Solar-driven co-thermolysis of CO 2 and H 2 O and in-situ oxygen removal across a non-stoichiometric ceria membrane

Abstract

We report on the first-ever experimental demonstration of simultaneous thermolysis of CO2 and H2O with in-situ separation of fuel and oxygen in a solar-driven membrane reactor. Gaseous CO2/H2O mixtures at molar ratios from 3:4 to 2:1 were fed to a mixed ionic-electronic conducting non-stoichiometric ceria (CeO2-δ) membrane enclosed in a solar cavity-receiver and exposed to simulated concentrated solar radiation of up to 4200 suns. Reaction rates were measured at isothermal and isobaric conditions in the range of 1723-1873 K and 0.2-1.7 Pa O2, yielding a maximum combined CO and H2 fuel production rate of 2.3 µmol cm⁻² min⁻¹ at 1873 K and 0.2 Pa O2 at steady state, which corresponded to a conversion of reactants of 0.7%. At all conditions tested, CO production was favored over H2 production, as expected from theory. Experimental results followed the same trends as the thermodynamic equilibrium limits of membrane-assisted thermochemical fuel production.

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1
Solar-driven co-thermolysis of CO2 and H2O and in-situ oxygen removal
across a non-stoichiometric ceria membrane
Maria Tou1, Jian Jin2,3, Yong Hao2,3, Aldo Steinfeld1, Ronald Michalsky1*
1 Department of Mechanical and Process Engineering, ETH Zürich, 8092 Zürich, Switzerland
2 Institute of Engineering Thermophysics, Chinese Academy of Sciences, 11 Beisihuanxi Rd.,
Beijing 100190, P. R. China
3 University of Chinese Academy of Sciences, No.19A Yuquan Rd., Beijing 100049, P. R. China
*corresponding author
Abstract
We report on the first-ever experimental demonstration of simultaneous thermolysis of CO2 and
H2O with in-situ separation of fuel and oxygen in a solar-driven membrane reactor. Gaseous
CO2/H2O mixtures at molar ratios from 3:4 to 2:1 were fed to a mixed ionic-electronic
conducting non-stoichiometric ceria (CeO2-δ) membrane enclosed in a solar cavity-receiver and
exposed to simulated concentrated solar radiation of up to 4200 suns. Reaction rates were
measured at isothermal and isobaric conditions in the range of 1723-1873 K and 0.2-1.7 Pa O2,
yielding a maximum combined CO and H2 fuel production rate of 2.3 µmol cm-2 min-1 at 1873 K
and 0.2 Pa O2 at steady state, which corresponded to a conversion of reactants of 0.7%. At all
conditions tested, CO production was favored over H2 production, as expected from theory.
Experimental results followed the same trends as the thermodynamic equilibrium limits of
membrane-assisted thermochemical fuel production.
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2
Introduction
The utilization of the vast solar energy resource for electricity, heat, and fuels has become a key
objective in research and development.1 The conversion and storage of solar energy in fuels is
especially appealing as a means to transition from fossil fuels to a “CO2 economy”.2 For this, a
solar refinery needs to be developed where solar energy is collected and used to convert CO2 and
H2O to fuels by some method. Existing research spans solar-driven electrochemical, photo-
electrochemical, and photocatalytic paths for direct conversion, as well as indirect routes via the
solar thermochemical production of syngas (H2 and CO).3
Solar thermochemical redox cycles utilize the entire spectrum of solar radiation concentrated to
high-temperature process heat to drive the splitting of CO2 and H2O and produce CO and H2 at
high rates, selectivity, mass conversions, and efficiencies.4-6 However, the temperature swing
required between the redox steps induce significant material stresses and energy irreversibilities,
which prompted the search for alternative isothermal processes.7-9 One promising approach is the
use of a dense, ceramic, mixed ionic-electronic conducting (MIEC) membrane for the continuous
separation of oxygen and fuel (H2 and/or CO) derived from the thermolysis of CO2 and H2O at
high temperatures, as pioneered for solar water splitting by Fletcher and co-workers.10, 11 We
recently demonstrated the proof-of-concept utilizing a solar-driven membrane reactor for
splitting of CO2.12 Other investigations of thermochemical membrane reactors, both theoretical
and experimental, have also only focused on either CO2- or H2O-splitting.10, 13-18 This work goes
further and demonstrates the feasibility of co-feeding both CO2 and H2O and assesses the relative
favorability between the two thermolysis reactions occurring simultaneously. The desired
dissociations are chemical equilibrium reactions in the gas phase described by:
(1)
12
2 2
CO CO + O
(2)
12
2 2 2
H O H + O
The reactions are analogous; that is, both are endothermic and thermolytic, but their reaction
energetics differ. This is described by the standard Gibbs free energy changes at equilibrium
(ΔG=0):
(3)
2
2
121
( )
CO O 2
1 1
CO
ln ln
p p
G RT p RT K
p

