Reactive Particles for Efficient Solar Thermochemical Hydrogen Production

REACTIVE PARTICLES FOR EFFICIENT SOLAR THERMOCHEMICAL HYDROGEN PRODUCTION

BRIAN DAVID EHRHART

B.S., Rensselaer Polytechnic Institute, 2010

M.S., University of Colorado, 2013

A thesis submitted to the

Faculty of the Graduate School of the

University of Colorado in partial fulfillment

of the requirement for the degree of

Doctor of Philosophy

Department of Chemical and Biological Engineering

2016

This thesis entitled:

Reactive Particles for Efficient Solar Thermochemical Hydrogen Production

written by Brian David Ehrhart has been approved for the Department of Chemical and Biological Engineering

______

Alan W. Weimer, Committee Chair

______

David E. Clough, Committee Member

Date ______

The final copy of this thesis has been examined by the signatories, and we find that both the content and the form meet acceptable presentation standards

of scholarly work in the above mentioned discipline.

Ehrhart, Brian David (PhD, Chemical and Biological Engineering)

Reactive Particles for Efficient Solar Thermochemical Hydrogen Production

Thesis directed by Professor Alan W. Weimer

Hydrogen is very important and useful for chemical products, ammonia production, and as a transportation fuel. Future production will need to be done renewably and carbon-free to avoid further damage from climate change. Production of hydrogen directly from solar energy has the potential to be highly efficient. Some metal oxides can be thermally reduced in one reaction and then re-oxidized with steam to produce hydrogen in a two-step reduction/oxidation cycle.

A comprehensive solar-to-hydrogen (STH) efficiency model is developed for two-step solar thermochemical redox water splitting. Two redox materials are considered and compared in order to assess the impact of the rate of oxidation and the hydrogen productivity per cycle on STH efficiency. Near-isothermal redox processing is beneficial for materials with slower kinetics, especially with moderate to high gas heat recuperation. Gas heat recuperation is critical for high efficiency cycles, especially at conditions that lead to high steam and inert gas flowrates. Three methods for achieving low oxygen partial pressures for reduction are compared, and the effect of vacuum pump efficiency and inert gas/oxygen separation efficiency are quantified. Currently available vacuum pump technologies have very low thermodynamic efficiencies at low pressures and are unlikely to provide efficient hydrogen production relative to other oxygen partial pressure lowering technologies. A novel recycled inert gas sweep with high temperature separation is proposed and STH efficiency values are shown to vary significantly depending on the inert gas flowrate required. A high separation temperature for the recycled inert gas is beneficial, especially for cases of lower gas heat recuperation and increased inert gas flowrates.

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Particles are an important aspect of many proposed solar thermochemical water splitting reactor concepts. These particles must maintain physical integrity and chemical performance at very high temperatures while moving around a system, which has the potential for highly efficiency continuous operation. Spray drying is used to produce iron aluminate (hercynite) particles using a pH-modified charge-stabilized sol. Nanoparticle suspensions are mixed and the pH is modified in order to induce partial flocculation. This procedure gives larger particles that are more spherical and structurally robust. Manganese oxide-based mixed metal oxide particles have been designed and tested for thermochemical energy storage. Al2O3, Fe2O3, and ZrO2 are tested in

Mn2O3-based spray-dried particles, and results are compared with thermodynamic predictions. A mixture of 2:1 Fe2O3:Mn2O3 formed iron manganese oxide spinel (FeMn2O4) on calcination, and demonstrated the highest thermochemical activity despite particle agglomeration and deformation.

Conversely, zirconia was an inert support that does not react with manganese oxide. Differences in redox performance between materials with different Fe to Mn ratios have been attributed to ion diffusion and secondary phase formation.

Co-doped hercynite materials are examined for two-step solar thermochemical water splitting. Density functional theory (DFT+U) for iron and cobalt aluminate predict that electron density transfer on the creation of oxygen vacancies is almost exclusively to the atoms that are first-nearest-neighbor to the oxygen vacancy. Cases that have a single cobalt next to the oxygen vacancies tend to be very favorable in terms of both a lower oxygen vacancy formation energy and also a lower host structure energy, while sites with more than one Co next to the O-vacancy tend to be less favorable, due to cobalt being less favorable on the octahedral sites.

Cobalt appears to preferentially change oxidation state on formation of oxygen vacancies.

Current XPS results seem to contradict the computational results, as the Co-doped hercynite

sample does not show Co oxidation state changes while the un-doped hercynite shows Fe oxidation state changes. However, excess cobalt in the Co-doped sample makes oxygen vacancies less favorable, while samples with less cobalt should show more oxidation state change for Co. Un- doped hercynite has demonstrated more hydrogen production capacity but slower reaction rates than the Co-doped hercynite. This could be due to kinetic/catalytic impacts of Co on the reaction mechanism, but also partially affects the thermodynamics of the reaction enthalpy change. This work suggests that lowering the amount of Co in the material or by introducing new dopants that can assume the octahedral site more easily will improve the hercynite cycle.

To my wife Sarah

And

To my parents David and Julie

ACKNOWLEDGMENTS

There are many people that have been an enormous help over the years. I do not have the space here to fully express my gratitude, but I want to quickly list a number of important specific contributions. First, I would like to thank some of my external collaborators. Eric Coker at Sandia

National Laboratories was very helpful in discussing TGA measurements and helping perform

HT-XRD measurements. Yahya Al-Salik from SABIC was very helpful in performing and discussing HT-XPS measurements and results. Michael Takacs at ETH Zürich was very helpful in hosting and supporting me during experiments in Switzerland. Ivan Ermanoski from Sandia

National Laboratories was very helpful and patient in explaining and discussing thermodynamics and other aspects of system efficiency calculations.

There are many people within the University of Colorado system that have been enormously helpful. Paul Boni of the Department of Geological Sciences and Ilya Lisenker in the

Stoldt Lab in Mechanical Engineering at CU Boulder were both very helpful in training and support for XRD measurements. Various staff at the Nanomaterials Characterization Facility

(NCF) at CU Boulder were very helpful for training and troubleshooting SEM images. Fred

Luiszer in the Department of Geology was incredibly helpful in performing measurements and answering questions about ICP measurements. Outside of CU Boulder, Aditya Gandhi in the

Skaggs School of Pharmacy and Pharmaceutical Sciences at the CU Anschutz Medical Campus was very helpful in providing training and support for the zeta potential measurements.

There are also many people within the Chemical and Biological Engineering department at CU Boulder that have been hugely helpful to me. Dragan Mejic is incredibly helpful with just about everything in lab; whether it is fixing something that is broken or building/modifying a piece

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of equipment, Dragan can get it done. Dana Hauschulz was enormously helpful to me in helping to diagnose any sort of electrical issue, reading wiring diagrams, come up with modifications for power requirements to custom equipment, answering questions, and even wiring up some equipment for me when we were up against a deadline. Colleen Courtney in the Chatterjee Lab was very helpful with centrifuging samples and discussing zeta potential measurements. Ben

Richardson in the Anseth lab was extremely generous in running viscosity measurements for me.

Finally, Dominque De Vangel is an absolute epitome of competency and helpfulness, and was helpful beyond words in navigating requirements, rules, and answering every single question I could come up with on any topic.

Many undergraduate students have worked with me over the last few years, and have all contributed greatly to the success of my dissertation. The senior design group that helped with part of my master’s degree ended up having even more of a contribution, as they found some of the first literature I had read about solid electrolyte oxygen separation membranes: Gabriel Draper,

Vinh Le, and Nigel Wang. Kayla Weston was an all-star researcher who did much of the earlier work for Chris Muhich on the chemistry of the hercynite cycle, and helped me out quite a bit in my early days in the lab. Vanessa Witte worked with me personally the longest, and was incredibly helpful and a hard worker in making and characterizing all kinds of powders. Ben Mousseau was also critical in getting some of our spray dried powders made and characterized, and figuring out how to work the sonic sifter. Kevin Sun was very generous with his time in running many different

XRD samples for me. Aiden Coffey was also very generous with his time, and very understanding of my somewhat hectic schedule in the final months of my graduate school career.

Team Weimer is an amazing research group to be a part of, and many members have contributed a lot to my work. Aside from being able to bounce ideas off of a kind and helpful

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audience, people have answered questions, analyzed samples, and helped me out when I needed it. Dr. Torrie Aston was around for my first few years in the group, and made it a very welcoming place; she helped me a lot in lab, helped me get started on a number of topics, and together we started TWITT. Dr. Janna Martinek was always very kind and available to discuss various minutia about national labs, and later helped me quite a bit in discussing thermal losses from solar receivers. Dr. Darwin Arifin also helped me get going, and had some very important insights to solar thermochemistry. Dr. Alia Lubers was a very close friend and helped get TWITT and lab chores going; she was always very helpful and made sure people were being safe in lab. Dr. Chris

Muhich was there to get me started my very first time in lab, and subsequently we became both very good friends and very close colleagues; from attending conferences, writing papers, and trying to get some science done in-between, Chris was instrumental to my dissertation. Dr. Casey

LaMarche was helpful for a number of troubleshooting discussions in lab, and let me use the centrifuge on short notice. Scott Rowe was very helpful in setting up and diagnosing lab computers, LabView software, and we had a few very interesting discussions about numerical solvers. Ibraheam Al-Shankiti has worked very closely with me in a variety of situations, from setting up efficiency computer models to testing at NREL to showing me around Saudi Arabia, and has always been helpful. BJ Ward worked closely with and put up with me for a long time, and we got a lot of very good work done for many different projects. Ryan Trottier was helpful in running, parsing, and answering questions about DFT calculations. Samantha Miller has been a huge help with running samples in lab, and especially in running and helping me understand some critical DFT calculations in the final months of my dissertation. Amanda Hoskins has been very helpful in lab, in many, many discussions, and helping to make the STCH project a success. Boris

Chubukov and Wilson McNeary were also very helpful in lab and in a number of discussions on

results. Chip Fisher, Mark Wallace, and Caitlin Majlinger were very helpful in grading sections of the senior design tests and final reports during my Advanced TA. Caitlin was also incredibly helpful in almost every lab technique, and in running technoeconomic models for multiple projects.

Research would be impossible without funding, and so I would like to acknowledge my funding sources. I have been supported by the U.S. Department of Energy (DOE) Fuel Cell

Technologies Program through the Solar Thermochemical Hydrogen (STCH) directive and the

U.S. DOE Office of Energy Efficiency and Renewable Energy (EERE), Fuel Cell Technologies

Office under Award Number DE-EE0006671. I also want to gratefully acknowledge financial support from Award P200A120125 of the U.S. Department of Education Renewable and

Sustainable Energy Graduate Assistance in Areas of National Need (GAANN) Program. Finally,

I have been supported by funding from the Saudi Basic Industries Corporation (SABIC).

Additionally, this work utilized the Janus supercomputer, which is supported by the NSF (award number CNS-0821794) and the University of Colorado Boulder. The Janus supercomputer is a joint effort of the University of Colorado Boulder, the University of Colorado Denver and the

National Center for Atmospheric Research.

All of the members of my thesis committee have been very helpful over the past few years in providing invaluable feedback. They have all been very generous with their time and understanding of various requirements of my leave of absence. Dr. Andrea Ambrosini of Sandia

National Laboratories has shared some very useful insights in inorganic chemistry over the course of many discussions. Dr. Hicham Idriss from SABIC has also met with me multiple times and provided extremely useful feedback, in addition to providing support for XPS measurements and hosting me on a visit to KAUST. Dr. Ronggui Yang has provided some very useful feedback, and has been very flexible in meeting with me during his busy travel schedule. Dr. David Clough taught

one of the single most useful classes I have ever taken, and means I have a much better understanding of statistics and how they apply to research; he has furthered this by ensuring I have a good understanding of various experimental designs for my work. Dr. Charles Musgrave has been very helpful in keeping me focused on my dissertation, and very understanding when the work took a different path than we originally planned.

I want to especially thank my adviser and mentor, Dr. Alan Weimer. Al originally accepted and helped get me into the Master’s program, taking on the risk of committing to a student that he himself had not vetted. Al successfully guided me through my M.S., and enthusiastically accepted me back for a Ph.D. Throughout both, Al has always placed a great deal of trust in me. We have written multiple proposals together, and one or two even got funded! He has sent me to multiple meetings and conferences every year, when many graduate students are lucky to go on only a couple of these trips. Not only that, he has sent me abroad twice, once to Zürich, Switzerland and once to Jeddah, Saudi Arabia in order to meet with and learn from our collaborators. He has put me in touch with many different people form industry, government, and academia. He has let me work independently and explore things for myself, which while occasionally stressful is an excellent way to learn about topics more fully. He has never stopped supporting me throughout my graduate career.

Apart from my many professional colleagues, I could not have done any of this without my family and friends. My parents David and Julie Ehrhart have always taught me to learn and be curious, and the importance of hard work. Thank you both so much. My brother Mark and sister

Megan have always been very supportive; they are on their way to their own highly educated and technical careers, and I know they will do well. My good friends Ryan Mowbray, Kate (who also talked with me a lot about flocculation) and Luke Sooy, Kyle Lieurance, Vince and Keesha

Dorzweiler, Stacey Skaalure, Kelsey and Dave MacConaghy, Matt McBride, Steele Reynolds, and many others have helped make these years very enjoyable. Thank you all.

Finally, I want to thank my wonderful wife Sarah. She has been my partner for years, and is a great support for me and our dog Conall. She is supportive of everything I do, is a world-class chef, and is generally awesome. She is the most incredible woman in the world. Conall is a strange but amazing dog, always there to make me feel better and reminding me to exercise by taking him on walks all hours of the day and night. Thank you and I love you both.

