solid lines. Measurements are represented by dashed lines. The top tube assembly is the reactive element. The bottom tube assembly is one of the five passive elements. Gas flow and measurement configuration for the five passive elements are identical.

was supplied downstream of the reactive element to quantify outlet flow rates (see data analysis section 2.3). Temperatures, pressures, and gas composition were recorded at one second intervals. The radiative input power to the reactor was controlled during testing using a calibration of the total radiative flux at the aperture versus the electrical current of each simulator lamp. The input power during the reported steady state tests was measured after the test with a water-cooled black-body calorimeter (±5%) at the lamp current levels used during the tests. 2.2. Procedure The reactor was characterized in two test phases. Flow charts of the procedure for each test phase are provided in Fig. 3. During Phase I, operational parameters were varied to observe the impacts of temperature, CH4 concentration, total reduction flow rate, and cycle timing on carbon accumulation and efficiency. The radiative input power at the aperture was controlled to maintain average reactive element temperatures of 1223 and 1273 K. For each target temperature, the reactor was initially operated for reduction with 100% CH4 at 10.1 mmol s1 g1 CeO2 for 480 s and oxidation with 100% CO2 at 7.1 mmol s1 g1 CeO2 for 300 s. These initial operating condi-

tions were selected to achieve the highest efficiency based on a thermodynamic process model of the reactor. The model used rate data obtained in a fixed-bed of ceria particles heated in an electric tube furnace [9]. At the initial set of conditions, a material balance calculation indicates that carbon was accumulating in the reactor, most likely in the heat recuperator where ceria particles are not

present. As a result, the operating conditions were adjusted in an iterative manner (See Fig. 3) to identify conditions under which carbon accumulation was eliminated. We considered reduction gas flow rates from 3.7 to 11.0 mmol s1 g1 CeO2 , CH4 concentrations from 50 to 100%, reduction durations from 240 to 480 s, and oxidation durations from 300 to 120 s. Based on the results of Phase I, a set of optimized operating conditions were selected for each target temperature for longer tests in Phase II. During these tests, referred to as Test 1 and 2, the reactor was operated at thermal steady state over ten cycles with radiative inputs of 1.5 and 1.6 kW and corresponding nominal reactive element temperatures of 1228 and 1274 K, respectively. Operating conditions for the steady state tests are listed in Table 1.

2.3. Data analysis The data were evaluated to illustrate temporal behavior and to quantify the key overall performance metrics, including conversions and selectivities, energetic upgrade factor, heat recovery effectiveness, solar-to-fuel efficiency, and thermal efficiency. The molar flow rates of CH4, CO2, and N2 delivered to the reactor were determined from the feedback signal of the corresponding mass flow controllers. The transient specific molar flow rates of CH4, CO2, CO, and H2 in the product stream were quantified using the measured species mole fractions and the total feed rate of N2, including the reference flow of N2 added to the product gas stream.

J.R. Fosheim et al. / Energy 169 (2019) 597e612

601

Fig. 3. Flow charts of experimental procedure for Phase I (left) and Phase II (right) of the reactor experiments.

Table 1 Operating conditions for the two steady state tests. Test Solar Input, ðQ_ solar Þ [kW]

1 2

Nominal Temperature, ðTr Þ [K]

1.5 ± 0.1 1.6 ± 0.1

1228 ± 12 1274 ± 12

Reduction

Oxidation

CH4 Feed Rate, ðn_ CH4 ;in;rd Þ [mmol s1 g1 CeO2 ]

N2 Feed Rate, ðn_ N2 ;in;rd Þ [mmol s1 g1 CeO2 ]

Reduction CO2 Feed Rate, ðn_ CO2 ;in;ox Þ Duration, ðtrd Þ [s] [mmol s1 g1 CeO2 ]

5.5 ± 0.1

1.7 ± 0.4

240

7.1 ± 0.1

Oxidation Duration ðtox Þ [s] 120

Uncertainty is the root sum square of the measurement uncertainty and temporal precision at a 95% confidence interval.

0




n_ i;out ¼ xi

n_ N2 ;in þ n_ N2 ;ref xN2 mCeO2

 (1)

0

n_ fCH4 ;H2 g;out;rd ¼

1

trd

The reported average inlet specific molar flow rates of CH4 and CO2 (mmol s1 g1 CeO2 ) are the instantaneous measured molar flow rates integrated over the duration of the reduction and oxidation steps divided by the duration of the corresponding reaction step.

0

n_ CH4 ;in;rd ¼

0

n_ CO2 ;in;ox ¼

1

trd

1

tox

t0 þ ðtrd


t0

 n_ CH4 ;in dt mCeO2

t0 þtð rd þtox
t0 þtrd

 n_ CO2 ;in dt mCeO2

(2)

(3)

The start time for each cycle, t0 , is when the inlet gases are switched from CO2 to CH4. The durations of reduction and oxidation are trd ¼ 240 s and tox ¼ 120 s, respectively. Similarly, the average outlet specific molar flow rates during reduction and oxidation are the integrated instantaneous values divided by the corresponding duration of the reaction step.

0

n_ fCO;CO2 g;out;rd ¼

0

n_ fCO;CO2 g;out;ox ¼

1

trd

1

tox

t0 þtð rd þtox

0

n_ fCH4 ;H2 g;out dt

(4)

t0 t0 þtres ð þtrd

0

n_ fCO;CO2 g;out dt

(5)

t0 þtres t0 þtres þ ðtrd þtox

0

n_ fCO;CO2 g;out dt

(6)

t0 þtres þtrd

In Eq. (4), the rates of CH4 and H2 during reduction are integrated over the entire cycle to eliminate inaccuracy due to dispersion. (All CH4 and H2 in the outlet gas stream are products of reduction.) The residence time, tres , in Eqs. (5) and (6) is the time lapse between when the inlet gas composition is changed and when the corresponding change in the outlet gas composition is detected by the gas analyzer (see section 3.1). During this short transition, the contributions of CO and CO2 in the product stream from reduction and oxidation cannot be differentiated. The sensitivity of the average flow rates of CO and CO2 during reduction and

602

J.R. Fosheim et al. / Energy 169 (2019) 597e612

oxidation to variations in the residence time is analyzed for residence times from 35 to 45 s. The variations in the calculated average specific species molar flow rates in the product stream are within measurement uncertainty. Results are presented for a residence time of 40 s. The average outlet specific molar rates of unmeasured species (H2O and C) are determined by material balances on hydrogen and carbon.