 

 
  
 
 
o
o
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3
(4)
2 2
2
121
( )
H O 2
2 2
ln ln
H O
p p
G RT p RT K
p

 

 
  
 
 
o
o
where ΔGi° is related to the equilibrium constant Ki which in turn is a function of temperature
only. Figure 1a shows T- pO2 equilibrium contours of the separate thermolysis of CO2 and H2O
for various mole fractions of CO or H2 in the product gas, respectively. For both CO2 and H2O
thermolysis, products are favored with increasing T and decreasing pO2, i.e. higher mole fractions
of fuel are possible towards the upper-left corner of Fig. 1a. At such high T, dissociation of CO2
is more thermodynamically favorable than that of H2O at equal conditions.19, 20 A decrease in pO2
can be achieved without use of high-value electrical energy by removal of O2 utilizing a dense
membrane made of an oxygen-selective MIEC material.21 The pO2 is controlled to a low value on
the opposite side of the membrane. For each of the separate thermolysis reactions, if pO2 ≥ ½ pCO
or pO2 ≥ ½ pH2, the membrane provides no benefit.
The membrane reactor concept used in this work for the co-thermolysis of H2O and CO2 is
shown schematically in Figure 1b. CO2 and H2O are fed to the inner side of a capped tubular
non-stoichiometric ceria membrane. Ceria has become the benchmark material for oxygen-
cycling applications due to its stability and fast kinetics.22-24 In our previous work it was also
found to be an effective material for oxygen-conducting membranes.12 The supply of
concentrated solar process heat at high temperatures drives the thermolysis, producing CO, H2,
and O2. The latter adsorbs at the inner membrane surface, dissociates, and is transported across
the membrane in an ionic form along a chemical potential gradient. The O2- then associates into
O2 at the outer membrane surface and desorbs into an inert sweep gas contained in a shell tube.
This in-situ removal of one of the reaction products drives the reactions forward towards
dissociation and avoids downstream recombination. The counter-flow configuration of the
reactant and product gases favorably maximizes the gradient of pO2 along the length of the
membrane. By placing this reactor in a solar cavity-receiver, the high-temperature heat for the
reactions is provided by concentrated solar radiation incident on the shell tube.
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4
A B
Fig. 1 (a) Equilibrium contours for separate thermolysis of CO2 (light-colored) or H2O (dark-
colored) at 1 bar total pressure as a function of T and pO2 (according to Eqs. 3 and 4) for various
mole fractions of CO or H2, respectively. Contours extend until pCO or pH2 = 2· pO2, increasing
pO2 beyond this point no longer provides benefit over unperturbed thermolysis. (b) Schematic of
the tubular redox membrane reactor for splitting of CO2 and H2O. CO2 and H2O are fed to the
inner side of the membrane and dissociate into fuel and O2, with the latter selectively crossing
the membrane into Ar sweep gas.
Typical operating conditions demand temperatures around 1773 K and partial pressures of O2
down to 1 Pa. These high temperatures eliminate the need for catalysts but pose significant
constraints on the construction materials which must withstand these conditions over extended
periods of time. Materials must also resist thermal shock that may occur due to cooling
overnight, unless the reactor is equipped with an alternative heat source such as a high-
temperature thermal energy storage system.25 A modular tubular membrane design could avoid
costly maintenance by allowing for simple replacement of degraded membranes. Maintaining
low partial pressures of oxygen is crucial, requiring additional energy for vacuum pumping or
gas separation to regenerate the inert sweep gas (such as N2, though here we use Ar for gas
analytic considerations).26, 27 Alternatively, some studies have reported solar-driven production
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5
of pure O2 and inert gas with low partial pressures of oxygen using thermochemical oxygen
pumps driven by low-grade process heat.27-29
Experimental
Materials. Cerium (IV) oxide (ceria, CeO2, powder, particle size <5 μm, 99.9% purity), poly (oxy-
1,4-phenylene sulfonyl-1,4-phenylene) (PES, (C12H8O3S)n, pellets), polyvinylpyrrolidone (PVP,
(C6H9NO)n, powder, average M.W. 40,000), and 1-methyl-2-pyrrolidinone (NMP, C5H9NO,
liquid, ≥99.0% purity) were from Sigma Aldrich. Al2O3 membranes (Alsint 99.7, 7 mm outer
diameter (OD), 5 mm inner diameter (ID), 250 mm length) were from Intertechno-Firag AG. High-
purity alumina adhesive (Aremco Ceramabond 569) and glass-filled sealant (Aremco-Seal 617)
were from Kager Industrieprodukte GmbH. CO2 (99.998%), Ar (99.996%, 99.999%), He
(99.999%), and calibration gas mixtures, i.e., 1000 mol ppm H2 (99.999%) and 500 mol ppm CO
(99.997%) in Ar (99.999%), and 1000 mol ppm CO (99.997%), 500 mol ppm CO2 (99.995%), 500