Thank you all.

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CONTENTS

CHAPTER

I. INTRODUCTION ...... 1

1.1. Overview of Solar Thermochemical Water Splitting ...... 1

1.1.1. Importance of Hydrogen ...... 1

1.1.2. Hydrogen Generation ...... 2

1.1.3. Two-Step Solar Thermochemical Water Splitting Concepts ...... 5

1.2. Two-Step STWS Active Materials ...... 8

1.2.1. Thermodynamics of STWS Materials ...... 8

1.2.2. Current STWS Materials ...... 10

1.2.2.1. Volatile Stoichiometric Chemistries ...... 10

1.2.2.2. Non-Volatile Stoichiometric Reactions ...... 12

1.2.2.3. Oxygen Vacancy Mechanism Reactions ...... 14

1.3. Modes of Operation ...... 21

1.3.1. Reaction Temperature ...... 21

1.3.2. Oxygen Removal ...... 24

1.4. Reactor Design ...... 25

1.4.1. Monolithic Reactors ...... 27

1.4.2. Particle Reactors...... 30

1.4.3. The Solar-thermal Particle Flow Reactor ...... 33

1.5. Project Objectives ...... 35

1.5.1. System Efficiency Objectives ...... 36

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1.5.2. Active Particle Objectives ...... 36

1.5.3. Metal Ion Activity Objectives ...... 37

1.6. References ...... 37

II. EFFICIENCY MODELING ...... 49

2.1. Abstract ...... 49

2.2. Introduction ...... 51

2.3. Methods ...... 56

2.3.1. Efficiency Calculations for Thermodynamic Equilibrium ...... 57

2.3.1.1. Overall System Description ...... 57

2.3.1.2. Reduction and Oxidation Chemistry ...... 58

2.3.1.3. Overall System Efficiency ...... 60

2.3.1.4. Solar Energy Fluxes ...... 61

2.3.1.5. Thermal Energy Required to Produce Hydrogen...... 64

2.3.1.6. Auxiliary Heating Requirements and Benefits ...... 65

2.3.1.7. Recycled Inert Gas Terms ...... 68

2.3.1.8. Vacuum Pump Case ...... 71

2.3.1.9. Cascade Pressure Reduction Case ...... 72

2.3.1.10. Heat Benefits ...... 74

2.3.1.11. Heat Rejection ...... 75

2.3.2. Kinetic Limitations ...... 76

2.3.3. Calculation and Analysis ...... 76

2.4. Results...... 78

2.4.1. Thermodynamic Efficiency of Ceria and Ferrite/Zirconia Cycles ...... 78

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2.4.2. Illustration of Kinetics ...... 81

2.4.3. Inert Sweep Gas Flow Rate ...... 82

2.4.3.1. Effect of Non-Minimum Inert Gas Flowrates on System

Efficiency ...... 83

2.4.3.2. Impact of Gas Heat Recuperation on Non-Minimum Inert Gas

Flowrates ...... 85

2.4.4. Inert Sweep Gas/Oxygen Separation ...... 86

2.4.5. Impact of Water/Hydrogen Separation Temperature ...... 87

2.4.6. Efficiency of Vacuum Pumping ...... 89

2.4.7. Efficiency of Cascade Pressure Reduction ...... 93

2.4.8. Efficiency of Inert Sweep Gas Reduction ...... 96

2.5. Discussion of Model Assumptions ...... 99

2.5.1. Material Specific Effects on Thermodynamic Efficiency ...... 100

2.5.2. Impact of Reaction Rates on Process Design ...... 101

2.5.3. Practicality of Counter-Flow Inert Gas Sweep ...... 102

2.5.4. Implications for High Inert/Oxygen Separation Temperature ...... 102

2.5.5. Practicality of Gas and Solid Heat Recuperation ...... 103

2.5.6. Implications for High Hydrogen/Water Separation Temperatures ...... 105

2.5.7. Feasibility of High Vacuum Pump Efficiency ...... 106

2.5.8. Practicality of Cascade Pressure Reduction ...... 107

2.5.9. Practicality of Inert Gas Sweep Reduction ...... 109

2.6. Conclusions...... 110

2.7. References ...... 113

III. SPRAY DRYING ACTIVE PARTICLES ...... 119

3.1. Abstract ...... 119

3.2. Introduction ...... 120

3.2.1. Active Particles for Solar Thermochemical Water Splitting...... 120

3.2.2. Thermochemical Energy Storage ...... 120

3.2.3. Spray Drying ...... 122

3.2.4. Specific Novelty of This Work ...... 124

3.3. Methods ...... 125

3.3.1. Preparation of Spray Dried Particles ...... 125

3.3.2. Particle Characterization ...... 126

3.4. Results...... 128

3.4.1. Manganese Oxide Particle Characterization ...... 129

3.4.2. Thermodynamic Predictions of Manganese Oxide Particles ...... 134

3.4.3. Gravimetric Measurements of Manganese Oxide Particles ...... 137

3.4.4. Characterization of Hercynite Suspension ...... 138

3.4.5. Hercynite Particle Characterization ...... 144

3.5. Discussion ...... 145

3.5.1. Mass Change in Mixed-Metal Manganese Oxide Particles ...... 145

3.5.2. Effect of Flocculation on Droplet Formation in Hercynite Particles ... 148

3.6. Conclusions...... 150

3.7. References ...... 152

IV. ROLE OF METAL IONS IN HERCYNITE SPINEL ...... 157

4.1. Abstract ...... 157

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4.2. Introduction ...... 158

4.2.1. Doping in Solar Thermochemical Water Splitting Materials ...... 158

4.2.2. Hercynite Cycle History ...... 160

4.3. Method ...... 161

4.3.1. Computational Simulations ...... 161

4.3.2. Experimental Characterization ...... 163

4.4. Results...... 164

4.4.1. Computational Results ...... 165

4.4.2. Particle Characterization ...... 173

4.5. Discussion ...... 175

4.5.1. Differences Between Oxides and Aluminates...... 175

4.5.2. Preferential Charge Transfer On Aluminate Oxygen Vacancy Formation

...... 176

4.5.3. Octahedral and Tetragonal Site Energy Penalties ...... 179

4.5.4. Trade-Off Between Kinetics and Capacity ...... 182

4.6. Conclusion ...... 183

4.7. References ...... 185

V. CONCLUSIONS AND RECOMMENDATIONS FOR FUTURE WORK ...... 188

5.1. Conclusions...... 188

5.1.1. Solar Thermochemical Water Splitting Materials and Reactors ...... 188

5.1.2. System Efficiency Modeling ...... 189

5.1.3. Engineered Particles for High Temperature Solar Thermochemical

Cycles ...... 191

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5.1.4. Role of Metal Ions in Hercynite Thermochemical Cycles ...... 193

5.2. Recommendations for Future Work ...... 194

5.2.1. System-Level Considerations...... 194

5.2.2. Improvements in Active Particles ...... 195

5.2.3. Improved Doping in Aluminate Spinel Materials ...... 196

VI. BIBLIOGRAPHY ...... 198

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TABLES

Table 2.1: Reduction endotherms and heat capacity for reactive solid materials. The references for the ceria values are shown in the table. The values for the ferrite/zirconia composite were estimated using FactSage...... 64 Table 3.1. Overview of composition and sample ID for candidate materials...... 125 Table 3.2. Conditions used in the FactSage thermodynamic equilibrium calculations of manganese oxide-based mixed metal oxides...... 127 Table 3.3. BET surface area (before and after calcining) and particle size distribution for spray dried candidate materials after calcining at 1,200°C...... 132 Table 3.4. Thermodynamic predictions of slagging temperatures of candidate materials in air and inert environments...... 137

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FIGURES

Figure 1.1: Methods of concentrating solar irradiance using a) power tower and heliostats and b) a parabolic dish concentrator [31]. c) schematic of a generic two-step solar thermochemical water splitting cycle...... 6 Figure 1.2: Thermodynamic extent of oxygen non-stoichiometry (δ) of ceria, based on temperature and partial pressure of oxygen [38]...... 16 Figure 1.3: a) The doped-hercynite cycle H2 production rates after reduction at various temperatures and oxidation at 1,000°C. b) the production rates of CoFe2O4 on ZrO2 after reduction and oxidation under the same conditions as a) [103]. c) shows the H2 production rates of doped-hercynite operating under temperature swing water splitting conditions (left two peaks) and isothermal water splitting conditions (right peak) [40]...... 18 Figure 1.4: Perovskite based solar thermal water splitting cycles. SLMA materials are capable of producing significantly more H2 than ceria when reduced at 1,350°C and oxidized at 1,000°C [110]. b) Many other perovskite formulations have shown to be capable of undergoing STWS. From left to right are the H2 production capacities of Ba25Sr75Co80Fe20, Ba50Sr50Co80Fe20, La60Sr40Co20Fe80, LaSrCo, La65Sr35Mn, La50Sr50Mn, and La50Sr50Mn under operating conditions shown in the inset [113]...... 20 Figure 1.5: Three different methods to achieve low pO2 of oxygen and the associated energy requirements: a) vacuum pumping, b) direct inert gas sweep, and c) recycled inert gas sweep...... 25 Figure 1.6: Monolith-based solar thermal water splitting reactor concepts: a) the porous monolith cavity reactor,[32] b) the rotating piston reactor,[65] c) the CR5,[123] and d) the SurroundSun reactor.[124] ...... 28 Figure 1.7: Particle based solar thermal water splitting reactor concepts: a) rotating cavity particle reactor,[135] b) non-rotating particle flow reactor,[131] c) aerosol flow reactor,[137] d) internally circulating fluidized bed reactor,[138] e) moving particle packed bed reactor. [139] The labels in a) and b) point out: a) 1) rotating drum, 2) actuation, 3) aperture, 4) cavity, 5) screw feeder, 6) product outlet port, 7) rotary joint, 8) working fluids, 9) insulation, 10) quartz window; b) 1) water-cooled window mount and vortex-flow generation. 2) water-cooled cavity aperture, 3) BOP and data-acquisition cavity access ports, 4) alumina-tile reaction surface, 5) annular solid ZnO exit, 6) bulk insulation and cavity- shape support and 7) central product-vapor and gas exit...... 31 Figure 1.8: The Solar-thermal Particle Flow Reactor. a) an individual reduction/oxidation reactor unit and b) receiver configuration containing multiple individual reduction/oxidation reactor units. Reactors are not shown to scale...... 34 Figure 2.1: Model process schematic, not to scale. Line arrows indicate mass flows, while block arrows indicate energy flows. A block arrow directed into a block is a heat requirement, a block arrow directed out of a block is a heat benefit. A solid filled block arrow denotes an energy flux per unit time, and empty block arrow denotes an energy requirement per mole of

hydrogen. The grey cross-hatched block arrows denote energy flow per mole of hydrogen that is rejected. The empty block arrows sum to equal the QMOL term...... 58 Figure 2.2: A schematic (not to scale) of a cascade pressure reduction system for NC chambers. Each pressure reduction chamber has an equally sized vacuum pump attached to it, which pumps an equal volumetric flow rate of oxygen at an efficiency based on the pressure of that stage of the cascade. The molar flow rate of oxygen for each chamber is based on the temperature and pressure of that chamber. It is assumed the reactive solids reach thermodynamic equilibrium in each chamber. Chamber 0 is a chamber with no pump for when the equilibrium partial pressure of oxygen is >1 atm...... 72 Figure 2.3: Solar-to-hydrogen efficiency for a) ceria and b) ferrite/zirconia active redox materials at various levels of gas heat recuperation (εGG) and a reduction temperature of 1,600 K, a reduction pressure of 0.1 Pa, and a solid heat recuperation effectiveness of 50%...... 79 Figure 2.4: Heat loads and benefits for a) ceria and b) ferrite/zirconia active redox materials with gas heat recuperation of 90%, a reduction temperature of 1,600 K, a reduction pressure of 0.1 Pa, and a solid heat recuperation effectiveness of 50%...... 79 Figure 2.5: Values for Δδ for ceria and ferrite/zirconia at a reduction temperature of 1,600 K and a reduction pressure of 0.1 Pa...... 81 Figure 2.6: Efficiency values with efficiency discounts for various values of εGG for a) ceria and b) ferrite/zirconia both with a reduction temperature of 1,600 K, a reduction pressure of 0.1 Pa, and a solid heat recuperation effectiveness of 50%...... 82 Figure 2.7: System efficiency (ηSTH) for inert gas sweep system with various multipliers of the minimum inert flowrate. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5...... 83 Figure 2.8: System efficiency (ηSTH) for inert gas sweep system with various lower limits on the minimum inert gas flowrate. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5...... 84 Figure 2.9: System efficiency (ηSTH) for inert gas sweep system with various gas heat recuperation effectiveness values (εGG) with a a) the ideal counter-flow minimum and b) a lower limit of 100 on the inert gas flowrate. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, and a solid heat recuperation effectiveness (εSS) of 0.5...... 85 Figure 2.10: Combined separation and sensible heat required for inert/O2 separation and heating, based on a high separation temperature of 1,223 K and a low temperature of 90 K. All calculations are done for ceria at a reduction temperature of 1,800 K, a reduction partial pressure of 0.1 Pa, and a gas heat recuperation effectiveness of 0.9...... 86 H2 Figure 2.11: Solar-to-hydrogen efficiency for ceria with TSEP = 373 K – 1,000 K for a reduction partial pressure of a) 0.1 Pa and b) 1,000 Pa. All cases have a reduction temperature of 1,800 K and a solid heat recuperation effectiveness of 50%...... 88