XCO2 ¼

n_ CO2 ;in;ox  n_ CO2 ;out;ox

(15)

n_ CO2 ;in;ox

The energetic upgrade factor (U) is the ratio of the (higher) heating value of the outlet species to the heating value of the CH4 feedstock. (HHVH2 ¼ 286 kJ mol1, HHVCO ¼ 283 kJ mol1, HHVCH4 ¼ 889 kJ mol1).



trd n_ H2 ;out;rd HHVH2 þ trd n_ CO;out;rd þ tox n_ CO;out;ox HHVCO þ trd n_ CH4 ;out;rd HHVCH4 U¼ trd n_ CH4 ;in;rd HHVCH4

  0 0 0 0 n_ H2 O;out;rd ¼ 2n_ CH4 ;in;rd  2n_ CH4 ;out;rd þ n_ H2 ;out;rd

(7)

  0 0 0 0 0 n_ C;net;rd ¼ n_ CH4 ;in;rd  n_ CH4 ;out;rd þ n_ CO;out;rd þ n_ CO2 ;out;rd

(8)

  0 0 0 0 n_ C;net;ox ¼ n_ CO2 ;in;ox  n_ CO;out;ox þ n_ CO2 ;out:ox

(9)

The left hand sides of Eqs. (8) and (9) are the average specific rates of carbon accumulation during reduction and oxidation, respectively. A positive value indicates deposition of carbon and a negative value indicates gasification of carbon. Similarly, the bedaveraged nonstoichiometry swing over reduction and oxidation is determined from a material balance on oxygen.

 0  0 0 hDdird ¼ MCeO2 trd n_ CO;out;rd þ 2n_ CO2 ;out;rd þ n_ H2 O;out;rd 

0

0

0

hDdiox ¼ MCeO2 tox n_ CO;out;ox þ 2n_ CO2 ;out;ox  2n_ CO2 ;in;ox

(10)

 (11)

(16)

A gas-phase heat recuperator is used to reduce the sensible heating load for gas flows through the reactive element. Krenzke and Davidson [28] showed that gas-phase heat recuperation becomes increasingly important for realizing high efficiency when reactants are not completely converted to syngas. The average heat recovery effectivenesses are estimated with respect to the minimum enthalpy fluid flow. The effectiveness during reduction is defined with respect to the inlet gas flow (Eq. (17)) and during oxidation with respect to the outlet gas flow (Eq. (18)).

erd ¼

1

trd

t0 þ ðtrd

t0

     1 0 _ n h T  h T r;z¼0:022 i i r;in @Pi¼CH4 ;N2 i;in;rd      Adt 0 _ i¼CH4 ;N2 ni;in;rd hi Tr;z¼0:330  hi Tr;in 0P

(17)

eox ¼

1

tox

t0 þtres þ ðtrd þtox t0 þtres þtrd

0P      1 0 _ i;out;ox hi Tr;z¼0:330  hi Tr;out n i¼CO;CO 2 @P      Adt 0 _ i¼CO;CO ni;out;ox hi Tr;z¼0:330  hi Tr;in

(18)

2

The bed-averaged nonstoichiometry swings are calculated using the cumulative amounts of oxygen uptake/release during each reaction step. Spatial nonstoichiometry gradients are expected to develop in the bed due to non-uniform oxygen transfer rates [9]. The performance metrics for both steady state tests are evaluated over ten cycles. Performance during reduction is quantified by the conversion of CH4 (XCH4 ) and selectivities towards H2 (SH2 ) and CO (SCO ). 0

XCH4 ¼

0

n_ CH4 ;in;rd  n_ CH4 ;out;rd 0

n_ CH4 ;in;rd

(12)

The gas temperatures at the inlet and outlet of the ceria bed are assumed equal to the temperatures measured at the corresponding axial positions along the outer wall of the reactive element (Tr;z¼0:022 and Tr;z¼0:330 for the inlet and outlet of the ceria bed, respectively). This assumption is supported by results of a twodimensional, axisymmetric numerical model of the reactor [23,29]. The solar-to-fuel efficiency (hsol ) correlates inversely with the size and cost of the solar concentrating field and is used commonly as a proxy for cost to project the commercial viability of nascent solar thermochemical technologies. In the present study, the solarto-fuel efficiency is defined as the difference in the heating value of the produced syngas (Q_ ) and the converted CH4 (Q_ ) divided syn

0

n_ H2 ;out;rd  SH2 ¼  0 0 2 n_ CH4 ;in;rd  n_ CH4 ;out;rd

(13)

SCO ¼

n_ CO;out;rd 0

0

n_ CH4 ;in;rd  n_ CH4 ;out;rd

vac

Q_

 Q_

CH4 hsol ¼ _ syn Q solar þ Q_ vac

0

(14)

Performance during oxidation is quantified by the conversion of CO2 (XCO2 ).

CH4

by the sum of the radiative input power to the reactor (Q_ solar ) and the equivalent solar heat to provide vacuum pumping work (Q_ ).

(19)

The work to produce the N2 used to dilute the CH4 during reduction is excluded because N2 provides no useful work to drive the process, unlike the inert gas used to maintain low oxygen partial pressures during reduction in sweep gas driven metal oxide cycles [21,22,30,31]. The cycle-averaged heating values of the

J.R. Fosheim et al. / Energy 169 (2019) 597e612

syngas produced during reduction ðQ_ syn;rd Þ, the syngas produced Þ, and the converted CH4 (Q_ ) are given during oxidation ðQ_ syn;ox

CH4

603

The cycle-averaged energy balance on the reactor for operation with a single reactive element is given by Eq. (26).

by Eqs. (20)e( 22) , respectively.

Q_ syn;rd ¼

Q_ syn;ox ¼ Q_ CH4 ¼




Q_ solar ¼ Q_ chem þ Q_ sens;r þ Q_ sens;p þ Q_ rad þ Q_ loss

 0  mCeO2 trd  0 n_ H2 ;out;rd HHVH2 þ n_ CO;out;rd HHVCO trd þ tox




  mCeO2 tox  0 n_ CO;out;ox HHVCO trd þ tox

(20)

The solar input to the reactor (Q_ solar ) is the measured radiative power at the aperture. The terms on the right hand side are the net power consumed by the chemical reactions (Q_ ), the sensible

(21)

heating of gases flowing through the single reactive element (Q_ ), the sensible heating of N2 flowing through the five passive

chem

sens;r


0  mCeO2 trd  0 n_ CH4 ;in;rd  n_ CH4 ;out;rd HHVCH4 trd þ tox

(22)

The cycle-averaged solar thermal power to provide vacuum pumping work during reduction ðQ_ vac;rd Þ and oxidation (Q_ vac;ox ) is the theoretical pump work divided by the product of the efficiency of the vacuum pump (hpump ), and the solar-to-electric conversion efficiency ðhs/e Þ.

Q_ vac;rd ¼

!


1

hpump hs/e 0

mCeO2 trd trd þ tox

X

@

elements (Q_ sens;p ), the heat lost via reflection and thermal emission through the aperture (Q_ ), and the heat lost by conduction rad

through the reactor insulation and natural convection in the open cavity (Q_ ). loss

The chemistry terms for reduction (Eq. (27)) and oxidation (Eq. (28)) are the average rates of enthalpy change between the reactants and products at the nominal temperature of the reactive

 element ( T r ).



0

n_ i;out;rd Ru Tamb

i¼CH4 ;N2 ;H2 ;CO;CO2

Q_ chem;rd ¼

1 
pamb A log pr




mCeO2 trd trd þ tox 2

4



X

i¼CH4 ;H2 ;H2 O;CO;CO2 ;C

3 0  0     n_ i;out;rd  n_ i;in;rd hi T r 5

(23) Q_ vac;ox ¼

!