mol ppm N2 (99.999%), and 100 mol ppm O2 (99.999%) in Ar (99.999%) were from Messer
Schweiz AG. According to the manufacturer, Ar (99.996%) contained < 5 ppm O2 on volume
basis, equivalent to a limiting pO2 < 0.5 Pa.
Membrane Fabrication. Capped tubular ceria membranes were produced as reported previously.12
Briefly, membranes were fabricated using a phase-inversion/sintering method.30, 31 Two polymers,
namely PES (5.7 wt%) and PVP (0.5 wt%), were dissolved in NMP (22.0 wt%). Ceria powder
(71.8 wt%) was suspended in the polymer solution. The ceria slurry was coated onto membrane
templates (High-Flexible silicone tubing, 3 mm ID, 7 mm OD, RCT Reichelt Chemietechnik
GMbH & Co.), which were placed into a water bath for phase inversion (unfiltered tap water
coagulant at ambient conditions). The silicone templates were removed, and the dried membrane
precursors were then sintered for 8 hours at 1873 K (oven model HTL 20/17, ThermConcept). The
sintered membranes were 6-7 mm OD, 5-6 mm ID, and 150-250 mm in length. Typically, the
membrane walls were about 0.5 mm thick. In-depth solid-state characterization of ceria
membranes before and after use in the reactor was performed previously.12 SEM analysis in Fig.
S1 shows that ceria membranes exposed to both CO2 and H2O in thermolysis experiments do not
change morphologically, consistent with membranes used in pure-CO2 experiments.
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Experimental setup. The solar membrane reactor system is depicted schematically in Figure 2. CO2
and/or H2O was fed through a feeder tube into the inner side of the membrane while Ar sweep gas
was fed into the reactor shell tube in a counter-current flow. This assembly was placed in a
thermally insulated solar cavity-receiver with an aperture 4 cm in diameter. A compound parabolic
concentrator (CPC) was incorporated onto the aperture to boost the solar flux concentration and
generate a more uniform directional distribution of concentrated radiation entering the cavity.32
Experimentation was performed at the High-Flux Solar Simulator (HFSS) of ETH Zurich: an array
of seven Xe arcs, close-coupled to truncated ellipsoidal reflectors, provided an external source of
intense thermal radiation that closely approximated the heat transfer characteristics of highly
concentrating solar energy facilities. The radiative flux distribution at the focal plane was
measured optically using a calibrated CCD camera focused on a Lambertian (diffusely reflecting)
target. The solar radiative power input to the cavity was calculated by integration of the radiative
flux over the aperture area and verified with a water calorimeter. Temperatures were measured at
the outer surface of the reactor shell at two heights along the tube (indicated in Figure 2) using B-
type thermocouples. Gas flow rates were regulated by electronic mass flow controllers (MFC,
Bronkhorst F-201 C, accuracy 0.5%Rd + 0.1%FS), whereas steam flow was generated by a liquid
flow controller (LFC, Bronkhorst Liqui-Flow L23-AAD-33-K-305, accuracy 1%FS) and steam
generator (Bronkhorst CEM W-202A-333-K). Product gas composition in each stream was
monitored on-line by gas chromatography (GC, Agilent 490 MicroGC).
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Fig. 2 Schematic and flow diagram of the experimental setup used to test co-thermolysis of CO2
and H2O in a membrane reactor. Simulated concentrated solar radiation from the HFSS enters the
cavity-receiver through the aperture and heats the reactor. Gaseous CO2 and/or H2O flows through
a feeder tube into the membrane, then flows upward through the annulus between the membrane
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and feeder tubes before exiting the reactor. The membrane removes O2 produced from thermolysis.
Unreacted H2O is removed in the condenser and the composition of the remaining gas is analyzed
by GC1. In counter-current to the oxidant stream, a sweep gas (Ar) flows through the shell tube,
taking up O2 crossing the membrane, and exits below the cavity-receiver to be analyzed by GC2.
Not to scale.
Experimental runs. All volumetric flow rates are given at standard conditions (1 bar and 273 K).
The HFSS heated the reactor to the desired temperatures in the range 1723-1873 K with a radiative
power input of 2.5-3.0 kW. For water-splitting experiments, steam was fed to the inner side of the
membrane at a rate of 140 mL min−1 (5 g h−1 water to steam generator), carried in 80-100 mL
min−1 Ar. For co-feed experiments, the same flow of steam was carried by 75-200 mL min-1 CO2.
At the outer side of the membrane, the reactor shell was purged with 200-1000 mL min−1 Ar. The
compositions of both gas streams exiting the solar reactor were analyzed simultaneously using two
gas chromatographs (GC). Steady state was defined as the condition at which the measured gas
concentration was within 2% of the mean over the previous five consecutive measurements
collected at a frequency of one every two minutes:
(5)
 