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Figure 2.12: Solar to hydrogen thermal efficiency (ηSTH) values for various ΔT values and various gas heat recuperation effective values (εGG). All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 10 Pa, and a solid heat recuperation effectiveness (εSS) of 0.5...... 89 Figure 2.13: ηSTH values for various ΔT values and various reduction pressures (pRED). All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5...... 91 Figure 2.14: ηSTH values for various vacuum pump efficiency values for 1) pRED = 0.1 Pa and b) pRED = 10,000 Pa. All calculations were done for a reduction temperature of 1,800 K, a gas heat recuperation effectiveness of 0.9, and a solid heat recuperation effectiveness of 0.5. .. 92 Figure 2.15: System efficiency (ηSTH) for a cascade pressure reduction system with 5 chambers. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 1 Pa, and a solid heat recuperation effectiveness (εSS) of 0.5. . 93 Figure 2.16: System efficiency (ηSTH) for a cascade pressure reduction system with five chambers. All of these values are calculated for ceria at a reduction temperature (TRED) of 1800 K, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5...... 94 eff Figure 2.17: a) effective pump efficiency (ηPUMP ) and b) solar to hydrogen efficiency (ηSTH) values for cascade pressure reduction systems with various numbers of chambers. All the calculations here are done for ceria at a reduction temperature (TRED) of 1,800 K and a final reduction pressure (pRED) of 1 Pa...... 94 eff Figure 2.18: a) effective pump efficiency (ηPUMP ) and b) solar to hydrogen efficiency (ηSTH) values for cascade pressure reduction systems with various numbers of chambers. All the calculations here are done for ceria at a reduction temperature (TRED) of 1,800 K and a final reduction pressure (pRED) of 100 Pa...... 96 Figure 2.19: System efficiency (ηSTH) for inert gas sweep system. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, and a solid heat recuperation effectiveness (εSS) of 0.5...... 96 Figure 2.20: System efficiency (ηSTH) for inert gas sweep system with different reduction pressures. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5...... 97 Figure 2.21: Inert gas flowrate ratio (nio) for use in an inert gas sweep for reduction at various reduction oxygen partial pressures. Calculations were done for ceria with a reduction temperature of 1,800 K...... 98 Figure 2.22: ηSTH values for ceria at various inert/oxygen separation efficiency values. All calculations were done at a reduction temperature of 1,800 K, a reduction oxygen partial pressure of 0.1 Pa, a gas heat recuperation effectiveness of 0.9, and a solid heat recuperation effectiveness of 0.5...... 99

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Figure 3.1. SEM images of spray-dried particles before calcine: a) NS, b) Al30, c) Zr30, d) Fe30, e) Fe67. White scale bars represent 5 m lengths...... 130 Figure 3.2. SEM images of spray-dried particles after calcine at 1,200°C for 8 hours: a) NS, b) Al30, c) Zr30, d) Fe30, e) Fe67. White scale bars represent 5 m lengths...... 131 Figure 3.3. Particle size distribution of candidate materials after calcining at 1,200°C. The x-axis on the inset figure is zoomed in on range of particles from 0-100 m, and larger figure’s axis is extended over the entire particle range...... 133 Figure 3.4. X-ray diffraction spectra for spray-dried candidate materials after calcining at 1,200°C ...... 134 Figure 3.5. Thermodynamic predictions of reduction and oxidation behavior of candidate materials in air and inert environments...... 136 Figure 3.6. Thermogravimetric analysis for candidate active materials over six redox cycles, with oxidation at 650°C and reduction at 900°C, 1,050°C, and 1,200°C. Plots at the top of each column show the temperature profile, and plots below show the mass change of the sample during redox cycling...... 138 Figure 3.7: Measured viscosity of pH-tuned hercynite slurries at various shear rates. Final pH of slurry is given in legend, and pure water is included for comparison. 95% confidence intervals are shown for multiple measurements. The data for pH 11 is not shown due to measurement issues...... 139 Figure 3.8: Measured stress of pH-tuned hercynite slurries at various shear rates. Final pH of slurry is given in legend, and pure water is included for comparison. 95% confidence intervals are shown for multiple measurements. The data for pH 11 is not shown due to measurement issues...... 140 Figure 3.9: Calculated yield stress of hercynite slurries modified to various pH values. 95% confidence intervals are shown for each calculated value. The data for pH 11 is not shown due to measurement issues...... 140 Figure 3.10: Measured viscosity of pH-tuned boehmite slurries at various shear rates. Final pH of slurry is given in legend, and pure water is included for comparison. 95% confidence intervals are shown for multiple measurements. The data for pH 5 is not shown due to measurement issues...... 141 Figure 3.11: Measured stress of pH-tuned boehmite slurries at various shear rates. Final pH of slurry is given in legend, and pure water is included for comparison. 95% confidence intervals are shown for multiple measurements. The data for pH 5 is not shown due to measurement issues...... 141 Figure 3.12: Calculated yield stress of boehmite slurries modified to various pH values. 95% confidence intervals are shown for each calculated value. The data for pH 5 is not shown due to measurement issues...... 142 Figure 3.13: Zeta potential of combined iron oxide and aluminum oxide slurry. Average measurement is shown with 95% confidence intervals...... 143

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Figure 3.14: Zeta potential of boehmite slurry. Average measurement is shown with 95% confidence intervals...... 144 Figure 3.15: SEM images of spray dried hercynite particles, each with 10wt% excess alumina and a pH modified to a) 3.7, b) 5.4, c) 7.4, and d) 10.1. All images were done at a magnification of 1,000X, and the scale bar represents 10 μm...... 145 Figure 3.16. Mass change for top performing candidate materials over six redox cycles. Theoretical maximum mass change for complete reduction is shown by solid lines...... 148 Figure 4.1: Calculated oxygen vacancy formation energy for pure oxides and aluminates for both cobalt and iron ...... 165 Figure 4.2: Structure of un-doped a) FeAl2O4, b) CoAl2O4, and FeAl2O4 doped with a single Co. All images display oxygen atoms as red, aluminum atoms as light blue, iron atoms as gold, and cobalt atoms as dark blue. In each image, charge density change iso-surfaces are shown as semi-transparent to illustrate charge transfer: the teal iso-surface denotes electron density loss, while the yellow iso-surface shows electron density gain...... 166 Figure 4.3: Net charge transfer for the 30 closet nearest neighbors (NN) of the oxygen vacancy. Each atom is labeled by the location relative to the oxygen vacancy, as well as the type (A or B) of the metal ion site. Bader charge analysis was used to assign charge to specific atoms...... 166 Figure 4.4: Calculated oxygen vacation formation energy and relative structure energy values...... 167 Figure 4.5: Structure of a) FeAl2O4, b) CoAl2O4, and FeAl2O4 doped with a single Co. Each of the above images are for a 50% inverse structure, in which 2 Al atoms are on tetrahedral sites that are 1st nearest neighbor to the O-vacancy. All images display oxygen atoms as red, aluminum atoms as light blue, iron atoms as gold, and cobalt atoms as dark blue. In each image, charge density change iso-surfaces are shown as semi-transparent to illustrate charge transfer: the teal iso-surface denotes electron density loss, while the yellow iso-surface shows electron density gain...... 168 Figure 4.6: Net charge transfer for inverse spinel structures for the 30 closet nearest neighbors (NN) of the oxygen vacancy. Each atom sorted by the location relative to the oxygen vacancy. Bader charge analysis was used to assign charge to specific atoms...... 169 Figure 4.7: Structures of a) FeAl2O4, b) CoAl2O4, and Co-doped FeAl2O4. Each of the above images are for a fully inverse structure, in which 0 Al atoms are 1st nearest neighbor to the O-vacancy. CoAl2O4 is an exception, where 1 Al is nearest neighbor. All images display oxygen atoms as red, aluminum atoms as light blue, iron atoms as gold, and cobalt atoms as dark blue. In each image, charge density change iso-surfaces are shown as semi-transparent to illustrate charge transfer: the teal iso-surface denotes electron density loss, while the yellow iso-surface shows electron density gain...... 169 Figure 4.8: Projected density of states for a) FeAl2O4, b) FeAl2O4 after O-vacancy creation, c) Co-doped FeAl2O4, d) Co-doped FeAl2O4 after O-vacancy creation, and e) Co-doped nd st FeAl2O4 after O-vacancy creation with the doped Co as a 2 -nearest neighbor instead of 1 .

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All simulations are for 2 Al first nearest neighbors. Energy values on the horizontal axis are relative to the Fermi energy for each system. Positive and negative vertical axis values are for spin-up and spin-down states, respectively. Density of states are scaled using different values for ease of viewing, and are not meant to be directly additive to obtain the total DOS...... 171 Figure 4.9: XPS measurements with fitted component curves for a) Fe2p in un-doped hercynite, b) Fe2p for Co-doped hercynite, and c) Co2p for Co-doped hercynite...... 174 Figure 4.10: Hydrogen and oxygen production values per cycle measured in stagnation flow reactor. The average production per unit mass of solid sample is shown for the second to the twelfth cycle, as typically the first cycle is much different. Error bars show the 95% confidence interval...... 175 Figure 4.11: Calculated oxygen vacation formation energy and relative structure energy values, sorted by sum of total energy for each case...... 178 Figure 4.12: Projected density of states for CoAl2O4 with a) 1 and b) 2 Co as a nearest neighbor to the oxygen that will be removed. Energy values on the horizontal axis are relative to the Fermi energy for each system. Positive and negative vertical axis values are for spin-up and spin-down states, respectively. Density of states are scaled using different values for ease of viewing, and are not meant to be directly additive to obtain the total DOS...... 181 Figure 4.13: Hydrogen production rates for un-doped and Co-doped hercynite for an example oxidation reaction. For both materials, the fifth cycle oxidation is shown...... 183

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CHAPTER I

INTRODUCTION

Hydrogen has a number of very important uses, but future production will need to be done renewably and carbon-free to avoid further damage from climate change. This work focuses on the production of hydrogen from solar energy. It is important to understand why a specific route of hydrogen production (out of the many that have been studied) is beneficial and why it is challenging. This chapter will give an overview of solar thermochemical hydrogen production

(including a comparison of alternatives), an overview of general concepts, a description of active materials, as well as discussions about the reactors and operating conditions used.

1.1. Overview of Solar Thermochemical Water Splitting

A general overview of current and future hydrogen production methods will be given to identify the importance of solar thermochemical water splitting as an area of current and future study. First, the importance and uses of hydrogen will be briefly discussed, followed by a look at the various ways hydrogen might be produced more renewably. Finally, general concepts around two-step solar thermochemical water splitting reaction cycles will be introduced and discussed.

1.1.1. Importance of Hydrogen

Hydrogen (H2) is an attractive carbon-free alternative fuel, due to a high specific energy density and the fact that combustion of H2 only produces water, both of which make it an attractive carbon-free alternative fuel. The oxygenation (e.g., burning) of H2 produces only H2O.

Additionally, compressed H2 gas is one of the few materials that can achieve an energy density

>2.5 times the energy density of gasoline by mass [1]. This fuel is also no longer only a futuristic

possibility. Major car companies are producing consumer models of hydrogen fuel cell cars [2], making this technology much more mainstream and wide spread.

Hydrogen is also used in a variety of other ways. In addition to a transportation fuel, hydrogen is widely used in various chemical industries, especially in the processing of fossil fuels

[3] and production of ammonia [4]. The transportation sector in the U.S. relied on fossil fuels for

95% of its energy in 2015 [5], meaning that hydrogen will still be required for fossil fuel processing for a long time into the future. Additionally, even if hydrogen fuel cells were ubiquitous, large amounts of hydrogen will still be required for ammonia production for the use of agricultural fertilizers. 9.3 million metric tons of ammonia were produced in the U.S. in 2014 (145 million tons worldwide) [6], and future increases in the global population will increase this number dramatically.

1.1.2. Hydrogen Generation

Current hydrogen production methods require reaction with other materials, many of which produce CO2. Unlike hydrocarbons, no large quantity of H2 exists on Earth to exploit; therefore,

H2 must be produced from other compounds. Steam methane reforming and other fossil fuel-based processes are currently the main method used to produce H2 [7, 8]. However, because these approaches also generate CO2 as a by-product, the produced H2 has a significant carbon footprint.

The reactions for the production of H2 from steam methane reforming are shown in Equation (1.1), and the subsequent water-gas shift reaction is shown in Equation (1.2). These show that for every

4 moles of H2, 1 mole of CO2 is generated; this corresponds to 5.46 kg of CO2 produced per kg of

H2. It should be noted that this amount of CO2 generated is just based on the chemistry alone and does not include the heat required to drive the reaction.

퐶퐻4 + 퐻2푂 → 퐶푂 + 3 퐻2 (1.1) 퐶푂 + 퐻2푂 → 퐶푂2 + 퐻2 (1.2)

Current renewable H2 generation methods have low process efficiencies, and therefore require further advances to be commercially viable [9, 10]. Generally, renewable H2 production methods can be categorized as biological, electrolytic (driven by renewable electricity), or solar

[11-15]. These categories will be briefly discussed.

Biological methods include fermentation and photosynthesis, both of which tend to require a lot of energy for production and separation of H2. Some algae or bacteria produce hydrogen in anaerobic conditions when excess energy from sunlight is collected; hydrogen production in this process is called direct photolysis. However, the simultaneous production of oxygen requires downstream H2/O2 separation, which is an energetically expensive process on an explosive mixture. Furthermore, the presence of O2 can cause the biologic organism to halt H2 production altogether. Thus, while these technologies are promising due to the fact that the only required inputs are sunlight and water (both abundant), direct photolysis is limited to very low efficiencies

[14]. Fermentation produces hydrogen by breaking down of biomass, and can be done with or without sunlight. However, both of these types of processes tend to be slow, require a lot of energy

(low efficiency), and produce impure H2 (requiring purification) [14].