1

hpump hs/e 0

X

@

mCeO2 tox trd þ tox

0

1 
pamb A log pr

Q_ chem;ox ¼ (24)


mCeO2 tox trd þ tox 2 3 0 X  0     4 n_ i;out;ox  n_ i;in;ox hi T r 5

(28)

i¼CO;CO2 ;C

The efficiency of the vacuum pump is assumed to be 40% based on the pressure-dependent correlations provided by Marxer et al. [30] and Bulfin et al. [32]. A solar-to-electric conversion efficiency of 25% is used to match the performance of typical solar dish Stirling systems based on solar energy intercepted at the aperture [33]. The reported solar-to-fuel efficiency does not account for the optical efficiency of the solar concentrating system. The thermal efficiency (hth ) is a useful metric for comparing solar reactors to conventional combustion driven systems. It is defined as the ratio of the heating value of the produced syngas (Q_ ) to the sum of the radiative input power (Q_ ), the solar syn

solar

heat to drive the vacuum pump (Q_ vac ), and the heating value of the converted CH4 (Q_ ). CH4

Q_

syn hth ¼ _ Q solar þ Q_ vac þ Q_ CH

(27)



n_ i;out;ox Ru Tamb

i¼CO;CO2

(26)

 The nominal temperature of the reactive element ( T r ) is the temporal and spatial average of the temperatures measured axially along the outer wall of the reactive element within the reactor cavity.



  T r ¼ Tr;z¼0:022 þ Tr;z¼0:127 þ Tr;z¼0:227 þ Tr;z¼0:330 4

(29)

The sensible heating terms for the reactive element during reduction (Eq. (30)) and oxidation (Eq. (31)) are the average power requirements to heat the gas flows at the inlet of the ceria bed to the nominal temperature of the reactive element plus the power required to heat the product gases from the nominal reactive element temperature to the temperature at the outlet of the ceria bed.

(25) 4

2.4. Energy balance and prediction of scale-up


mCeO2 trd Q_ sens;r;rd ¼ trd þ tox 2 0 X       n_ i;in;rd hi T r hi Tr;z¼0:022 4 i¼CH4 ;N2

An energy balance is performed on the reactor to determine the contributions of chemistry, sensible heating, and thermal losses to the solar-to-fuel and thermal efficiencies. Results of the energy balance for single element operation are extrapolated to forecast performance of the prototype reactor for operation with the full complement of six reactive elements and performance of a largerscale commercial solar reactor with reduced thermal losses.

þ

X

i¼CH4 ;N2 ;H2 ;H2 O;CO;CO2

3    hi T r 5

0    n_ i;out;rd hi Tr;z¼0:330

(30)

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6Q_


mCeO2 tox Q_ sens;r;ox ¼ trd þ tox 2

syn hth;fullscale ¼ _ Q solar;fullscale þ 6Q_ vac þ 6Q_ CH

0        4n_ CO2 ;in;ox hCO2 T r  hCO2 Tr;z¼0:022

X

þ

(36) 4

The solar input for full-scale operation of the reactor is given by Eqn. (37).

  Q_ solar;fullscale ¼ 6 Q_ chem þ Q_ sens;r þ Q_ rad;fullscale þ Q_ loss

3        5 n_ i;out;ox hi Tr;z¼0:330  hi T r 0

(37)

i¼CO;CO2

(31) To elucidate the impact of heat recuperation on efficiency, the maximum sensible heating load for the reactive element is estimated assuming gases enter the ceria bed at the average gas temperature at the inlet of the reactive element ðTr;in Þ (i.e. no heat recuperation). The sensible heating term for the five passive elements is the sum of the average power required to heat the N2 flow from the temperature at the inlet of the ceria bed to the temperature at the outlet of the ceria bed for each passive element.

The reaction enthalpy and sensible heating for each reactive element are assumed to be equal. The radiative emission and reflection losses are given by Eq. (38).

4 Q_ rad;fullscale ¼ εa sAap T cav þ ð1  εa ÞQ_ solar;fullscale

(38)

We assume the temperature distribution within the reactor with six reactive elements is the same as the temperature distribution with one reactive element. Consequently, the effective

 cavity temperature ( T cav ) and the conduction/convection losses ) are the same as with one reactive element. (Q_ loss

Q_ sens;p ¼

5 X

n_ N2 ;j



    hN2 Tj;z¼0:330  hN2 Tj;z¼0:022

(32)

j¼1

The radiative loss term is given by Eq. (33).

4 Q_ rad ¼ εa Aap s T cav þ ð1  εa ÞQ_ solar

(33)

The radiative losses include the average thermal emission (first term) and reflection (second term) from the solar cavity. Optical losses associated with the solar concentrating optics are not included. The apparent emissivity of the cavity (εa ) is 0.99 based on an analysis of radiative exchange between gray, diffuse surfaces.

 The cycle-averaged effective cavity temperature ( T cav ) is the temporal and spatial average of all temperature measurements along on the outer surface of the reactive element, the five passive elements, and the cavity wall. N

 1 X T cav ¼ Tj N

(34)

j¼1

Conduction/convection losses to the ambient are taken as the value required to close the energy balance. The losses predicted by the energy balance were validated using a 2-D axisymmetric conduction model of the reactor. The equivalent overall loss coefficient of ca. 1 W K1 is consistent with prior reactor tests [22]. In the present work, the prototype reactor is operated using one-sixth of the full reactor capacity. Improved performance is expected for scaled-up operation due to a higher fraction of the input power used to drive the chemical reactions and lower thermal losses relative to the solar input power. Performance of the prototype reactor with all six reactive elements is projected by extrapolating fuel production and energy requirements from the single-element tests. The inlet and outlet molar flow rates of N2, CH4, H2, H2O, CO, and CO2 for each reactive element are assumed equal to the measured performance of the single reactive element. The solar-to-fuel efficiency is projected using Eq. (35).

  6 Q_ syn  Q_ CH4

hsol;fullscale ¼ _ Q solar;fullscale þ 6Q_ vac The thermal efficiency is projected using Eq. (36).

(35)

Performance projections are also made for a commercial-scale solar reactor with lower thermal losses due to a smaller surfaceto-volume ratio. In this case, the approach follows the steps provided above, but the conduction/convection term is scaled as a fraction of the radiative solar input.