 
 
1
5
1
5
1
50.02
1
5
n
i n i j
j n
n
i j
j n
c t c t
c t

 

 




where ci(tj) is the volumetric concentration of species i at time point j. Steady-state data was
collected for at least 19 minutes and the arithmetic mean was used to summarize the results at
each experimental condition.
Thermodynamic Analysis
We calculated the thermodynamic equilibrium limits of thermolysis of CO2 and H2O in a
membrane reactor to compare to experimental results. Relatively fast rates are expected for each
reaction step: gas-phase thermolysis, heterogeneous surface reaction, and oxygen bulk diffusion.
When each of these serial processes are sufficiently fast, the global kinetics are fast, and
thermodynamics govern the net reaction. In this case, kinetics can be neglected.
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Previous observations with a solar cavity-receiver containing a porous ceria structure directly
exposed to high-flux irradiation reported that the overall kinetics are not controlled by solid-state
diffusion within the crystal lattice.33 This is also expected for a ceria membrane because the
measured values of ambipolar diffusion coefficients of oxygen in ceria (1.5×10-5-4×10-4 cm2 s-1
in the range 1673-1823 K22) translate to diffusion times on the order of seconds for the length
scales across the 0.5 mm-thick membrane. Thus, as far as solid-state diffusion is concerned, the
transport of oxygen vacancies through the membrane is almost instantaneous compared with the
time scales of data collection.
Reaction rates have an exponential dependence on temperature, scaling with exp(-EA/RT), as
seen in the Arrhenius equation. The high temperature in the range 1723-1873 K and consequent
high activity of reactive oxygen vacancies at the surface of the membrane are expected to lead to
fast surface exchange of oxygen from the gas into the solid phase.19 While studies of non-
isothermal processes show that heating rate limits the oxygen release rate, heat transfer should
not be limiting in this isothermal process because the heat of reaction is much lower than the heat
input.23, 33 Likewise, the high temperature, along with the catalytic effect of ceria, implies very
fast thermolysis reactions in the gas-phase.34 Therefore, a purely thermodynamic model is
expected to adequately predict reactor performance.
However, to reach the thermodynamic limit in the countercurrent-flow reactor, there must be
sufficient membrane area and sweep gas relative to the flowrate of reactant. To account for the
oxygen capacity in a given flow of sweep gas, a thermodynamic model described by Bulfin was
applied, which is specific to countercurrent-flow reactors.35 This approach guarantees
compliance with the second law of thermodynamics and conservation of mass along the entire
reactor by means of a dimensionless oxygen exchange coordinate, κ, defined as the number of
moles of O2 crossing the membrane up to a certain point along the length, x, per mole of oxidant
fed:
(6)
 