Electrolytic methods currently exist commercially, but with a low efficiency. Research efforts attempt to increase the current 56-73% efficiency [16] by using a proton exchange membrane or solid oxide cell and by increasing the temperature and pressure at which these cells operate [14]. However, all electrolytic methods suffer from the same two fundamental issues. First, all of these methods first require electricity to be produced, which is an inherently inefficiency process. Thus, the inherent inefficiencies with electricity production are then compounded with whatever inefficiencies exist in these electrolytic processes to result in an overall system efficiency that is much lower. Even if the electricity production and electrolysis can be done efficiently, the

electricity used to drive the electrolysis must be renewable itself if the H2 produced is to be considered renewable. If fossil fuels are used to drive electrolysis, the produced emissions can be even greater than methane reforming [14]. Renewable electricity generation suffers from additional efficiency and reliability challenges, meaning that overall system efficiencies will be even lower for renewable hydrogen produced via electrolysis.

Solar methods include photocatalytic and thermal-driven reactions, which use the energy inherent in solar radiation in different ways to drive chemical reactions. Photocatalytic methods use a semiconductor material in which an electron is excited above the band gap of the material, leaving an electron hole that splits water into O2 and H2 (with the specifics differing based on the material used) [14, 15]. This technology is thus conceptually simple, as the only inputs required are sunlight and water and no complex controls nor extreme conditions are needed. However, these methods tend to produce a H2/O2 mixture, which leads to safety and separation requirements as discussed above. It should be noted that some variations of this technology, such as the use a proton exchange membrane to separate the two reactions, produce H2 and O2 separately, but these processes can suffer from additional issues, such as increased complexity and the need for improved membrane materials [17]. Additionally, since only wavelength of light that are above a certain energy used in a particular material, much of the spectrum of sunlight is either not used or used much less efficiently (due to thermal relaxation of energies greater than the band gap) [14].

Solar thermal methods benefit from using the entire solar spectrum of light by concentrating sunlight and using the resulting energy flux to drive a thermal process. These thermal processes can be biomass gasification or solar thermochemical processes. Biomass gasification is a mature technology, but suffers from low thermal efficiencies due to water in the biomass, and often requires purification steps to obtain pure H2 [11, 14]. In contrast, solar thermochemical

processes can involve any number of process steps, from one-step thermolysis to a complex multi- reaction network [13]. These will be discussed further below.

Solar-based routes are of particular interest because while effective collection of sunlight can be challenging, using such an abundant renewable resource directly leads to potentially highly efficiency processes. Earth-incident sunlight has the capacity to provide ample energy to meet human needs [18]. However, the efficient collection and storage of this diffuse and transient energy resource presents significant technical and economic challenges [18]. While it could be argued that all energy pathways start with solar energy in some form, here the focus is no direct use of solar energy in some form. Of the current methods proposed for solar energy collection and conversion, solar thermal water splitting (STWS) driven by concentrated sunlight is especially promising because it utilizes the entire solar spectrum to split water without the inherent losses in growing biomass or generating current. Consequently, STWS has the potential to achieve high theoretical solar to H2 efficiencies [12, 19]. This is the focus of the remainder of this work.

1.1.3. Two-Step Solar Thermochemical Water Splitting Concepts

Solar thermochemical water splitting uses concentrated sunlight from mirrors that can be configured in a variety of ways, as shown in Figure 1.1. The collected solar energy is focused on a reactor, heating it to high temperatures to drive the endothermic decomposition of H2O into H2 and O2. This direct splitting water in a single reaction step is called direct thermolysis. Although this single-step process is simple, it is impractical because of the high reaction temperatures

(>2,200°C) required for even minimal extents of reaction [20]. Additionally, because thermolysis produces H2 and O2 simultaneously, it requires high-temperature H2/O2 separation steps to prevent product recombination and an explosive mixture.

Water splitting can be divided into two or more steps in which the H2 and O2 are produced separately, thus avoiding the issues of high-temperature gas separations [21], but these reaction cycles typically come at some other cost. Two-step thermochemical water splitting cycles rely on the reduction and subsequent re-oxidation of metal oxides and require reduction temperatures

>1,000°C, but much lower than 2,000°C [12, 13, 20, 22]. Multi-step (>2 steps) cycles can operate at a maximum temperature of <900°C, which can remove difficulties associated with operation at higher temperatures. However, these cycles typically utilize a metal in conjunction with harsh acids or bases and often include an electrolysis step, which necessitates a conversion to electricity

[21, 23-30]. Hazardous chemicals, complicated process designs, and compounding inefficiencies from numerous process steps make multi-step cycles unlikely to achieve the high efficiencies required for economical H2 generation [12, 20].

In two-step STWS, a metal oxide is heated to a high temperature (TRED) by concentrated sunlight to reduce and generate O2, as shown in Equation (1.3). Here, δ is the extent of non- stoichiometry, and in this case can be thought of as the number of oxygen vacancies in a metal oxide per mole of the base oxide. Thus, MOx is the oxidized state of a generic metal oxide in which

δ = 0, and MOx-δ is the reduced state of a metal oxide, which has δ metal oxides per X oxygen molecules in the original oxide. The value of δ depends on the material and operating conditions.

In the second step, the reduced metal oxide is exposed to steam that re-oxidizes the material and forms H2, as shown in Equation (1.4). The two half-reactions form a full reduction and oxidation

(redox) cycle, shown schematically in Figure 1.1c. Oxidation has traditionally been done at a temperature (TOX) that is at least 500°C lower than TRED [20, 32-39]; this is indicated in the schematic in Figure 1.1c. However, oxidation can occur at temperatures up to and including the reduction temperature [40-44]. The choice of these temperatures will be discussed further in

Section 1.3.

훿 푀푂 → 푀푂 + 푂 (1.3) 푥 푥−훿 2 2 푀푂푥−훿 + 훿 퐻2푂 → 푀푂푥 + 훿 퐻2 (1.4) A similar set of reactions analogous to Equations (1.3) and (1.4) describes the splitting of

CO2 to produce CO. Solar thermochemical CO2 splitting has been examined for some time as a way to produce a liquid solar fuel, using solar thermochemical reactions to produce syngas (a mixture of CO and H2), and then using established methods (e.g., Fischer-Tropsch process) to produce a liquid hydrocarbon fuel [45-47]. While there are some obvious differences, many works treat H2O splitting and CO2 splitting as more or less analogous [48]. This work focuses on water splitting specifically due to the advantages of carbon-free hydrogen in a long-term energy economy, but it should be noted that most (if not all) of the trends and effects outlined in this and other works should also apply directly to CO2 splitting as well.

While conceptually simple, implementation of STWS is complicated and many improvements are required if it is to become an efficient and commercially viable process. Three important and interrelated aspects of STWS need to be considered and further developed: 1) the optimal active material that undergoes the redox reactions; 2) the system operating conditions; and

3) the design of the reactor in which the materials will be used, including delivery of solar thermal energy to the active materials, containment of these materials, and control of the reactions. This

thesis includes all three of these aspects of research to greater or lesser extents; they will be discussed further below.

1.2. Two-Step STWS Active Materials

The identification of active and robust materials that efficiently undergo redox reactions to drive STWS at practical conditions and rates is an area of substantial research [31, 49]. The ideal

STWS material has high H2 production capacity, a low reduction temperature, fast kinetics, a long lifetime, compatibility with containment materials, non-toxic composition, and low cost. The first two desirable characteristics are thermodynamic properties of the material, which, together with fast kinetics, are the major attributes that determine the overall efficiency of the system. The other desirable characteristics affect the costs and potential hazards of operating a STWS system based on a particular redox material.

1.2.1. Thermodynamics of STWS Materials

Any chemical reaction is subject to thermodynamic limitations, and the two reactions that make up a redox cycle can combine into a net water splitting reaction. The overall energy input and entropy gain of a two-step STWS cycle must be at least that of direct water splitting, ΔH ≥

286 kJ/mol and ΔS ≥ 44.4 J/mol∙K, and the oxidation and reduction steps must each be spontaneous

(ΔG < 0). Based on these requirements, both the individual reaction steps and the overall STWS cycle are governed a number of thermodynamic relationships (Equations (1.5)-(1.8)), in which the subscripts OX and RED indicate the water splitting (oxidation) reaction and the thermal reduction reaction, respectively.

1 ∆퐺 = ∆퐻 − 푇 (∆푆 + 푆푂2 ) ≤ 0 (1.5) 푅퐸퐷,푇푅퐸퐷 푅퐸퐷 푅퐸퐷 푅퐸퐷 2 푇푅퐸퐷 ∆퐺 = −∆퐻 − ∆퐻퐻2푂 − 푇 (−∆푆 + 푆퐻2 − 푆퐻2푂) ≤ 0 (1.6) 푂푋,푇푂푋 푅퐸퐷 푓,푇푂푋 푂푋 푅퐸퐷 푇푂푋 푇푂푋 ∆퐻 푐푦푐푙푒 = ∆퐻푅퐸퐷 + ∆퐻푂푋 ≥ 286 푘퐽/푚표푙 (1.7) ∆푆푐푦푐푙푒 = ∆푆푅퐸퐷 + ∆푆푂푋 ≥ 44.4 퐽/푚표푙 ∙ 퐾 (1.8)

Equation (1.6) states that the oxidation reaction must be exothermic (ΔHOX < 0) because the overall water splitting step is entropically unfavorable (ΔS < 0) and both steps must be exergonic (ΔG < 0). The entropy decrease of the oxidation reaction arises because the entropy loss in reducing H2O to H2 (ΔS0 ≈ -58 J/mol∙K) is greater than any possible entropy gain from re- oxidizing the STWS materials [33]. The exothermic oxidation reaction (ΔHOX < 0) thus requires the thermal reduction step must be endothermic by at least 286 kJ/mol in order to satisfy the overall energetic requirements of water splitting, as shown in Equation (1.7). Hence, materials with reduction enthalpies <286 kJ/mol will require additional electrical or mechanical work to drive water splitting, that must be included in the overall process evaluation. This explains why some materials (e.g., Co3O4 and CdO) which have the desirable property of reducing at relatively low temperatures do not re-oxidize with water and therefore do not drive STWS [50, 51]. Conversely, materials with a highly exothermic oxidation readily split water but have correspondingly large endothermic reduction enthalpies. Thus, these materials only undergo reduction at very high temperatures, as observed for WO3 or Nb2O5 [50, 51]. This can be explained by a simple examination of the equation of Gibbs free energy: ΔG = ΔH – T·ΔS. Assuming a near-constant ΔS

(which most thermal reduction of metal oxides have), a larger ΔH value will require a larger T value in order to still make the overall ΔG spontaneous (<0). Additionally, when re-oxidation of the material is highly exothermic, not all of the released heat can be recovered (as we describe in

Section 1.3), leading to a less efficient process. Therefore, an ideal material is one where the reduction reaction is sufficiently endothermic to drive the water splitting reaction, but not so endothermic as to require impractical reduction temperatures or squander valuable process heat.

1.2.2. Current STWS Materials

Two-step water splitting cycles can be categorized by their reaction mechanisms: volatile stoichiometric chemistries, non-volatile stoichiometric chemistries, or oxygen vacancy chemistries. The two stoichiometric mechanisms involve the generation of a stoichiometric quantity of oxygen (0.5 moles) and hydrogen (1 mole) for each mole of reacting oxide as it undergoes reduction or oxidization, respectively. Both stoichiometric chemistries are decomposition reactions which produce either gaseous products, as in the volatile stoichiometric chemistries (e.g., ZnO → Zn(g) + ½O2 [52]), or solid products with an altered crystal structure, as in the non-volatile stoichiometric chemistries (e.g., Fe3O4 → 3FeO + ½O2 [35]). O-vacancy chemistries release O2 by the formation of O-vacancies in the metal oxide lattice (e.g., CeO2 →

CeO2-δ + δ/2 O2 [53]), while the overall crystal structure of the solid remains unchanged. Generic chemical reactions for each of these mechanisms are shown in Equations (1.9)-(1.11).

1 푀푂 → 푀 + 푂 (1.9) (푠) (푔) 2 2 1 (1.10) 푀푂 → 푀푂 + 푂 푥 (푠1) (푠2) 2 2 훿 (1.11) 푀푂 → 푀푂 + 푂 푥 (푠1) 푥−훿 (푠1) 2 2 Most of the early STWS cycles were based on stoichiometric chemistries, but more recently the focus has shifted to O-vacancy based mechanisms. The review below briefly highlights research advances for each of these mechanisms, along with general descriptions of the benefits and drawbacks to each.

1.2.2.1. Volatile Stoichiometric Chemistries

STWS volatile stoichiometric chemistry cycles offer high H2 production per unit mass of active material, albeit generally with some operational challenges. Volatile stoichiometric cycles are attractive as they offer high specific H2 production capacities. For example, BeO has the highest theoretical H2 production capacity of any material, with just under 40 mmol of H2/g of

BeO (39,982 μmol/g). All of the material in volatile stoichiometric cycles is active (participates in the redox cycle) and no excess mass stabilizes the reduced material as a condensed phase. This necessitates difficult gas handling steps, especially quenching of the reduction product gases to separate the reduced material from the liberated O2 and preventing the reverse re-oxidation [31].