Q_ loss;commerial ¼ FQ_ solar;commercial

(39)

Efficiency projections are made for thermal loss fractions of F ¼ 0.05 and F ¼ 0.20, which have been suggested as reasonable estimates of thermal losses for commercial scale reactors [34e36]. 3. Results We first illustrate the behavior of the reactor during steady state cycling by showing temporal and spatial distributions of temperature and gas flows for a single representative cycle from Tests 1 and 2 (section 3.1). Second, we demonstrate the repeatability of cycling by presenting average data over ten cycles for both steady state tests (section 3.2). Third, an energy balance on the reactor is provided to guide the discussion of key areas for improvement (section 3.3) and to predict performance of the prototype reactor for full-scale operation with six reactive elements and for a largerscale commercial solar reactor (section 3.4). 3.1. Representative cycles To illustrate the temporal behavior of the reactor over a complete cycle, data are plotted for representative cycles from the two steady state tests (see Table 1). Transient specific species molar flow rates, inlet pressure, and reactive element temperatures from Test 2, which provides the best performance, are plotted in Fig. 4 and discussed in detail. Results for Test 1 follow the same general trends as Test 2, albeit with lower fuel production at decreased temperature, and are shown in Fig. 5. Reported uncertainty is the root sum square of measurement uncertainty and precision error based on a 95% confidence interval. For Test 2, the input power is 1.6 ± 0.1 kW at a mean flux concentration of 900 ± 50 kW m-2. The nominal temperature of the

 reactive element is T r ¼ 1274 ± 12 K. At the beginning of the cycle (t ¼ t0 ¼ 0), the inlet gas flow is switched from CO2 (used for the preceding oxidation step) to CH4 and N2. The inlet flow rate of 0 CH4 increases rapidly over 2e3 s and stabilizes at n_ CH ;in;rd 4

J.R. Fosheim et al. / Energy 169 (2019) 597e612

605

Fig. 4. Data for a representative sCL-MR cycle during Test 2 (solar input: 1.6 kW, reduction CH4 inlet flow rate: 5.5 mmol s1 g1, reduction N2 inlet flow rate: 1.7 mmol s1 g1, reduction duration: 240 s, oxidation CO2 inlet flow rate: 7.1 mmol s1 g1, oxidation duration: 120 s): (a) transient inlet and outlet specific molar flow rates; (b) gauge pressure at the inlet of the reactive element; (c) axial temperatures along the outer wall of the reactive element.

¼ 5.5 ± 0.1 mmol s1 g1 CeO2 . The inlet flow of N2 increases over the 0

first 25 s and stabilizes at n_ N2 ;in;rd ¼ 1.77 ± 0.03 mmol s

1

g1 CeO2 .

The

average inlet flow rates of CH4 and N2 over the entire reduction 0

period are n_ CH4 ;in;rd

¼ 5.5 ± 0.1 mmol s1 g1 CeO2

0

and n_ N2 ;in;rd

¼ 1.7 ± 0.4 mmol s1g1 CeO2 . After the gas valves are switched, the

inlet pressure fluctuates for 30e60 s due to the unsteady response of the downstream pressure controller and vacuum pump. The low amplitude cyclic variations in pressure after t ¼ 60 s are attributed to the cyclic behavior of the downstream vacuum pump and

pressure controller. The average gauge pressure at the inlet of the reactive element over the full reduction period is pr ¼ 9.9 ± 0.1 kPa. Fluctuations in the outlet flow rates follow the fluctuations in pressure. The transition from oxidation to reduction products at the gas analyzer begins approximately 30 s after the inlet gases are switched. For t  30 s, the outlet gas comprises the products of the preceding oxidation. From 30 s  t < 40 s, the outlet gas is a mixture of CH4, H2, H2O, CO, and CO2 due to dispersion between the oxidation and reduction product streams. For 40 s  t  240 s, the

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Fig. 5. Data for a representative sCL-MR cycle during Test 1 (solar input: 1.5 kW, reduction CH4 inlet flow rate: 5.5 mmol s1 g1, reduction N2 inlet flow rate: 1.7 mmol s1 g1, reduction duration: 240 s, oxidation CO2 inlet flow rate: 7.1 mmol s1 g1, oxidation duration: 120 s): (a) transient inlet and outlet specific species molar flow rates; (b) gauge pressure at the inlet of the reactive element; (c) axial temperatures along the outer wall of the reactive element.

outlet gas comprises exclusively reduction products. After the 0 0 initial rise, n_ CH4 ;out ¼ 3.3 ± 0.5 mmol s1 g1 n_ H2 ;out CeO , 2

0

1 1 _ ¼ 3.6 ± 0.7 mmol s1 g1 gCeO2 . CeO2 , and nCO;out ¼ 1.7 ± 0.3 mmol s

The average H2/CO ratio is 2.1 ± 0.2. The production rate of CO2 0 decreases as the ceria bed is reduced from n_ CO ;out;rd 2

¼ 0.8 ± 0.1 mmol s1

g1 CeO2

at

t ¼ 40 s

to

0

n_ CO2 ;out;rd

¼ 0.1 ± 0.1 mmol s1 g1 CeO2 at t ¼ 240 s. The increase in CO selectivity with increasing bed-averaged ceria nonstoichiometry is consistent with prior experimental findings [9].

The temperatures along the outer surface of the reactive element vary axially due to the non-uniform distribution of radiative flux along the length of the reactive element. The temporallyaveraged temperatures of the reactive element at the axial locations corresponding to the front and back of the ceria bed are Tr;z¼0:022 ¼ 1255 ± 6 K and Tr;z¼0:330 ¼ 1291 ± 8 K, respectively. The spatially-averaged reactive element temperature decreases as reduction proceeds from hTir ¼ 1278 ± 12 K at t ¼ 40 s to hTir ¼ 1274 ± 12 K at t ¼ 240 s due to the endothermic reduction reaction.

J.R. Fosheim et al. / Energy 169 (2019) 597e612

At t ¼ 240 s, operation is switched to oxidation. The inlet flow 0 CO2 increases rapidly and stabilizes at n_ CO ;in;ox

rate of

607

the average values for each of the ten cycles at a 95% confidence interval. The uncertainties in reported nominal temperatures are dominated by measurement uncertainty. The measurement uncertainty and precision error for the average specific molar rates, nonstoichiometry swings, and performance metrics are of the same order of magnitude. The transient behavior of the reactor over ten cycles for each test is shown in Fig. 6. As shown in Fig. 6, thermal steady state conditions were achieved and fuel production is consistent over repeated cycling during Test 1 (left column) and Test 2 (right column). The temperatures corresponding to the inlet and outlet of the ceria bed, the spatiallyaveraged reactive element temperature, and the spatially-averaged cavity temperature are plotted in Fig. 6(a). The gas temperatures at the inlet and outlet of the reactive element are plotted in Fig. 6(b). The heat recovery effectivenesses for reduction and oxidation are shown in Fig. 6(c). The average outlet species molar rates for reduction are shown in Fig. 6(d) and for oxidation in Fig. 6(e). (Measurement uncertainty is indicated on each plot.) The cycleaveraged temperatures of the reactive element, reactor cavity, and inlet and outlet gases remain constant over the ten cycles (within measurement uncertainty). The cyclic peak-to-trough temperatures, previously discussed for the representative cycles in Section 3.1, are invariant from cycle to cycle. Likewise, average heat recovery effectivenesses and fuel production rates remain constant within measurement uncertainty over the ten cycles. For reduction at 1228 K, the average production rates of H2 and CO are 2.1 and 1.03 mmol s1 g1 CeO2 , respectively, and increase to 3.6