2
O
0
oxidant
( ) ,
xj x dx
xn


&
where ṅoxidant is the molar flow rate of H2O and/or CO2 (ṅCO2+ṅH2O), and jO2(x) is the molar flux of
O2 from the oxidant flow to the sweep gas as a function of the length along the membrane. Then
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pO2 in each flow can be formulated as a function of κ, and pO2,oxidant is determined by the
thermodynamic equilibrium of thermolysis, as described by equations 3 and 4 in the
introduction. For O2 to spontaneously flow from the oxidant flow to the sweep gas, partial
pressures must satisfy pO2,oxidant(κ) > pO2,sweep(κ), and equilibrium is reached if there exists a value
of κ in [0, κtotal] where pO2,oxidant(κ) = pO2,sweep(κ). Bulfin details the full methodology in an
example tailored to thermolysis in a membrane reactor with a counter-current sweep flow.35
The countercurrent-flow thermodynamic model was implemented and solved numerically in
Matlab using thermodynamic data from NIST JANAF.36 The model input parameters are T, p,
pO2, the relative flow of sweep to oxidant (ṅsweep/ṅoxidant), and the relative flow rates of CO2 and
H2O (ṅCO2/ṅH2O). Note that pO2 refers to the O2 impurity at the inlet of the sweep gas. The values
were generally chosen to match experimental conditions as determined from mass flow
controller, thermocouple, and GC measurements. For example, the reactor operated at ambient
pressure and 1 bar total pressure was set on both sides of the membrane. It was not possible to
measure the reaction temperature of the gas inside the membrane. Instead, the measured
temperature at the outer wall of the shell tube (certainly greater than the reaction temperature)
and a simple heat transfer model were used to estimate the reaction temperature. Consequently,
this analysis presents the thermodynamic limits as a shaded region, where the upper and lower
bounds are the limits at the maximum measured shell temperature and the estimated reaction
temperature, respectively. In the range of operation, the difference between these temperatures is
approximately 50 K. The simple heat transfer model used to calculate this temperature difference
is described in the Electronic Supplemental Information (ESI). The outputs of the
thermodynamic model are the partial pressures of the products and conversion of the reactants at
equilibrium.
Results and Discussion
Figure 3 summarizes the steady-state specific production rates measured experimentally for fuel,
comprising of CO (light shading) and H2 (dark shading), and O2 as a function of three process
variables: (a) time, (b) T, and (c) pO2. These experiments used a roughly equimolar feed of CO2
and H2O and produced fuel with a relatively higher fraction of CO than H2. This product
proportion is consistent with the more favorable change in Gibbs free energy for thermolysis of
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CO2 at high temperatures seen in Figure 1a. The average molar ratio O2:fuel over 19 experiments
was 0.53 ± 0.07, corroborating a closed mass balance. Furthermore, no other byproducts were
detected, indicating the absence of undesired side reactions. Specifically, Figure 3a shows
production rates over time at steady-state at a nominal T of 1873 K and pO2 of 0.4 Pa. Note that T
refers to the maximum measured shell temperature and pO2 refers to the partial pressure of O2 at
the Ar inlet, which is the minimum pO2 in the system and determined by the impurity in the
sweep gas. Gas evolution rates leveled off at a constant incident radiative flux of about 3500
suns (1 sun = 1 kW/m2) over the cavity’s aperture, and thus the reactor demonstrated continuous,
steady-state operation at isothermal conditions. Figure 3b shows the steady-state average
production rates as a function of T in the range 1723-1873 K. Mass flow rates were kept constant
at 5 g h-1 H2O with 100 mL min-1 CO2 to the inner side and 500 mL min-1 Ar to the outer side of
the membrane (L denotes standard liters). The measured pO2 ranged from 0.4 – 1.7 Pa due to
variation in the small amount of air leakage into the reactor during different experimental runs.