After the reduced material solidifies, complicated processing steps are required to transport the reduced solid material out of the quenching zone and into the oxidation reactor. Despite these challenges, two volatile stoichiometric cycles are still being studied, the ZnO and the SnO2 cycles

[52, 54].

The ZnO/Zn system (shown in Equations (1.12) and (1.13)) was an early promising redox cycle for splitting water [52]. Thermal reduction of ZnO at atmospheric pressure without an inert gas requires high temperatures (~1,700-2,000°C) [55], but lowering the O2 partial pressure via vacuum pumping or the use of an inert sweep gas lowers the reduction temperatures to ~1,300°C

[56, 57]. However, this lower reduction temperature leads to a slow reaction rate [56, 58-60].

Oxidation with H2O can be done with solid, liquid, and gaseous Zn, although the reaction with either liquid or gaseous Zn is significantly faster than with solid Zn [56, 61-64].

1 푍푛푂 → 푍푛 + 푂 (1.12) (푠) (푔) 2 2 푍푛 + 퐻2푂 → 푍푛푂 + 퐻2 (1.13) While high reduction temperatures and a slow reduction rate are detrimental to efficient

STWS, the most challenging impediment to efficient hydrogen generation is the back-reaction of

Zn(g) to ZnO(s) with the released O2 on cooling of the reduction product stream. This reaction is both thermodynamically favored and has fast kinetics [65], meaning the produced gases must be separated rapidly, typically by quenching. In addition to the significant improvements needed to the quenching step, transport of the reduced and oxidized materials (which often deposit on the

walls of the quenching or oxidization reactors) will have to be considered if the ZnO/Zn process is to become a viable STWS system.

Similar to ZnO, SnO2 dissociates to form SnO and O2 (as seen in Equations (1.14) and

(1.15)), but at a more practical temperature (~1,200°C at low pO2) than ZnO reduction. However, unlike ZnO which forms metallic Zn, SnO2 decomposition produces gaseous SnO. Further reduction to form Sn metal is possible, but undesirable due to higher reduction temperatures, additional reaction steps [66], faster oxidation kinetics, and fewer oxidation competing reactions

[67].

1 푆푛푂 → 푆푛푂 + 푂 (1.14) 2 (푠) (푔) 2 2 푆푛푂 + 퐻2푂 → 푆푛푂2 + 퐻2 (1.15) The SnO2/SnO cycle has many advantages over the ZnO/Zn cycle, but some critical problems remain. SnO condenses at 1,527°C while Zn condenses at 907°C, meaning a slower quench rate can be employed to limit gas phase re-oxidation during SnO/O2 separation [54].

Additionally, the undesirable re-oxidation of SnO(s) by O2 is slower than the re-oxidation of Zn(s), with a Zn re-oxidation reaction which is roughly twice as fast as the SnO re-oxidation reaction

[65]. Furthermore, water almost completely re-oxidizes SnO back to SnO2 [54, 68, 69], significantly higher than the re-oxidation of Zn, albeit at a slower rate [68]. While the SnO2/SnO

STWS cycle appears to be more promising than the ZnO/Zn cycle, the SnO2/SnO cycle is still plagued by low recovery of the reduced material and the difficulties inherent in quenching the reduction product and handling the resulting reduced solids.

1.2.2.2. Non-Volatile Stoichiometric Reactions

Ferrite (Fe3O4) is attractive for STWS because of its relatively high theoretical H2 production capacity (~4,300 μmol/g) and the absence of a volatile reduced metal oxide, but suffers from low melting temperatures. Nakamura developed the first two-step non-volatile metal oxide

redox cycle based on the oxidation and reduction of Fe2+/3+ as it transitions between magnetite

(Fe3O4) and wüstite (FeO) [35]. The reduction temperature of Fe3O4 reduction can be lowered from ~2,200°C under atmospheric conditions to 1,300-1,400°C at reduced O2 partial pressures

[22]. However, Fe3O4 reduction is complicated by the low melting temperature of Fe3O4

(~1,700°C) and FeO (~1,350°C) [22, 35, 70]. This leads to extensive sintering of the active materials which causes material deactivation by decreasing the surface area [71]. This in turn leads to low yields of Fe3O4 from the transport-limited steam oxidation of FeO [72, 73]. These factors complicate the design and operation of STWS reactors based on the Fe3O4/FeO redox material.

Two modifications have been suggested to overcome the deficiencies of the Fe3O4/FeO redox cycle: introduction of an inert carrier or buffer to minimize sintering and doping Fe3O4 to lower the reduction temperature of the oxidized phase and/or to raise the melting temperature of the reduced phase. Supports such as silica particles [74] and yttrium stabilized zirconia (YSZ) have been tried, but these materials either did not do enough to prevent sintering or added too much inert mass. In the second suggested approach, Fe3O4 is doped with many different elements (Zn,

Sn, Mn, Co, Ni, and combinations thereof) in an attempt to lower the reduction temperature and to prevent the sintering and melting of the reduced species [75-80]. Unfortunately, Zn and Sn dopants form separate phases and Zn sublimes and deposits on the reactor walls [31, 75, 78], while.

Mn dopants lower the oxidation reaction rate with no corresponding increase in hydrogen production capacity [76]. Co and Ni dopants show promise, as the reduction products melt at temperatures roughly 125°C higher than pure Fe3O4 [77]. However, techniques to prevent sintering of the reduced product, such as the use of a support, are required. Adding a support lowers the specific hydrogen production capacity by adding to the thermal mass of the system.

1.2.2.3. Oxygen Vacancy Mechanism Reactions

Oxygen vacancy STWS cycles tend to have faster reaction rates and fewer problems associated with phase change, but typically have much lower hydrogen production capacities. In contrast to the stoichiometric STWS chemistries described above, redox materials operating through the O-vacancy mechanism maintain the same phase throughout reduction and oxidation; the O2 released during reduction results from the formation of vacancies within the material.

Therefore, many of these cycles represent incomplete reductions of the active materials in potentially stoichiometric reactions. In other words, the amount of O liberated is insufficient to drive complete decomposition of the material to a distinct reduced phase at the reduction temperatures employed. The materials involved in these redox cycles, in general, can drive water splitting at conditions where they do not suffer the problems of melting or formation of gaseous reduced products. Consequently, handling the reduced materials produced by the O-vacancy mechanism is simpler than handling the materials produced by stoichiometric chemistry mechanisms. However, the H2 production capacities of O-vacancy mechanism materials are generally lower than those of stoichiometric chemistries because only a fraction of the metal cations participates in the redox process, while the others provide the structural stability required to avoid a phase transition.

In O-vacancy mechanism STWS cycles, the relevant measure of the activity is the change in the extent of O non-stoichiometry between the oxidation and reduction steps (Δδ), as shown in

Equations (1.16)-(1.18). The Δδ is determined by the temperature, the oxygen or water partial pressure operating conditions of the individual reaction steps, and the degree to which the thermodynamic equilibrium is achieved. As will be described in Section 1.2.1, a high reduction temperature and low oxygen partial pressure drive reduction while a low oxidation temperature

and high H2O partial pressure drive oxidation and H2 generation. This is illustrated graphically in

Figure 1.2, which shows the O non-stoichiometry for CeO2 at various temperatures and O2 partial pressures. As this plot demonstrates, cycles can operate isothermally, isobarically, or by changing both temperature and oxygen partial pressure. The amount of non-stoichiometry achieved for a given set of operating conditions depends on the material undergoing STWS.

훥훿 푀푂 → 푀푂 + 푂 (1.16) 푥−훿푂푋 푥−훿푅퐸퐷 2 2 (1.17) 푀푂푥−훿푅퐸퐷 + 훥훿 퐻2푂 → 푀푂푥−훿푂푋 + 훥훿 퐻2 ∆훿 = 훿푅퐸퐷 − 훿푂푋 (1.18) 1.2.2.3.1. Ceria-Based Cycles

Cerium(IV) oxide (CeO2 or ceria) is a material that undergoes an oxygen vacancy redox cycle and has a number of benefits, although it must be reduced at high temperatures. The originally proposed cycle was the reduction of CeO2 to Ce2O3, but the required reduction temperature (>2,000°C) is close to the melting temperature of Ce2O3, which led to sintering and substantial evolution of gaseous Ce [81]. Kaneko et al. suggested that ceria could function as an

O-vacancy mechanism STWS material after testing various mixed-metal cerium oxides [53].

Chueh et al. were the first to suggest un-doped ceria as an O-vacancy mechanism STWS material and found that ceria produced ~379 μmol H2/g when reduced at 1,500°C and oxidized at 800°C

[82]. Ceria also benefits from rapid reduction kinetics, which is generally limited by the heating rates of the reactor. Additionally, ceria has been shown to be stable over hundreds of cycles [32,

82] and is capable of undergoing isothermal water splitting [42, 43]. Many works [48, 83, 84] have quantified the thermodynamics of ceria over a wide range of temperatures and pressures, as shown in Figure 1.2. However, the practicality of pure ceria in STWS applications is limited by the high reduction temperatures (>1,500°C) required to produce substantial quantities of H2.

Figure 1.2: Thermodynamic extent of oxygen non-stoichiometry (δ) of ceria, based on temperature and partial pressure of oxygen [38].

Many elements have been doped into ceria in an attempt to increase the H2 production and lower the required reduction temperature. Divalent [85-87] and trivalent [53, 85, 88-94] dopants have not achieved these goals. The di- and trivalent elements decrease the oxygen vacancy formation enthalpy to such an extent that the oxygen vacancies no longer possess the energy required to split water. Some tetravalent (4+ oxidation state) elements have proven to be successful

4+ 4+ at increasing the H2 production of ceria. While some dopants (Sn and Ti ) form phases inert to

STWS [90], Zr4+ and Hf4+ dopants increase the STWS capacity of ceria but lower the oxidation rate [93, 95]. Tetravalent dopant ions that are physically smaller than Ce4+ improve STWS by inducing lattice strain in the material, rather than changing oxidation states [96]. The tensile strain imposed by doping Zr4+ and Hf4+ counteracts the compressive strain of forming an oxygen vacancy, thus lowering the overall energy required to break the O-Ce bonds to form the vacancy.

For Zr4+ and Hf4+, the decrease in O-vacancy formation energy is sufficient to facilitate reduction, but not so large as to hinder water splitting (oxidation).

Additional dopants (Mg, Ca, Ni, Fe, and rare earth elements) have been mixed into Ce1- xZrxO2 in attempts to further increase H2 yields [94, 97-99]. Of these elements, Ni and Fe had no effect on H2 production, while Ca and Mg both increased H2 yields [97]. Co-doping Pr, La, and

Tb produced more H2 than single-doped Ce0.75Zr0.25O2 [99]. The benefit of including a second dopant is potentially attributable to changes in the oxygen vacancy formation energy from a slight modification to the reducibility of the Ce4+ ions [96] or to alterations in the Zr-induced strain field by the variously sized lanthanides. Additionally, La and Gd were found to increase cycling stability of the doped ceria [94].

Of the other ceria dopants examined, W was found to not facilitate STWS [91]. However,

U was found to increase STWS, either due to its ability to enable over-oxidation, i.e. Ce1-xUxO2+y

[100], or its ability to take on multiple oxidation states which then aid Ce4+ reduction [96, 101].

While U is unlikely to be used in commercial processes due to the radioactivity of its various isotopes, it may serve as a useful tool for identifying other dopants.

Despite the extensive effort to dope ceria to achieve better STWS materials, the relatively low H2 production capacity at reasonable reduction temperatures (≤1,500°C) still limits its potential application in commercial STWS processes. Either very efficient heat recuperation is needed so that ceria can be repeatedly cycled efficiently, or new dopants from the small list of elements not already studied must be identified to either lower the reduction temperature, increase

H2 production, or preferably both.

1.2.2.3.2. The Hercynite Cycle

The hercynite cycle is now considered to operate via an oxygen vacancy mechanism [102], but was originally thought to be a stoichiometric reaction. In 2010, Scheffe et al. discovered that when cobalt ferrite was deposited on alumina it would produce H2 after reduction at 1,200°C,

which is 200-300°C lower than when cobalt ferrite was deposited on a zirconia support, as shown in 1a and b [103]. This reaction was hypothesized to operate through a stoichiometric mechanism, where the CoFe2O4 reacts with the Al2O3 support during reduction to produce a solid solution of

FeAl2O4 (hercynite) and CoAl2O4: CoFe2O4 + 3Al2O3  CoAl2O4 + 2FeAl2O4 + ½ O2. This solid solution was expected to then revert back to two separate phases upon oxidation by steam [103].

In addition to producing H2 at relatively low reduction temperatures [103, 104], on-sun experiments have demonstrated that the doped-hercynite material produces ~10 times more H2 produced and does not sinter, in contrast to Fe3O4 materials [71]. However, at reduction temperatures above 1,500°C, CoFe2O4 produces more H2 per gram of active material than doped- hercynite [103] because a larger fraction of the material undergoes active reduction and oxidation.

Figure 1.3: a) The doped-hercynite cycle H2 production rates after reduction at various temperatures and oxidation at 1,000°C. b) the production rates of CoFe2O4 on ZrO2 after reduction and oxidation under the same conditions as a) [103]. c) shows the H2 production rates of doped-hercynite operating under temperature swing water splitting conditions (left two peaks) and isothermal water splitting conditions (right peak) [40]. Doped-hercynite active materials not only reduce at relatively low temperatures, but they are also capable of producing substantial quantities of H2 isothermally; operating isothermally at

1,350°C, the doped-hercynite material generated >12 times more hydrogen than ceria after reduction at 1,350°C and oxidation at 1,000°C [40]. The larger extent of re-oxidation at the higher water splitting oxidation temperatures is attributed to the kinetic limitations of the re-oxidation of doped-hercynite at lower temperatures. Similar to other STWS materials, the rate-limiting step of

the doped-hercynite oxidation half-cycle is attributed to surface reaction [105]. For doped- hercynite active materials, slow reaction rates are the most significant limitation to the productivity and overcoming the slow kinetics of the surface reactions of O2 generation and H2O spitting during the reduction and oxidation half-cycles, respectively, is the major challenge to improving this material.