2

¼ 7.1 ± 0.1 mmol s1 g1 CeO2 . Similar to when operation was switched

from oxidation to reduction, the outlet flow rates of product gases fluctuate for 30e60 s after the gas valves are switched due to the response of the downstream pressure controller and vacuum pump. Oxidation products are first detected by the gas analyzer at t ¼ 275 s. From 240 s  t  275 s, the outlet gas comprises reduction products. From 275 s  t  290 s, the outlet flow rates of CH4 and H2 decrease and CO and CO2 increase as the outlet gas composition transitions from reduction to oxidation products. After t ¼ 290 s, the outlet gas comprises exclusively oxidation products. The relative flow rates of CO and CO2 during oxidation reflect the CO2 0 conversion rate. CO production reaches a peak value of n_ CO;out;ox ¼ 6.6 ± 0.1 mmol s1 g1 CeO2 at t ¼ 295 s. The slight decrease in CO2 conversion (i.e. the sight increase in the outlet CO2 flow rate and slight decrease in the outlet CO flow rate) from 295 s  t  360 s is due to the reduced thermodynamic impetus for CO2 splitting as oxygen vacancies in the ceria lattice are filled. Oxidation is terminated at t ¼ 360 s prior to complete oxidation of the ceria bed. The cycle is operated with short oxidation durations to achieve high CO2 conversion during oxidation and high H2 and CO selectivities during reduction [9]. The average production rate of CO over the full oxidation period is 5.5 ± 0.1 mmol s1 g1 CeO2 . The reactive element temperatures slowly rise as the oxidation step precedes due to the exothermic reaction. The spatially-averaged reactive element reaches 1277 ± 12 K at t ¼ 360 s. For Test 1 (Fig. 5), the nominal temperature of the reactive

 element is T r ¼ 1228 ± 12 K with input power of 1.5 ± 0.1 kW. These data follow the same transient trends as the representative cycle shown in Fig. 4. The impact of the reduced input power and temperatures on fuel production and performance metrics are discussed in Section 3.2.

and 1.61 mmol s1 g1 CeO2 at 1274 K. The CH4 conversion increased

from 0.25 at 1228 K to 0.36 at 1274 K. H2 and CO selectivity are 0.76 and 0.76, respectively, at 1228 K and increase to 0.90 and 0.82, respectively, at 1274 K. The increase in CH4 conversion and H2 and CO selectivities with increasing temperature is consistent with expectations based on chemical thermodynamics [8,17,28] and previously reported rates for fixed-beds of ceria particles [9,17]. The reduction products of both tests have a H2/CO ratio of ca. 2, making them desirable feedstocks for gas-to-liquids processes. For oxidation, the average production rate of CO is 4.8 mmol s1 g1 CeO2 at

3.2. Performance over ten cycles

1228 K and increases to 5.5 mmol s1 g1 CeO2 at 1274 K, corresponding

The reactor was operated for ten contiguous cycles for both Tests 1 and 2. The average performance during each test is listed in Tables 2 and 3. Table 2 includes the average outlet specific molar rates and the bed-averaged nonstoichiometry swings over reduction and oxidation. Table 3 lists the average performance metrics for both tests, including the CH4 conversion, H2 and CO selectivities, CO2 conversion, energetic upgrade factor, heat recovery effectiveness for reduction and oxidation, solar-to-fuel efficiency, and thermal efficiency. The uncertainties reported in the tables are the root sum square of the measurement uncertainty and precision of

to an increase in CO2 conversion from 0.61 to 0.69. The increase in CO2 conversion with increasing temperature is attributed to faster kinetics [37] and the thermodynamic benefit of larger reduction nonstoichiometries. The increase in CO2 conversion along with the increase in CH4 conversion and H2 and CO selectivities with increasing temperature improve the energetic upgrade factor from 1.07 at 1228 K to 1.10 at 1274 K. For both tests, the average gasphase heat recovery is 0.955 during reduction and 0.985 during

Table 2 Average outlet specific species molar rates and nonstoichiometry swings over ten cycles for the two steady state tests. Test

Average Outlet Specific Molar Rates [mmol s1 g1 CeO2 ] Reduction

1 2

Nonstoichiometry Swing [molOvac mol1 Ce ]

Oxidation

CH4

H2

H2O

CO

CO2

C

CO2

CO

C

〈Dd〉rd rd

〈Dd〉ox ox

4.1 ± 0.1 3.5 ± 0.1

2.1 ± 0.2 3.6 ± 0.1

0.7 ± 0.3 0.4 ± 0.3

1.03 ± 0.07 1.61 ± 0.04

0.30 ± 0.02 0.20 ± 0.01

0.0 ± 0.1 0.2 ± 0.2

2.8 ± 0.1 2.2 ± 0.1

4.8 ± 0.1 5.5 ± 0.1

0.5 ± 0.2 0.6 ± 0.2

0.09 ± 0.01 0.10 ± 0.01

0.08 ± 0.01 0.09 ± 0.01

Uncertainty is the root sum square of the measurement uncertainty and precision of the average values for each of the ten cycles at a 95% confidence interval. Table 3 Average performance metrics over ten cycles for the two steady state tests. Test

XCH4 []

SH2 []

SCO []

XCO2 []

U []

εrd []

εox []

hsol [%]

hth [%]

1 2

0.25 ± 0.02 0.36 ± 0.02

0.76 ± 0.07 0.90 ± 0.06

0.76 ± 0.07 0.82 ± 0.06

0.61 ± 0.02 0.69 ± 0.01

1.07 ± 0.02 1.10 ± 0.0 3

0.955 ± 0.005 0.955 ± 0.005

0.985 ± 0.005 0.985 ± 0.005

5±2 7±2

19 ± 2 25 ± 2

Uncertainty is the root sum square of the measurement uncertainty and precision of the average values for each of the ten cycles at a 95% confidence interval.

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Fig. 6. Data for ten consecutive solar chemical-looping methane reforming cycles for Test 1 (left column) and Test 2 (right column): (a) reactive element temperatures corresponding to the inlet and outlet of the ceria bed, the spatially-averaged reactive element temperature, and the spatially-averaged cavity temperature; (b) gas temperatures at the inlet and outlet of the reactive element; (c) average heat recovery effectiveness for each reduction and oxidation; (d) average outlet specific species molar rates for each reduction; (e) average outlet specific species molar rates for each oxidation.

oxidation. The solar-to-fuel efficiency increases from 5% at 1228 K to 7% at 1274 K. The thermal efficiency increases from 19% to 25%. For both tests, the cycle-averaged material balances on oxygen and carbon are within measurement uncertainty (hDdird þ hDdiox

¼ 0.01 ± 0.01, n_ C;net;cycle ¼ 0.1 ± 0.1 mmol s1 g1 CeO2 ), indicating repeatable cycling of the ceria bed with reversible oxygen exchange and no accumulation of carbon in the reactor. Carbon balances over each reaction step indicate that a small amount of carbon may have

J.R. Fosheim et al. / Energy 169 (2019) 597e612

deposited during reduction but is gasified in the subsequent oxidation. Additionally, the bed-averaged nonstoichiometry swings for both tests are equal within measurement uncertainty (jhDdifrd;oxg j ¼ 0.09 ± 0.01), implying that oxygen exchange between the ceria particles and the bulk gas phase is likely not a ratelimiting factor. This finding is consistent with prior suggestions that syngas production via sCL-MR is limited by surface kinetics [16,38e40].