The specific fuel and O2 production rates at steady state increased with temperature, in
accordance with the thermodynamic dependence on exp(-ΔG°/RT). Finally, Figure 3c shows the
steady-state average production rates as a function of pO2 in the range 0.2 – 0.9 Pa at 1873 K.
Although the range of pO2 tested was small, the gas production rates indeed decreased at higher
pO2, as expected from the equilibrium relationship for the thermolysis of CO2 and H2O described
in equations 3 and 4. However, to achieve this range of pO2, the flow rate of Ar was adjusted,
thus changing the relative flow of sweep gas, ṅsweep/ṅoxidant, which is also a key thermodynamic
parameter in this reactor system.
A B C
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Fig. 3 Trends in steady-state specific gas production rates. (a) Steady-state production rates of
fuel (CO and H2) and O2 vs. time at 1873 K and 0.4 Pa O2. Steady-state average production rates
of fuel and O2 vs. (b) T, keeping all mass flow rates constant (c) pO2, keeping T constant at 1873
K. Feed was equimolar CO2 and H2O in all cases. Fuel production is composed of two
contributions distinguished by shade: CO (light) and H2 (dark).
Figure 4 shows the experimental steady-state conversion of CO2 to CO and H2O to H2 as a
function of (a) T in the range 1723-1873 K at pO2 = 1 Pa, ṅsweep/ṅoxidant = 2.4, and (b) ṅsweep/ṅoxidant
ranging 1-5 at T = 1873 K, pO2 = 0.5 Pa. The thermodynamic limits are also indicated, denoted by
a band whose upper and lower bounds are calculated as described above at Tshell and the
estimated reaction temperature, Tshell – 50 K, respectively. The band is wider for CO2 than H2O
because its reaction favorability changes more steeply with T. Like the gas production rates,
conversion of reactants in Figure 4a increased with T, as expected. The conversion also increased
at higher relative sweep rates in Figure 4b, because ṅsweep/ṅoxidant determines the total amount of
O2 that can be removed across the membrane at equilibrium. In fact, the thermodynamic analysis
reveals that the trend in production rates observed in Figure 3c is more attributable to varying
ṅsweep/ṅoxidant than pO2. The pO2 in the sweep gas must be lower than pO2 in the oxidant stream at all
points along the membrane to drive transfer of O2. At sufficiently low pO2, however, this
parameter does not have a strong influence on the conversion of reactants, shown in Figure S3 in
the ESI. In the range of pO2 observed in experiments (0.2-0.9 Pa), the theoretical conversion of
each reactant at constant ṅsweep/ṅoxidant is almost flat, while the experimental data points exhibit a
trend due to varying relative sweep rates. In contrast, the same experimental data plotted against
ṅsweep/ṅoxidant in Figure 4b match the shape of the equilibrium limit curves.
In general, the results indicate that the reactor performance indeed approaches the
thermodynamic limit for a countercurrent flow reactor. Importantly, the experimental conversion
does not exceed the theoretical limit. Furthermore, the experimental points lie closer to the lower
bound of the equilibrium region, suggesting that the simple heat transfer model is necessary and
effective to estimate the reaction temperature. While the experimental conversion of H2O closely
follows the lower bound of the predicted thermodynamic limit, the experimental conversion of
CO2 falls short. The discrepancy is less than a factor of two and may be a result of the water-gas
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shift (WGS) reaction occurring at lower T downstream of the reactor, before the GC analysis.
WGS consumes some CO to produce additional H2 and is thus the difference of the CO2 and
H2O dissociation reactions (Eq. 2 minus Eq. 1):
(7)
2 2 2
CO + H O H + CO
In this case, the GC measurements may not be representative of the composition in the reactor;
the actual conversion of CO2 may have been higher, and the conversion of H2O lower,
potentially equalizing the difference seen in Figure 4 between the experimental data and the limit
for each gas.