1.2.2.3.3. Perovskite Cycles

Perovskites are a broad set of materials having the formula ABO3 which is highly amenable to doping, composed of many different elements, and stable at relatively high levels of oxygen non-stoichiometry. These characteristics led Nalbandian et al. and Evdou et al. to suggest perovskites as oxygen exchange materials in a high temperature methane membrane reactor and in two-step chemical looping cycles [106, 107]. This prompted Scheffe et al. to suggest using

LaxSr1-xMnO3 (SLM) in STWS [108], which was followed soon thereafter by McDaniel et al. who used a slightly modified LaxSr1-xMnyAl1-yO3 (SLMA) perovskite [109]. Both the SLM and SLMA materials produce more H2 than ceria at low temperatures; where SLMA produced 2.3 times the amount of H2 after reduction at 1,350°C (307 μmol/g) as ceria produced after reduction at 1,500°C, as shown in Figure 1.4 [108, 110].

Doping the perovskite structure with other elements led to a variety of effects and trade- offs. The SLM materials appear to suffer from sintering issues, while the SLMA material produced

3+ consistent quantities of H2 over 80 cycles [108, 109]. This suggests that Al acts to stabilize the

SLMA perovskite and prevent sintering. While doping the B-site stabilized the SLMA materials, doping the A-site with 10-40% Sr increased the overall productivity of the SLM materials from

40.6 to 397.7 µmol of H2/g [111]. However, similar to other STWS materials, the increased H2 production capacity was coupled with a decreased reaction rate [111]. The ability of Sr to decrease

the reduction enthalpy stems from reducing or eliminating the band gap of the perovskites which reduces the energy increase of the two O2- electron as they transfer to the conduction band metal states when O leaves the lattice. This was demonstrated by Deml et al. who used PBE+U density functional theory (DFT) to predict the oxygen vacancy formation energies of several perovskite materials including SLM and LaMnO3 [112].

Figure 1.4: Perovskite based solar thermal water splitting cycles. SLMA materials are capable of producing significantly more H2 than ceria when reduced at 1,350°C and oxidized at 1,000°C [110]. b) Many other perovskite formulations have shown to be capable of undergoing STWS. From left to right are the H2 production capacities of Ba25Sr75Co80Fe20, Ba50Sr50Co80Fe20, La60Sr40Co20Fe80, LaSrCo, La65Sr35Mn, La50Sr50Mn, and La50Sr50Mn under operating conditions shown in the inset [113]. La- and Fe-based perovskites have also been tested for solar thermal gas splitting capabilities, namely LaxA1-xFeyB1-y (A= Sr, Ce and B=Co, Mn). Although these materials evolved significant quantities of O2 during reduction, almost no gas splitting behavior was observed. This is because the low reduction enthalpies, as calculated by Deml et al. [112], are insufficient to reduce water, which is consistent with the thermodynamic argument made in Section 1.2.1. While these perovskites were incapable of STWS, CO2 splitting does occur when these materials are supported on SiO2 [114]. This suggests that a reaction occurs between the SiO2 support and the perovskite, producing a material with a reduction enthalpy sufficiently high to reduce CO2.

Non-La based perovskites can also potentially undergo STWS; McDaniel et al. showed that CaTi1-xFexO3 has similar O exchange behavior as CeO2 [110], and Demont et al. showed that

BaySr1-yCoxFe1-x splits water, although with a lower H2 production capacity than LaySr1-yMnO3, as shown in Figure 1.4b [113]. This suggests that the SLM materials are just a starting point, and new perovskites with higher H2 production capacities are possible either through further doping of existing materials or from as yet unidentified combinations of elements. The only major drawback to using perovskites identified so far is their relatively high heat capacity [108]. This means that perovskite cycles should preferably operate isothermally or near-isothermally to reduce the heating and cooling losses associated with large temperature swing cycles.

1.3. Modes of Operation

There are several broad choices to be made for solar thermochemical operation, and the optimal set of choices is not immediately clear. While some combinations of choices (“modes”) may have higher theoretical efficiencies, the implementation of others may result in higher practical efficiencies. These are important choices, and are very closely linked with other aspects of material and reactor design. Two of these parameters are the oxygen partial pressure and operating temperature.

1.3.1. Reaction Temperature

One of the most important considerations for STWS operation is the temperature at which the reduction and oxidation reactions occur. Higher reduction temperatures often lead to higher overall efficiencies, even when re-radiation losses are considered [35, 37, 38, 41, 42, 115, 116].

This occurs because many materials reach higher extents of reduction at higher temperatures, which leads to an increased hydrogen production capacity of the material. The higher extent of reduction also leads to an increased re-oxidation driving force, thus reducing the amount of excess steam required for oxidation. Overall this lowers the material flow rates required to produce equivalent quantities of hydrogen [37, 38, 41, 42, 44, 115, 116]. While higher reduction

temperatures lead to higher theoretical efficiencies, reduction cannot be conducted at arbitrarily high temperatures due to practical and economic concerns such as the temperature limits of the reactor containment materials, the melting point of the active materials, and the diminishing thermal and economic efficiency of solar collection as temperature increases. Given these constraints, the highest possible operating temperature of the H2 production plant should be used for the reduction step in order to achieve high solar to H2 efficiencies, although higher temperatures can introduce operational problems.

In addition to the temperature of the reduction step, the temperature of the oxidation step significantly influences the efficiency of the system. STWS operation methods can be categorized by the temperature difference between the reduction and oxidation steps (ΔT). The three modes of operation are: temperature swing water splitting (TSWS) where ΔT > 300°C, isothermal water splitting (ITWS) where ΔT ≈ 0°C, or near-isothermal water splitting (NITWS) where 0 < ΔT <

~300°C. Traditionally, only TSWS was studied because it was believed that large temperature swings of >400°C were necessary for both the reduction and oxidation reactions to be thermodynamically favorable [34-39]. However, it was recently demonstrated that the oxidation reaction occurs at temperatures up to and including the reduction temperature, proving that isothermal operation is possible [40-44]. In this mode of operation, changes in the chemical potential of the reactant and product gasses are controlled by altering the partial pressure of gasses rather than varying the temperature of the system [40]. The maximum driving force at a given temperature can be achieved by minimizing the O2 partial pressure during reduction, and maximizing the H2O/H2 partial pressure ratio during oxidation.

Isothermal conditions offer increased materials longevity, better efficiency in reactor heating, and faster oxidation kinetics than temperature swing systems. Under ITWS operation, the

active materials and reactor containment materials do not undergo temperature cycles, which limits the thermal stresses and resulting thermal fatigue experienced during TSWS. Additionally, whereas each TSWS cycle involves heating the active material, and potentially the reactor material, from the oxidation temperature to the reduction temperature using solar energy, this is not required in ITWS. Also, because ITWS is conducted at higher oxidation temperatures, the rate of the oxidation reaction is increased, which is especially important for kinetically limited materials.

While isothermal operation presents a number of thermal benefits, the chemical potential cycling requirement presents challenges. ITWS is driven by the chemical potentials of reactant and product gases [40], and so lower oxygen and higher water partial pressures are required to drive the respective reduction and oxidation reactions than in TSWS. Low O2 partial pressures are achieved by either large vacuum pumping or large excess of inert gas sweep. This increases the mechanical work required for reduction because operating under high vacuum is difficult at the high temperatures required for reduction [117], and large inert flow rates require significant pumping work to move excess gas and significant separations work to remove the O2 from the inert. The high levels of excess steam required to drive the oxidation reaction lead to higher pump work requirements and higher heat duties for steam heating, the effects of which can have substantial detrimental effects on the overall efficiency of the system [115].

Operating under NITWS conditions is a trade-off between the benefits and detriments of

ITWS and TSWS, is potentially the best mode of operation, and has been shown to offer the highest theoretical efficiencies [115]. In NITWS, both temperature and pressure drive the water splitting reaction forward. By slightly decreasing the oxidation temperature from ITWS, the required steam partial pressure is lowered, minimizing the gas heating duty. Furthermore, the energy required to

reheat the oxidized material to the reduction temperature is lowered over TSWS by maintaining a relatively high oxidation temperature. The optimal NITWS ΔT is found where the steam heating and solids reheating duties are balanced. The optimal NITWS oxidation temperature depends on active material and the efficiencies of solid and gas heat recuperation [115].

1.3.2. Oxygen Removal

Low oxygen partial pressures (pO2) increase the thermodynamic driving force of the reduction reaction. Vacuum pumping and inert gas sweeping are two methods for achieving the required low pO2. Generally, vacuum pumping is assumed to require only one energy input, the pump work (Qpump), as shown in Figure 1.5a. The use of a sweep gas is more complex, as the inert gas is either continually generated, or re-conditioned by removing the oxygen (typically via a membrane). In the continually generated inert gas case, there are three energy inputs: producing the inert gas (Qinert), pumping the inert gas to the reduction reaction pressure (Qpump), and heating the inert gas to the reduction reaction temperature (Qheat), as shown in Figure 1.5b. In the recycled inert gas case, after the gas is initially generated, there are only two energy inputs: the O2 separation energy (Qsep), and the reheating energy (Qheat), as shown in Figure 1.5c. In the case of inert regeneration, the temperature at which the inert sweep gas is separated from the generated O2 (Tsep) affects the system efficiency; increasing Tsep increases the energy requirements for separation

(Qsep) but decreases the heating energy (Qheat) more. This indicates that separation should occur at the highest possible temperatures. High temperature oxygen separation membrane technologies currently exist and require temperatures around 900°C to operate [118].

Figure 1.5: Three different methods to achieve low pO2 of oxygen and the associated energy requirements: a) vacuum pumping, b) direct inert gas sweep, and c) recycled inert gas sweep. When comparing vacuum pumping and inert gas sweeping, Ermanoski et al. showed that at pO2 = 1 Pa the overall system efficiency is almost 40% for ceria with vacuum pumping and is

0% for continuously produced inert gas sweeping [119]. The large efficiency difference arises because of the large quantity of inert gas required to achieve low pO2 and the related energy inputs, i.e. Qinert, Qpump and Qheat [119]. However, Bader et al. showed that the inert gas regeneration method significantly decreases the quantity of fresh inert required to achieve the necessary pO2 and thus increases system efficiency to about 31.6% for ceria cycles [42] which is closer to the theoretical efficiency achievable using vacuum pumping. Although the theoretical efficiency with vacuum pumping is 40%, the effect of mass and heat transport limitations inside the reactor could reduce the overall efficiencies if reduction driven by a vacuum requires significantly longer reduction times and, therefore, incurs additional re-radiation losses. Understanding the effects of mass and heat transport on the O2 removal rate is essential for predicting the performance of the solar water splitting process, and requires further investigation.

1.4. Reactor Design

An active material with optimal thermodynamics, kinetics, and durability will be unable to meet the economic, efficiency, or production capacity milestones necessary for commercialization

unless it is utilized in an efficient reactor. In order to generate H2 efficiently, a STWS reactor must deliver solar heat and steam to the reactive materials without dissipating unnecessary energy or requiring significant amounts of external work, while also being resistant to structural failure.

Thermal management is essential for a process requiring high temperatures and heat loads. To minimize unnecessary heat loss, the reactor design should limit the number of solar reflections, limit heat loss by re-radiation or convection from the light absorbing material, rapidly transfer heat to the active material; and avoid unnecessary temperature changes of the reactive material and reactor system because heat recuperation is inefficient between solids [19]. An optimal design also requires efficient gas mass transport to maintain thermal efficiency and minimize electrical work.

Furthermore, reactants (H2O) and product gases (H2 and O2), must be delivered and removed rapidly to prevent back reaction. This necessitates convective transport, high material void space, or porosity. Moving the reactant and product gases efficiently requires a low pressure drop across the material. Finally, an efficient reactor must effectively transport solids and maintain the structure of active solid materials.

In addition to managing the parasitic losses from heat and mass transport, several other key design considerations must be met in order to achieve efficient H2 production. The reactor should be scalable, eliminating the need for many small reactors which do not provide for an economy of scale and suffer from higher heat losses because of their small volume to surface area ratios. The reactor should also ideally decouple the reduction and oxidation reaction times, because the oxidation and reduction kinetics are unlikely to be identical. Additionally, the reactor should spatially or temporally decouple the reduction and oxidation steps to separate the O2 and H2 product gases. Minimizing moving parts is very useful, as these are liable to fail in the high temperature environment of a STWS reactor. Next, reactors should avoid the use of windows as

these are prone to particle deposition and subsequent degradation. Finally, the reactors should be constructed of materials which are compatible with the active material and stable under the operating conditions. There has been significant pioneering work done in designing STWS reactors based on these design goals, and these reactors can broadly be classified as monolithic or particle systems based on the configuration of the active material.