3.3. Energy balance An energy balance on the reactor is performed to quantify how the radiative input power is used in the reported experiments with a single reactive element. The results of the energy balance are extrapolated to predict efficiencies for full-scale operation of the prototype reactor and for a larger-scale commercial solar reactor with reduced thermal losses (see section 3.4). Fig. 7 shows the components of the solar-to-fuel and thermal efficiencies for both steady state tests with a single reactive element. The left bar for each test is the heating value of the produced syngas quantified separately for H2 and CO produced during reduction and CO produced during oxidation. The middle-left bar is the heating value of the converted CH4. The middle-right bar contains the components of the radiative input, including the fractions used for chemistry, sensible heating of the gases flowing through the reactive element, sensible heating of the N2 flowing through the five passive elements, radiative losses through the aperture, and conduction/ convection losses from the reactor. The right bar is the estimated solar heat for vacuum pumping. The solar-to-fuel efficiency is the difference between the heating value of the syngas product (left bar) and the heating value of the converted CH4 (middle-left bar) divided by the sum of the radiative input (middle-right bar) and the equivalent solar heat for vacuum pumping (right bar). The thermal efficiency is the heating value of the syngas product (left bar) divided by the sum of the heating value of the converted CH4 (middle-left bar), the radiative input (middle-right bar), and the solar heat for vacuum pumping (right bar). At 1228 K, with a radiative input of 1.53 kW, 0.27 kW of CH4 is converted into 0.13 and 0.07 kW of H2 and CO, respectively, during reduction. During oxidation, 0.15 kW of CO is produced. The net energetic gain in the products is 0.08 kW. Conduction/convection

609

losses account for 1.06 kW (69% of Q_ solar ), while 0.34 kW (22% of Q_ solar ) is lost by radiative reflection and emission from the reactor cavity. Sensible heating requirements are 0.01 kW for the reactive element and 0.05 kW for the five passive elements (4% of Q_ solar

total). The low sensible heating requirements are due to highly effective gas-phase heat recuperation. Without heat recuperation, the sensible heating load for the reactive element would be to 0.12 kW. The solar heat to drive the vacuum pump is less than 1% of the total input. Increasing the nominal temperature to 1274 K with an increase in the radiative input to 1.63 kW provides a large benefit in performance primarily due to the proportionally high energetic gain in the products (42% increase compared to 1228 K) with an increase in the radiative input of only 7%. At 1274 K, 0.39 kW of CH4 is converted into 0.50 kW of syngas, for a net energetic gain of 0.11 kW. The conduction/convection are 1.06 kW (65% of Q_ ) and the solar

radiative losses are 0.39 kW (24% of Q_ solar ). Sensible heating requirements are 0.01 kW and 0.06 kW for the reactive element and passive elements, respectively (4% of Q_ total). Without heat solar

recuperation, the sensible heating requirement for the reactive element would be 0.13 kW. The solar heat for vacuum pumping remains less than 1% of the total input. The energy balance shows that for operation with a single reactive element, the thermal losses from the reactor are the largest fraction of the input power for both tests. High thermal losses are typical of prototype reactors because of the high surface-to-volume ratio. Thus, efficiency will increase markedly with implementation of all six reactive elements (2016 g of ceria total) and even more so for a larger commercial-scale reactor. 3.4. Performance projections The extrapolated performance of the reactor for operation with all six reactive elements is shown in Fig. 8. The projected performance is for operation at 1274 K (Test 2), which provided the best performance for single-element operation. With a radiative input of 2.15 kW, the projected energetic content of the converted CH4 is 2.35 kW and the energetic content of the produced syngas is 3.03 kW. The projected solar-to-fuel efficiency is 31% and the thermal efficiency is 67%. Conduction/convection losses are still the most significant energy sink, accounting for 49% of Q_ . Reflection solar

Fig. 7. . Energetic breakdown for the two steady state tests. For each test, the left bar breaks down the heating value of the produced syngas. The middle-left bar is the heating value of the converted CH4. The middle-right bar breaks down the total radiative input into the components of the reactor energy balance. The right bar is the solar power to provide vacuum pumping work.

610

J.R. Fosheim et al. / Energy 169 (2019) 597e612

Fig. 8. Projected energetic breakdown for full scale operation with six reactive elements based on the results of Test 2. From left to right, the first bar breaks down the heating value of the produced syngas. The second bar shows the heating value of the converted CH4. The third bar breaks down the total radiative input into the components of the reactor energy balance. The fourth bar is the solar power to provide vacuum pumping work.

and emission from the solar cavity account for 18% of Q_ solar . The solar energy required for vacuum pumping is insignificant in comparison to the other terms (2% of total input). Due to the highly effective gas-phase heat recuperator, sensible heating of the gas constitutes only 2% of Q_ . Without heat recuperation, sensible solar

heating would increase to 27% of Q_ solar and the projected solar-tofuel and thermal efficiencies would decrease to 23 and 57%, respectively. Performance projections are also made for a representative commercial-scale solar reactor with thermal loss fractions of F ¼ 0.05 and F ¼ 0.20. A reduction of conduction/convection losses to 20% of the radiative input would yield a solar-to-fuel efficiency of 48% and a thermal efficiency of 80%. At 5% thermal loss, the solarto-fuel efficiency would be 56% and the thermal efficiency would be 85%.

reforming in one tube assembly and to apply these data to project efficiency for operation of the reactor with all six tube assemblies and for a larger-scale commercial solar reactor with reduced thermal losses. The reduction step was conducted with 75% CH4, which to-date is the highest CH4 concentration that has been utilized for solar chemical-looping methane reforming. Reduction with high CH4 concentrations is important as the use of an inert carrier gas can introduce sizeable economic and energetic costs. Gas-phase heat recuperation was utilized to reduce sensible heating requirements and improve efficiency for the first time for solar chemical-looping methane reforming. Operating conditions were selected to enable contiguous operation of the full cycle with the aim of demonstrating the commercial viability of the approach. The reported data provide favorable results. Reactor performance was consistent during the two steady state tests with repeatable fuel production, reversible oxygen exchange, and no accumulation of carbon over ten cycles. At 1274 K, CH4 conversion is 0.36, H2 selectivity is 0.90, CO selectivity is 0.82, CO2 conversion is 0.69, and the upgrade factor is 1.10. The H2/CO ratio of the syngas is near 2, which is suitable for the synthesis of liquid hydrocarbon fuels. Heat recovery effectiveness is 0.955 during reduction and is 0.985 during oxidation. The measured solar-to-fuel and thermal efficiencies are 7% and 25%, respectively, for single-element operation. Based on an energy balance on the reactor during singleelement operation, extrapolated solar-to-fuel and thermal efficiencies for full-scale operation of the prototype reactor are 31 and 67%, respectively. Heat recuperation raises the projected solar-tofuel efficiency from 23 to 31% and the projected thermal efficiency from 57 to 67%. The higher efficiencies compared to prior work are attributed in part to the selection of operating conditions that achieve high H2 and CO selectivity and CO2 conversion while avoiding carbon accumulation and effective recuperation of the sensible heat of the product gases. If the reactor was scaled up to a commercial scale, the solar-to-fuel and thermal efficiency could reach 56 and 85%, respectively. The present work represents a major step towards assessing the scalability and commercial viability of solar chemical-looping methane reforming. Repeatable performance of sCL-MR during steady operation was demonstrated for the first time in a solar reactor. The reported solar-to-fuel efficiencies are the highest reported to-date for this process. The projected efficiency for a scaled-up commercial reactor suggests solar chemical-looping methane reforming could be a competitive approach for production of solar fuels.