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Fig. 4 Comparison of average experimental conversion of CO2 and H2O at steady state versus
equilibrium limits as a function of (a) T, at ṅsweep/ṅoxidant = 2.4 and 1 Pa O2, (b) ṅsweep/ṅoxidant, at
1873 K and 0.5 Pa O2.
Figure 5 illustrates the effect of the molar feed ratio, CO2:H2O, on the: (a) steady-state average
production rates of fuel and O2, (b) conversions of CO2 and H2O, and (c) molar ratio of the two
fuels produced, CO:H2. Figures 4b and c compare the results measured experimentally with
those predicted from thermodynamics. The relative flow rates of CO2 and H2O were varied while
maintaining steam feed rate at 5 g h-1 H2O, temperature at 1873 K, and pO2 at 0.5 Pa. Figure 4a
shows that the overall fuel production (sum of H2 and CO) increased with CO2:H2O, as did the
production rate of CO, which occurred because the total feed rate and CO2 feed rate both
increased. The production rate of H2, on the other hand, decreased with CO2:H2O. It must be
emphasized that ṅsweep/ṅoxidant did not remain constant over experimentation, but rather decreased
with CO2:H2O because the flow rate of sweep gas remained constant. A constant flowrate of
sweep gas is less effective at maintaining low pO2 as the amount of O2 to be removed increases.
In the case of CO, the effect of increasing CO2 feed rate compensated for the decreasing sweep
ratio, so that the net production rate increased. However, the H2 production decreased.
The confounding factors of changing both the total feed rate and the relative sweep rate are
accounted for by plotting conversion instead of production rate in Figure 5b. As in Figure 4, the
absolute values of H2O conversion match the model results better than for CO2. The
experimental conversions of both CO2 and H2O decreased slightly with CO2:H2O, in agreement
with the trend predicted at equilibrium. There are two contributions to the negative trend in
conversion. First, as already mentioned, ṅsweep/ṅoxidant decreased with CO2:H2O, which decreased
the conversion. However, conversion of each reactant is predicted to decrease slightly with
CO2:H2O even with a constant ṅsweep/ṅoxidant. The second, smaller reason for the trend is related to
the difference in favorability of thermolysis of CO2 and H2O. As the feed ratio increases, a
higher proportion of the feed is CO2, which has a higher conversion than H2O at these
conditions. In fact, the overall conversion of reactants to products actually increased slightly with
CO2:H2O. However, the O2 capacity of the sweep gas was unchanged and therefore the
conversion of each individual reactant must decrease to balance the production of O2 with its
removal. This result indicates that a feed of CO2 requires a higher relative sweep rate than an
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equal feed of H2O, because its higher potential conversion leads to a larger amount of O2 that
must be removed.
Consistent with Figure 5a, Figure 5c shows that the product ratio CO:H2 increased with feed
ratio CO2:H2O, as expected intuitively. Interestingly, CO:H2 is always greater than the
corresponding CO2:H2O, which further confirms that dissociation of CO2 is more favorable than
that of H2O at equivalent conditions. The observed experimental trend qualitatively matches
equilibrium thermodynamics, though with a smaller slope because the experimental conversion
of CO2 is lower than predicted. In consideration of downstream processing, the Fischer-Tropsch
synthesis favors a syngas fed with 1:2 moles CO:H2.38, 39 According to Figure 5c, the product
ratio can be adjusted via the feed ratio, and a 1:2 product ratio would require a feed ratio smaller
than the minimum tested here.
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Fig. 5 Effect of molar feed ratio CO2:H2O on: (a) steady-state average production rate of fuel and
O2, where fuel is composed of two contributions distinguished by shade: CO (light) and H2
(dark); (b) molar conversion of CO2 and H2O calculated from experiments and their limits at
equilibrium; and (c) molar fuel ratio CO:H2 measured experimentally and at equilibrium.