1.4.1. Monolithic Reactors

In monolithic-type solar thermal reactors, the active material is self-supporting and the reduction and oxidation steps are spatially separated either by mechanical motion (e.g. rotation) of the material or re-direction of the solar beam. Stationary monolith cavity (SMC) reactors are the simplest of the monolithic reactors, an example of which is shown in Figure 1.6a. SMC reactors have one [32] or more [120] reaction chambers where either reduction or oxidation occurs. The reactive material is either free standing or supported on a scaffolding [121]. During reduction, the active material is directly illuminated by concentrated sunlight via a quartz window or indirectly heated using a containment structure. After the reduction step ends, the irradiation of the reduced material is decreased either by redirection onto different reaction chambers [122] or by directing the solar beam away from the receiver altogether [32]. Steam is then injected into the chamber containing the reduced material, generating H2 and re-oxidizing the STWS material. Although simple and therefore less prone to mechanical failure, solar cavity receiver designs have no inherent way to recuperate the heat released during the temperature change between the reduction and oxidation steps, unless the reduction and oxidation reactors are contained within a single cavity. Additionally, if a quartz window is used to introduce solar radiation into the reaction chamber, the potential size of the reactor concept is limited, meaning that these concepts cannot fully exploit the economies of scale.

Figure 1.6: Monolith-based solar thermal water splitting reactor concepts: a) the porous monolith cavity reactor,[32] b) the rotating piston reactor,[65] c) the CR5,[123] and d) the SurroundSun reactor.[124] The rotating piston reactor is an alternative version of a cavity reactor that uses the same general concepts as described above, but is designed for use with volatile stoichiometric reaction chemistries [58]. In this design, pellets of the active material are fed into the hot zone of the reactor by a piston as shown in Figure 1.6b.; as the material reduces and volatilizes, it is swept by an inert gas into a quenching chamber. As the irradiated material is consumed, the piston pushes fresh active material into the reactor [125]. This reactor has similar advantages and disadvantages as

SMC reactors, but enables the use of volatile chemistries. Thus, the advantages and disadvantages of the volatile chemistries are thus added to the advantages and disadvantages of an SMC reactor.

Another reactor design is the Counter-Rotating-Ring Receiver Reactor Recuperator (CR5)

[123, 126], shown in Figure 1.6c. The CR5 design attempted to overcome the solid-solid heat recuperation challenges of cavity-based reactors using counter-rotating reactive rings. The reactor is composed of a set of rings constructed of the reactive material or a support coated by reactive material where each ring rotated in the opposite direction to that of its neighbors. At any given time, a quarter of each ring is irradiated through a quartz window and undergoes reduction, the opposite quarter is exposed to steam and undergoes oxidation, while the two remaining quarter sections allow the reactive material to exchange and recuperate heat. Inert sweep gas prevents the mixing of the H2 and O2. While this reactor can theoretically recuperate heat, a major challenge is mechanical failure of the rings during operation due to thermal stress between the illuminated and dark sides of the ring that creates rapid temperature changes and gradients during redox cycling.

The SurroundSun reactor design avoided the use of a quartz window by using a “tube within a tube” design, as shown in Figure 1.6d [124, 127-129]. In this configuration, one or more reaction tubes are packed with reactive material and housed within an insulated cavity.

Concentrated sunlight enters through an open aperture. When operated in a temperature swing mode, at any given time half of the tubes are illuminated and undergo reduction while the other half are exposed to steam and undergo oxidation. In contrast, when operated in an isothermal mode, all of the tubes are continually illuminated and oxidation and reduction are controlled by the gasses in individual tubes. An inert sweep gas flows through the tubes undergoing reduction to remove generated O2, while steam is fed to the tubes undergoing oxidation. The lack of a closed window made of a transparent material or moving mechanical parts means that this concept is potentially scalable to large production sizes, provided that suitable containment materials are available.

However, this reactor design has several serious drawbacks. Regardless of the mode of operation,

there is uneven radial illumination of the reactor tubes [130] and poor thermal transport within the bed which results in lower redox reaction rates and, thus, lower production throughput [124]. The need to pump H2O and H2 and O2 through the packed bed may limit the diameter of each reaction tube; this problem could be alleviated by using many small tubes, but this potentially exacerbates the problem of uneven illumination of the reactor tubes. Additionally, hot and cold spots within the beds which stem from uneven solar flux on the reaction tubes and poor heat conductivity result in an underutilization of the material [124], which also limits the overall efficiency.

1.4.2. Particle Reactors

A second category of STWS reactors is particle reactors, which utilize moving particles to facilitate product transport away from the solids. These reactors can decouple the reduction and oxidation times, which is useful as the reduction and oxidation reaction kinetics are generally not identical. Often, these reactors utilize efficient direct irradiation of the reactive material. For instance, in rotating particle reactors, sunlight enters a rotating cylinder containing particles of the reactive materials along its main axis, as shown in Figure 1.7a [131, 132]. Because these reactors are typically used for volatile stoichiometric cycles, the reduced material is removed via a vacuum pump and transported to the quenching and oxidation units. Fresh particles are typically fed into the reactor by screw feeders [133]. While these concepts allow for good mass and heat transport properties, they have several limitations. Because these reactors use direct radiation, they suffer from the size limitations associated with the use of quartz windows. Additionally, the reaction drum along with the associated insulation and instrumentation must rotate, which poses challenges at temperatures >1,500°C. To avoid the issues of a mechanically moving reactor, several designs suggest feeding the reactive particles at the top of a slope and gravitationally transporting the particles through the incident beam, as shown in Figure 1.7b [134-136].

Figure 1.7: Particle based solar thermal water splitting reactor concepts: a) rotating cavity particle reactor,[135] b) non-rotating particle flow reactor,[131] c) aerosol flow reactor,[137] d) internally circulating fluidized bed reactor,[138] e) moving particle packed bed reactor. [139] The labels in a) and b) point out: a) 1) rotating drum, 2) actuation, 3) aperture, 4) cavity, 5) screw feeder, 6) product outlet port, 7) rotary joint, 8) working fluids, 9) insulation, 10) quartz window; b) 1) water-cooled window mount and vortex-flow generation. 2) water-cooled cavity aperture, 3) BOP and data-acquisition cavity access ports, 4) alumina-tile reaction surface, 5) annular solid ZnO exit, 6) bulk insulation and cavity-shape support and 7) central product-vapor and gas exit. Particle reactors have also been developed for non-volatile cycles. In aerosol-based designs, the particles are loaded at the top of a long tubular reactor and are gravity fed through the hot zone of the reduction chamber, as shown in Figure 1.7c [137]. An inert gas that flows counter to the reducing particles can be used to increase residence time and/or mass transfer. If the material is based on a volatile stoichiometric cycle, the gases are collected from the top of the reactor for quenching [140]. If the material is based on non-volatile chemistry, the particles are collected at the bottom and subsequently fed to the oxidation reactor. The particles are generally heated indirectly, where the reactor tube absorbs sunlight and then transfers heat to the active particles

via conduction through the reactor walls and radiation from the walls to the particles. While these reactors benefit from low mass transfer limitations and the lack of a quartz window, they lack a direct connection to an oxidation zone, solid-state heat recuperation, and a mechanism to move the re-oxidized material back to the top of the aerosol reactor.

Internally circulating fluidized bed reactors attempt to retain the low mass transfer limitations of the aerosol reactor while simultaneously enabling both STWS steps to occur in a single chamber [138]. In this reactor, particles are loaded into a reaction chamber containing a center draft tube, as shown in Figure 1.7d. Inert gas is fed from the bottom of the reactor below the draft tube, fluidizing the particles and forcing them to rise through the center and fall through the annulus. The particles are directly irradiated at the top of the particle bed through a quartz window. The circulating bed facilitates heat transfer from the irradiated top of the bed to the lower sections. To control whether the reduction or oxidation step is occurring, an inert gas or steam is flowed through the fluidized bed of active particles. This reactor design still requires a quartz window, and separate reduction and oxidation reactive chambers must be used if valuable concentrated sunlight is to be utilized consistently.

In the Moving Packed Particle Bed Reactor, Ermanoski et al. attempt to decouple the reduction and oxidation times while maximizing solid-solid heat recuperation [139]. The reactor consists of a directly illuminated reduction chamber at the top of a hollow ceramic screw located above an oxidation chamber, as shown in Figure 1.7e. Starting at the bottom of the reactor, fully oxidized particles are lifted out of the oxidation chamber via a screw elevator to the bottom of the reduction chamber. Then a rotating casing pushes the particles up a stationary ceramic screw which also serves as a heat exchanger. At the top of the reactor, the particles are irradiated with concentrated sunlight, causing them to reduce before they drop through the hollow center of the

screw elevator. The generated O2 is removed by vacuum pumping. As the particles move through the center of the stationary screw, they can exchange heat with the oxidized particles moving up the outer section of the reactor. The moving bed in the screw acts as a pressure buffer between the low pressure reduction zone and the roughly atmospheric pressure oxidation zone. Once in the oxidation chamber, the particles form a secondary moving bed through which H2O is pumped, re- oxidizing the particles and generating H2. This reactor decouples reduction and oxidation times, simplifies solid-solid heat transfer, houses both redox steps, and continuously processes material.

A recently updated design has been suggested which deemphasizes solid/solid heat recuperation and incorporates staged pressure reduction to facilitate O2 removal [117]. However, it utilizes a quartz window for direct irradiation, which in addition to limiting the reactor size based on the size of available quartz windows will likely attract fines through thermophoretic deposition. This could lead to diminishing window transparency or catastrophic window failure if the deposited particles produce hot spots on the window. Additionally, large parts of the reactor will have to rotate at the high reduction temperatures, straining the vacuum seals and stressing the materials of construction.

1.4.3. The Solar-thermal Particle Flow Reactor

Using concepts developed and lessons learned from past reactor designs [138, 139] and general design principles outlined above, the Solar-thermal Particle Flow Reactor (SPFR) is proposed based on well-established hydrocarbon thermal cracking unit designs currently used in the chemical processing industry [141], and is shown in Figure 1.8. The SPFR is a particle reactor without any mechanically moving parts. The design is based on a beam-up approach and consists of several sets of reduction/oxidation chambers arranged in an inner and outer circle where the reduction chambers form the inner ring of the reactor. Concentrated sunlight is directed up through the gap in the bottom of the receiver, as illustrated in Figure 1.8. The downward-facing cavity

receiver only requires one reflection and minimizes convective losses from hot gas rising out of the aperture. The entire apparatus is envisioned to sit atop a solar power tower.

Figure 1.8: The Solar-thermal Particle Flow Reactor. a) an individual reduction/oxidation reactor unit and b) receiver configuration containing multiple individual reduction/oxidation reactor units. Reactors are not shown to scale. In this reactor, the reactive particles flow between the reduction and oxidation reaction chambers. Starting from fully oxidized material, the particles fall through the Falling Particle

Reduction Reactor, which is heated indirectly through the reactor wall, where a low oxygen partial pressure is maintained using vacuum pumping. Baffles can be added to lengthen the particle residence time and enable staged pressure reduction [117]. The reduced particles are stored as a pseudo-packed bed at the bottom of the reduction chamber before they enter the oxidation reactor.

The packed bed moving storage section provides a pressure buffer, enabling for both a low pressure in the reduction reactor and a high steam partial pressure in the oxidation reactor [139]. After storage, the reduced particles are entrained in steam and conveyed upwards to a fluidized bed oxidation chamber, enabling the oxidation time to be decoupled from the height of the steam conveyance tube. Oxidized particles fall to the bottom of the oxidation chamber, and are stored in

a second packed bed before re-entering the reduction reactor. The reduction and oxidation reactions could be run at near-isothermal temperatures, reducing the need for solid-solid heat recuperation and lowering thermal stresses on the reactor due to cooling and reheating. Process heat recuperation is accomplished by utilizing the heat from both the liberated O2 and product H2 and unreacted H2O to generate electricity or pre-heat reactant steam. This heat recuperation allows excess steam to be used for fluidization without large efficiency losses, while concomitantly providing a large thermodynamic driving force for the oxidation reaction. Gas-solid mixing in a fluidized bed promotes rapid transport of gaseous reactants and products to and from the reactive solids, while providing heat transfer between the gases and solids, including both the reactive solid particles and the reactor walls (which are heated by the solar flux). This overcomes heat and mass transfer issues of packed beds; with better heat transfer and all of the solids passing through the hot zone, “hot” and “cold” spots are minimized. Key challenges that remain include the development of high temperature ceramic heat exchangers for steam/steam heat exchange, high- temperature and thermally shock resistant reactor containment materials that are compatible with active materials, robust flowable active particles, and tower/receiver/heliostat designs which allow for efficient beam-up solar heating.

1.5. Project Objectives

This dissertation addresses some of the critical challenges in the complex and interconnected field of solar thermochemical water splitting. It starts by taking an overall view of the system and then estimates the effect of various operational choices and values in order to illustrate when different aspects of the system as a whole become more or less important. Next, engineered particles are improved upon to increase the overall system efficiency. Finally, the all- important active material is considered, and a promising material is more fully explained. Specific

objectives for each of these three main areas are given below, and the order of these objectives will be reflected in the dissertation to follow.

1.5.1. System Efficiency Objectives

This work aims to determine the relative importance of different aspects of the solar thermochemical water splitting process. A numerical model will be developed in order to calculate and compare energy requirements of different aspects of the process. Specifically, the effect of oxidation kinetics will illustrate a more nuanced choice of operating conditions. The full possible range of gas and solid heat recuperation will be shown to be less important to an idealistic system but remain critical to realistic systems. The effect of separation temperatures of both reduction and oxidation products will be quantified for the first time and found to be less important for oxidation water but highly impactful on reduction for inert gas. The differences between open-loop and recycled inert gas reduction systems will be directly compared vis-à-vis the separation temperature.