4. Conclusion Solar chemical looping methane reforming is a two-step thermochemical cycle that promises to be a highly efficient means of storing solar energy in chemical bonds and producing H2 and CO for gas-to-liquid fuel synthesis and other industrial processes. In the present work, chemical-looping methane reforming is implemented in a prototype fixed-bed solar reactor operated in an indoor high-flux solar simulator. In the reactor, CH4 and CO2 flow alternately over a packed bed of ceria particles (336 g) within the annulus of a concentric assembly of alumina tubes. The tube assembly extends outside the solar cavity to integrate with a ceramic gas-phase heat recuperator. In the recuperator, the internal passages of the concentric tube are filled with reticulated porous alumina to enhance heat transfer. The reactor was operated at thermal steady state conditions for ten cycles at 1228 and 1274 K with solar concentrations at the aperture of 840 and 900 suns, respectively. The objectives of the experiments were to characterize the thermal and chemical performance of the reactor for operation of chemical-looping methane

Acknowledgement This work was supported by the University of Minnesota Institute on the Environment. Jesse R. Fosheim is supported by the National Science Foundation Graduate Research Fellowship Program (NSF-GRFP) under Grant 00039202. The authors acknowledge the assistance of Nathaniel J. Lewin during experiments in the solar simulator. Nomenclature

Latin Aap F h HHV mCeO2

Area of the aperture [m2] Thermal loss fraction [] Partial molar enthalpy [J mol1] Higher heating value [J mol1] Mass of ceria [g]

J.R. Fosheim et al. / Energy 169 (2019) 597e612

MCeO2 n_ p Q_ Ru S t to T hTi U x X z

Molar mass of ceria [172.115 g mol1] Molar flow rate [mol s1] Pressure [kPa] Cycle-averaged heat rate [W] Universal gas constant, [8.314 J mol1 K1] Selectivity [-] Time [s] Cycle start time [s] Temperature [K] Spatially-averaged temperature [K] Energetic upgrade factor [] Mole fraction [] Conversion [] Axial coordinate [m]

d

[3]

[4]

[5]

[6]

[7]

Greek hDdi

[2]

[molOvac mol1 CeO2 ]

Nonstoichiometry Bed-averaged nonstoichiometry swing

[8]

[molOvac mol1 CeO2 ] εa ε

h s t

Apparent emissivity [] Gas-phase heat recovery effectiveness [] Efficiency [] Stefan-Boltzmann constant [5.67E-8 W m2 K1] Duration [s]

[9]

[10]

[11]

Superscripts 0 Mass specific ½g1 CeO2 

[12]

Temporal-average Subscripts amb Ambient conditions (100 kPa, 298.15 K) cav Pertaining to the reactor cavity chem Chemical reaction commercial Commercial-scale solar reactor full-scale Full-scale operation of the prototype reactor i Species i j General index in Inlet of the reactor loss Conduction/convection to ambient out Outlet of the reactor ox Oxidation p Pertaining to the passive element(s) pump Electric-to-pumping work conversion rad Radiative emission and reflection r Pertaining to the reactive element rd Reduction ref Reference flow res Residence time s/e solar-to-electric conversion sens Sensible heat sol Solar-to-fuel conversion solar Radiative input syn Syngas th Thermal energy conversion vac Vacuum Abbreviations PPI Pores per inch RPC Reticulated porous ceramic sCL-MR Solar chemical-looping methane reforming

[13]

[14]

[15]

[16]

[17]

[18]

[19]

[20]

[21]

[22]

[23]

[24]

[25]

References [26] [1] Muhich CL, Ehrhart BD, Al-Shankiti I, Ward BJ, Musgrave CB, Weimer AW.

611

A review and perspective of efficient hydrogen generation via solar thermal water splitting. Wiley Interdiscip Rev Energy Environ 2016;5:261e87. https:// doi.org/10.1002/wene.174. Agrafiotis C, Roeb M, Sattler C. A review on solar thermal syngas production via redox pair-based water/carbon dioxide splitting thermochemical cycles. Renew Sustain Energy Rev 2015;42:254e85. https://doi.org/10.1016/ j.rser.2014.09.039. Scheffe JR, Steinfeld A. Oxygen exchange materials for solar thermochemical splitting of H2O and CO2: a review. Mater Today 2014;17:341e8. https:// doi.org/10.1016/j.mattod.2014.04.025. Roeb M, Neises M, Monnerie N, Call F, Simon H, Sattler C, et al. Materialsrelated aspects of thermochemical water and carbon dioxide splitting: a review. Materials (Basel) 2012;5. https://doi.org/10.3390/ma5112015. 2015e54. Miller JE, McDaniel AH, Allendorf MD. Considerations in the design of materials for solar-driven fuel production using metal-oxide thermochemical cycles. Adv Energy Mater 2014;4:1300469. https://doi.org/10.1002/ aenm.201300469. Wright HA, Allison JD, Jack DS, Lewis GH, Landis SR. ConocoPhillips GTL technology: the COPox process as the SynGas generator. ACS Div Fuel Chem Prepr 2003;48:791e2. Wood DA, Nwaoha C, Towler BF. Gas-to-liquids (GTL): a review of an industry offering several routes for monetizing natural gas. J Nat Gas Sci Eng 2012;9: 196e208. https://doi.org/10.1016/j.jngse.2012.07.001. Krenzke PT, Fosheim JR, Davidson JH. Solar fuels via chemical-looping reforming. Sol Energy 2017;156:48e72. https://doi.org/10.1016/ j.solener.2017.05.095. Krenzke PT, Fosheim JR, Zheng J, Davidson JH. Synthesis gas production via the solar partial oxidation of methane-ceria redox cycle: conversion, selectivity, and efficiency. Int J Hydrogen Energy 2016;41:12799e811. https://doi.org/ 10.1016/j.ijhydene.2016.06.095. Zinkevich M, Djurovic D, Aldinger F. Thermodynamic modelling of the ceriumeoxygen system*. Solid State Ionics 2006;177:989e1001. https:// doi.org/10.1016/j.ssi.2006.02.044. Kümmerle E, Heger G. The structures of CeCe2O3þd, Ce7O12, and Ce11O20. J Solid State Chem 1999;147:485e500. https://doi.org/10.1006/ jssc.1999.8403. Panlener RJ, Blumenthal RN, Garnier JE. A thermodynamic study of nonstoichiometric cerium dioxide. J Phys Chem Solid 1975;36:1213e22. https:// doi.org/10.1016/0022-3697(75)90192-4. Mogensen M. Physical, chemical and electrochemical properties of pure and doped ceria. Solid State Ionics 2000;129:63e94. https://doi.org/10.1016/ S0167-2738(99)00318-5. Otsuka K, Ushiyama T, Yamanaka I. Partial oxidation of methane using the redox of cerium oxide. Chem Lett 1993;22:1517e20. https://doi.org/10.1246/ cl.1993.1517. Otsuka K, Sunada E, Ushiyama T, Yamanaka I. The production of synthesis gas by the redox of cerium oxide. Stud Surf Sci Catal 1997;107:531e6. https:// doi.org/10.1016/S0167-2991(97)80386-2. Otsuka K, Wang Y, Sunada E, Yamanaka I. Direct partial oxidation of methane to synthesis gas by cerium oxide. J Catal 1998;175:152e60. https://doi.org/ 10.1006/jcat.1998.1985. Warren KJ, Reim J, Randhir K, Greek B, Carrillo R, Hahn DW, et al. Theoretical and experimental investigation of solar methane reforming through the nonstoichiometric ceria redox cycle. Energy Technol 2017;5:2138e49. https://doi.org/10.1002/ente.201700083. Welte M, Warren K, Scheffe JR, Steinfeld A. Combined ceria reduction and methane reforming in a solar-driven particle-transport reactor. Ind Eng Chem Res 2017;56:10300e8. https://doi.org/10.1021/acs.iecr.7b02738. Chuayboon S, Abanades S, Rodat S. Syngas production via solar-driven chemical looping methane reforming from redox cycling of ceria porous foam in a volumetric solar reactor. Chem Eng J 2018;356:756e70. https:// doi.org/10.1016/j.cej.2018.09.072. Wilhelm D, Simbeck D, Karp A, Dickenson R. Syngas production for gas-toliquids applications: technologies, issues and outlook. Fuel Process Technol 2001;71:139e48. https://doi.org/10.1016/S0378-3820(01)00140-0. Hathaway BJ, Bala Chandran R, Sedler S, Thomas D, Gladen A, Chase T, et al. Effect of flow rates on operation of a solar thermochemical reactor for splitting CO 2 via the isothermal ceria redox cycle. J Sol Energy Eng 2015;138. https://doi.org/10.1115/1.4032019. 011007. Hathaway BJ, Bala Chandran R, Gladen AC, Chase TR, Davidson JH. Demonstration of a solar reactor for carbon dioxide splitting via the isothermal ceria redox cycle and practical implications. Energy Fuels 2016;30:6654e61. https://doi.org/10.1021/acs.energyfuels.6b01265. Bala Chandran R, Davidson JH. Model of transport and chemical kinetics in a solar thermochemical reactor to split carbon dioxide. Chem Eng Sci 2016;146: 302e15. https://doi.org/10.1016/j.ces.2016.03.001. Krueger KR, Davidson JH, Lipinski W. Design of a new 45 kW[sub e] high-flux solar simulator for high-temperature solar thermal and thermochemical research. J Sol Energy Eng 2011;133. https://doi.org/10.1115/1.4003298. 011013.  ski W, Davidson JH. Operational performance of the uniKrueger KR, Lipin versity of Minnesota 45 kWe high-flux solar simulator. J Sol Energy Eng 2013;135. https://doi.org/10.1115/1.4023595. 044501. Banerjee A, Bala Chandran R, Davidson JH. Experimental investigation of a reticulated porous alumina heat exchanger for high temperature gas heat