Reaction conditions were ṅsweep/ṅoxidant varying from 1.6-2.5 and constant 5 g h-1 H2O feed rate,
1873 K, and 0.5 Pa O2.
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The maximum conversions observed experimentally were 1.0% CO2 and 0.4% H2O at 1873 K,
ṅsweep/ṅoxidant = 5, and 0.2 Pa O2 (0.7% overall conversion of reactants). In general, the absolute
values of both theoretical and experimental conversion were lower in the co-feed case tested here
than in the pure-CO2 feed case tested previously,12 because the relative sweep rates were lower in
this set of experiments. The base case ṅsweep/ṅoxidant was 8 in pure-CO2 experiments and 2.4 in
these co-feed experiments. The relative sweep rate is a significant thermodynamic parameter for
sweep gas operation, and the application of the countercurrent flow model was essential to
accurately predict the behavior of the reactor. Furthermore, although there was no effort to
optimize the efficiency of the reactor in these proof-of-concept experiments, ṅsweep/ṅoxidant also
impacts efficiency because it determines how much sweep gas must be heated and circulated per
unit fuel produced.
Conclusion
We have demonstrated steady-state splitting of a mixed feed containing CO2 and H2O into
separate streams of syngas fuel and O2 using an isothermal tubular ceria membrane reactor
driven by simulated concentrated solar radiation. The experimental results generally agreed with
trends predicted by thermodynamics. The conversion of CO2 to CO was favored over H2O to H2,
consistent with the energetics of the respective thermolysis reactions. The co-thermolysis of a
mixture of CO2 and H2O is more complex than feeding either CO2 or H2O separately to the
reactor, both experimentally and in the theoretical analysis. In the co-feed case, the mixture of
CO2, H2O, CO, and H2 could undergo additional reactions, especially the reverse water-gas shift
(RWGS). As seen in equation 7, the RWGS reaction is not independent of the two thermolysis
reactions, and thus does not change the equilibrium conversion of each reactant from the pure-
feed values. Other possible reactions were found to be negligible from the lack of byproducts
predicted by thermodynamics and measured experimentally at these conditions. Absence of
byproducts and a 2:1 ratio of fuel:oxygen together confirmed 100% selectivity for the desired
splitting reactions.
The solar thermochemical membrane reactor unifies both CO2 and H2O splitting in a single
modular and scalable device and offers a technically viable pathway to single-step syngas
production. However, determining an appropriate relative sweep rate is challenging in co-feed
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operation because the different favorability for thermolysis of CO2 and H2O implies different
optimums for each species. In addition, these energetic differences mean that H2O needs to be
fed in large excess to achieve a desirable syngas composition. Therefore, it may still be attractive
to produce CO and H2 separately and mix them into syngas as needed. Furthermore, the single-
step approach embodied in the membrane reactor must compete with multistep cycles currently
available. Thus, further R&D and alternative membrane configurations are needed to boost mass
conversions and consequently reach favorable solar-to-fuel energy efficiencies, a challenge
because T and pO2 determine the thermodynamic limits.
Conflicts of Interest
There are no conflicts to declare.
Acknowledgements
This work was funded by the Swiss National Science Foundation (Ambizione Energy Grant No.
166883) and by the Chinese Academy of Sciences (International Collaboration Key Program
award no. 182211KYSB20160043). We thank Brendan Bulfin for fruitful discussions and
Patrick Basler and Simon Minder for supporting the experimental campaign.
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Tou et al
Table of Contents entry
First-time experimental demonstration of simultaneous thermolysis of CO2 and H2O in a solar-
driven membrane reactor
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