612

[27]

[28]

[29]

[30]

[31]

[32]

J.R. Fosheim et al. / Energy 169 (2019) 597e612 recovery. Appl Therm Eng 2015;75:889e95. https://doi.org/10.1016/ j.applthermaleng.2014.10.033. Bala Chandran R, Smith RM De, Davidson JH, Bala R, Smith RM De, Davidson JH, et al. Model of an integrated solar thermochemical reactor/ reticulated ceramic foam heat exchanger for gas-phase heat recovery. Int J Heat Mass Tran 2015;81:404e14. https://doi.org/10.1016/ j.ijheatmasstransfer.2014.10.053. Krenzke PT, Davidson JH. Thermodynamic analysis of syngas production via the solar thermochemical cerium oxide redox cycle with methane-driven reduction. Energy Fuels 2014;28:4088e95. https://doi.org/10.1021/ ef500610n. Bala Chandran R, De Smith RM, Davidson JH, Smith RM De, Davidson JH, Bala R, et al. Model of an integrated solar thermochemical reactor/reticulated ceramic foam heat exchanger for gas-phase heat recovery. Int J Heat Mass Tran 2015;81:404e14. https://doi.org/10.1016/ j.ijheatmasstransfer.2014.10.053. Marxer D, Furler P, Takacs M, Steinfeld A. Solar thermochemical splitting of CO 2 into separate streams of CO and O 2 with high selectivity, stability, conversion, and efficiency. Energy Environ Sci 2017;10:1142e9. https://doi.org/ 10.1039/C6EE03776C. Marxer D, Furler P, Scheffe J, Geerlings H, Falter C, Batteiger V, et al. Demonstration of the entire production chain to renewable kerosene via solar thermochemical splitting of H 2 O and CO 2. Energy Fuels 2015;29:3241e50. https://doi.org/10.1021/acs.energyfuels.5b00351. Bulfin B, Call F, Lange M, Lübben O, Sattler C, Pitz-Paal R, et al. Thermodynamics of CeO2 thermochemical fuel production. Energy Fuels 2015;29: 1001e9. https://doi.org/10.1021/ef5019912.

[33] Mancini T, Heller P, Butler B, Osborn B, Schiel W, Goldberg V, et al. Dishstirling systems: an overview of development and status. J Sol Energy Eng 2003;125:135. https://doi.org/10.1115/1.1562634.  ski W, Steinfeld A. Heat transfer model of a solar receiver[34] Schunk LO, Lipin reactor for the thermal dissociation of ZnOdexperimental validation at 10kW and scale-up to 1MW. Chem Eng J 2009;150:502e8. https://doi.org/ 10.1016/j.cej.2009.03.012. [35] Hirsch D. Solar hydrogen production by thermal decomposition of natural gas using a vortex-flow reactor. Int J Hydrogen Energy 2004;29:47e55. https:// doi.org/10.1016/S0360-3199(03)00048-X. phane A, Florent L, Gilles F. Analysis of solar chemical processes [36] Charvin P, Ste for hydrogen production from water splitting thermochemical cycles. Energy Convers Manag 2008;49:1547e56. https://doi.org/10.1016/ j.enconman.2007.12.011. [37] Ackermann S, Sauvin L, Castiglioni R, Rupp JLMM, Scheffe JR, Steinfeld A. Kinetics of CO 2 reduction over nonstoichiometric ceria. J Phys Chem C 2015;119:16452e61. https://doi.org/10.1021/acs.jpcc.5b03464. [38] Fathi M, Bjorgum E, Viig T, Rokstad O. Partial oxidation of methane to synthesis gas. Catal Today 2000;63:489e97. https://doi.org/10.1016/S09205861(00)00495-8. [39] Pantu P, Kim K, Gavalas GR. Methane partial oxidation on Pt/CeO2eZrO2 in the absence of gaseous oxygen. Appl Catal Gen 2000;193:203e14. https:// doi.org/10.1016/S0926-860X(99)00429-9. [40] Warren KJ, Scheffe JR. Kinetic insights into the reduction of ceria facilitated via the partial oxidation of methane. Mater Today Energy 2018;9:39e48. https:// doi.org/10.1016/j.mtener.2018.05.001.