Finally, a cascade pressure system and a high temperature recycled inert gas system will both be considered as ways to improve the reduction reaction energy requirements. The high temperature separation recycled inert gas system will be identified as a very promising possibility, depending on inert gas flowrates and separation efficiency values.

1.5.2. Active Particle Objectives

Spray dried particles are produced with a mixture of ceramic starting particles. Specifically, partial flocculation driven by pH modification is induced in a known stable colloid suspension in order to produce solid spherical particles of a mixed metal oxide. Particle morphology is examined by microscopy. Zeta potential and yield stress measurements are used to show partial flocculation, while at the same time the suspension does not exhibit expected sedimentation behavior. This new

formulation and effects will be helpful for the future scalable production of solar thermochemical water splitting particles.

Three secondary metal oxides (Al2O3, ZrO2, and Fe2O3) are combined with manganese oxide in order to examine the impact on the redox behavior and sintering temperatures. Spray dried articles are tested in a thermogravimetric analyzer (TGA) over multiple redox cycles with multiple reduction temperatures to evaluate the impact of the different secondary metal oxides on chemical activity and robustness of the particles. The particles are characterized after spray-drying, and after calcining, with respect to their shape, crystal structure, specific surface area, and particle size distribution, using scanning electron microscopy (SEM), X-ray diffraction (XRD), Brunauer–

Emmett–Teller (BET) surface area, and laser diffraction. This characterization helps to explain various performance trends and trade-offs for future development of these particles.

1.5.3. Metal Ion Activity Objectives

This work uses computational simulations to further explore differences between hercynite

(FeAl2O4), cobalt aluminate (CoAl2O4), and hercynite doped with cobalt (CoxFe1-xAl2O4). Charge density transfer upon formation of oxygen vacancies is examined as a way to determine what role each metal ion plays in the material. Computational trends then compared to analytical results in which these materials are characterized with X-ray photoelectron spectroscopy (XPS) while being heated to high reduction temperatures in vacuum. This analysis allows for the determination of oxidation state changes in materials as they reduce, allowing for electron density charge transfer to be experimental observed directly in order to verify computational predictions.

1.6. References

1. U.S. Energy Information Administration (EIA). Few transportation fuels surpass the energy densities of gasoline and diesel. Today in Energy 2013; Available from: http://www.eia.gov/todayinenergy/detail.cfm?id=9991.

2. Yoshida, T. and K. Kojima, Toyota MIRAI Fuel Cell Vehicle and Progress Toward a Future Hydrogen Society. The Electrochemical Society Interface, 2015. 24(2): p. 45-49.

3. U.S. Energy Information Administration (EIA). Hydrogen for refineries is increasingly provided by industrial suppliers. Today in Energy 2016; Available from: http://www.eia.gov/todayinenergy/detail.cfm?id=24612.

4. Winter, C.-J., Hydrogen energy — Abundant, efficient, clean: A debate over the energy- system-of-change. International Journal of Hydrogen Energy, 2009. 34(14, Supplement 1): p. S1-S52.

5. U.S. Energy Information Administration (EIA), Monthly Energy Review. April 2016: Report No.

6. U.S. Geological Survey (USGS), Mineral Commodity Summaries 2016. 2016.

7. Simpson, A.P. and A.E. Lutz, Exergy analysis of hydrogen production via steam methane reforming. International Journal of Hydrogen Energy, 2007. 32(18): p. 4811-4820.

8. The Impact of Increased Use of Hydrogen on Petroleum Consumption and Carbon Dioxide Emissions. 2008, U.S. Energy Information Administration (EIA): Report No. SR/OIAF- CNEAF/2008-04.

9. Barreto, L., A. Makihira, and K. Riahi, The hydrogen economy in the 21st century: a sustainable development scenario. International Journal of Hydrogen Energy, 2003. 28(3): p. 267-284.

10. Ogden, J.M., Prospects for Building a Hydrogen Energy Infrastructure. Annual Review of Energy and the Environment, 1999. 24(1): p. 227-279.

11. Marbán, G. and T. Valdés-Solís, Towards the hydrogen economy? International Journal of Hydrogen Energy, 2007. 32(12): p. 1625-1637.

12. Steinfeld, A., Solar thermochemical production of hydrogen––a review. Solar Energy, 2005. 78(5): p. 603-615.

13. Perkins, C. and A.W. Weimer, Solar-thermal production of renewable hydrogen. AIChE Journal, 2009. 55(2): p. 286-293.

14. Holladay, J.D., et al., An overview of hydrogen production technologies. Catalysis Today, 2009. 139(4): p. 244-260.

15. Zou, Z., et al., Direct splitting of water under visible light irradiation with an oxide semiconductor photocatalyst. Nature, 2001. 414(6864): p. 625-627.

16. Turner, J., et al., Renewable hydrogen production. International Journal of Energy Research, 2008. 32(5): p. 379-407.

17. Wang, Z., et al., Comparison of thermochemical, electrolytic, photoelectrolytic and photochemical solar-to-hydrogen production technologies. International Journal of Hydrogen Energy, 2012. 37(21): p. 16287-16301.

18. Lewis, N.S. and D.G. Nocera, Powering the planet: Chemical challenges in solar energy utilization. Proceedings of the National Academy of Sciences, 2006. 103(43): p. 15729- 15735.

19. Siegel, N.P., et al., Factors Affecting the Efficiency of Solar Driven Metal Oxide Thermochemical Cycles. Industrial & Engineering Chemistry Research, 2013. 52(9): p. 3276- 3286.

20. Perkins, C. and A.W. Weimer, Likely near-term solar-thermal water splitting technologies. International Journal of Hydrogen Energy, 2004. 29(15): p. 1587-1599.

21. Abanades, S., et al., Screening of water-splitting thermochemical cycles potentially attractive for hydrogen production by concentrated solar energy. Energy, 2006. 31(14): p. 2805-2822.

22. Scheffe, J.R. and A. Steinfeld, Oxygen exchange materials for solar thermochemical splitting of H2O and CO2: a review. Materials Today, (0).

23. Naterer, G.F., et al., Thermochemical hydrogen production with a copper-chlorine cycle. I: oxygen release from copper oxychloride decomposition. International Journal of Hydrogen Energy, 2008. 33(20): p. 5439-5450.

24. Naterer, G., et al., Recent Canadian advances in nuclear-based hydrogen production and the thermochemical Cu-Cl cycle. International Journal of Hydrogen Energy, 2009. 34(7): p. 2901-2917.

25. Kasahara, S., et al., Flowsheet study of the thermochemical water-splitting iodine-sulfur process for effective hydrogen production. International Journal of Hydrogen Energy, 2007. 32(4): p. 489-496.

26. Yildiz, B. and M.S. Kazimi, Efficiency of hydrogen production systems using alternative nuclear energy technologies. International Journal of Hydrogen Energy, 2006. 31(1): p. 77- 92.

27. Nakamura, N., et al., Hydrogen production via thermochemical water-splitting by lithium redox reaction. Journal of Alloys and Compounds, 2013. 580, Supplement 1: p. S410-S413.

28. Marugán, J., et al., Study of the first step of the Mn2O3/MnO thermochemical cycle for solar hydrogen production. International Journal of Hydrogen Energy, 2012. 37(8): p. 7017-7025.

29. Kreider, P.B., et al., Manganese oxide based thermochemical hydrogen production cycle. International Journal of Hydrogen Energy, 2011. 36(12): p. 7028-7037.

30. Francis, T.M., P.R. Lichty, and A.W. Weimer, Manganese oxide dissociation kinetics for the Mn2O3 thermochemical water-splitting cycle. Part 1: Experimental. Chemical Engineering Science, 2010. 65(12): p. 3709-3717.

31. Roeb, M., et al., Materials-Related Aspects of Thermochemical Water and Carbon Dioxide Splitting: A Review. Materials, 2012. 5(11): p. 2015-2054.

32. Chueh, W.C., et al., High-Flux Solar-Driven Thermochemical Dissociation of CO2 and H2O Using Nonstoichiometric Ceria. Science, 2010. 330(6012): p. 1797-1801.

33. Meredig, B. and C. Wolverton, First-principles thermodynamic framework for the evaluation of thermochemical H2O or CO2 splitting materials. Physical Review B, 2009. 80(24): p. 245119.

34. Miller, J.E., A.H. McDaniel, and M.D. Allendorf, Considerations in the Design of Materials for Solar-Driven Fuel Production Using Metal-Oxide Thermochemical Cycles. Advanced Energy Materials, 2014. 4(2): p. n/a-n/a.

35. Nakamura, T., Hydrogen production from water utilizing solar heat at high temperatures. Solar Energy, 1977. 19(5): p. 467-475.

36. Diver, R.B., et al., Solar Thermochemical Water-Splitting Ferrite-Cycle Heat Engines. Journal of Solar Energy Engineering, 2008. 130(4): p. 041001-041001-8.

37. Chueh, W.C. and S.M. Haile, A thermochemical study of ceria: exploiting an old material for new modes of energy conversion and CO2 mitigation. Philosophical transactions. Series A, Mathematical, physical, and engineering sciences, 2010. 368(1923): p. 3269-94.

38. Lapp, J., J.H. Davidson, and W. Lipiński, Efficiency of two-step solar thermochemical nonstoichiometric redox cycles with heat recovery. Energy, 2012. 37(1): p. 591-600.

39. Scheffe, J.R. and A. Steinfeld, Thermodynamic Analysis of Cerium-Based Oxides for Solar Thermochemical Fuel Production. Energy & Fuels, 2012. 26(3): p. 1928-1936.

40. Muhich, C.L., et al., Efficient Generation of H2 by Splitting Water with an Isothermal Redox Cycle. Science, 2013. 341(6145): p. 540-542.

41. Hao, Y., C.-K. Yang, and S.M. Haile, High-temperature isothermal chemical cycling for solar-driven fuel production. Physical Chemistry Chemical Physics, 2013. 15(40): p. 17084- 17092.

42. Bader, R., et al., Thermodynamic Analysis of Isothermal Redox Cycling of Ceria for Solar Fuel Production. Energy & Fuels, 2013. 27(9): p. 5533-5544.

43. Venstrom, L.J., et al., Efficient Splitting of CO2 in an Isothermal Redox Cycle Based on Ceria. Energy & Fuels, 2014. 28(4): p. 2732-2742.

44. Krenzke, P.T. and J.H. Davidson, Thermodynamic Analysis of Syngas Production via the Solar Thermochemical Cerium Oxide Redox Cycle with Methane-Driven Reduction. Energy & Fuels, 2014. 28(6): p. 4088-4095.

45. Miller, J.E., A.H. McDaniel, and M.D. Allendorf, Considerations in the Design of Materials for Solar-Driven Fuel Production Using Metal-Oxide Thermochemical Cycles. Advanced Energy Materials, 2014. 4(2): p. 1300469.

46. Kim, J., et al., Comparative analysis of environmental impact of S2P (Sunshine to Petrol) system for transportation fuel production. Applied Energy, 2013. 111(0): p. 1089-1098.

47. Kim, J., et al., Fuel production from CO2 using solar-thermal energy: system level analysis. Energy & Environmental Science, 2012. 5(9): p. 8417-8429.

48. Bulfin, B., et al., Thermodynamics of CeO2 Thermochemical Fuel Production. Energy & Fuels, 2015. 29(2): p. 1001-1009.

49. Muhich, C.L., et al., A review and perspective of efficient hydrogen generation via solar thermal water splitting. Wiley Interdisciplinary Reviews: Energy and Environment, 2016. 5(3): p. 261-287.

50. Lundberg, M., Model calculations on some feasible two-step water splitting processes. International Journal of Hydrogen Energy, 1993. 18(5): p. 369-376.

51. Kodama, T. and N. Gokon, Thermochernical cycles for high-temperature solar hydrogen production. Chemical Reviews, 2007. 107(10): p. 4048-4077.

52. Steinfeld, A., Solar hydrogen production via a two-step water-splitting thermochemical cycle based on Zn/ZnO redox reactions. International Journal of Hydrogen Energy, 2002. 27(6): p. 611-619.

53. Kaneko, H., et al., Reactive ceramics of CeO2–MOx (M=Mn, Fe, Ni, Cu) for H2 generation by two-step water splitting using concentrated solar thermal energy. Energy, 2007. 32(5): p. 656-663.

54. Abanades, S., et al., Novel two-step SnO2/SnO water-splitting cycle for solar thermochemical production of hydrogen. International Journal of Hydrogen Energy, 2008. 33(21): p. 6021- 6030.

55. Loutzenhiser, P.G., A. Meier, and A. Steinfeld, Review of the two-step H2O/CO2-splitting solar thermochemical cycle based on Zn/ZnO redox reactions. Materials, 2010. 3(11): p. 4922-4938.

56. Weidenkaff, A., et al., Thermogravimetric analysis of the ZnO/Zn water splitting cycle. Thermochimica Acta, 2000. 359(1): p. 69-75.

57. Levêque, G. and S. Abanades, Kinetic analysis of high-temperature solid–gas reactions by an inverse method applied to ZnO and SnO2 solar thermal dissociation. Chemical Engineering Journal, 2013. 217(0): p. 139-149.

58. Charvin, P., et al., Experimental study of SnO2/SnO/Sn thermochemical systems for solar production of hydrogen. AIChE Journal, 2008. 54(10): p. 2759-2767.

59. Schunk, L.O. and A. Steinfeld, Kinetics of the thermal dissociation of ZnO exposed to concentrated solar irradiation using a solar-driven thermogravimeter in the 1800–2100 K range. AIChE Journal, 2009. 55(6): p. 1497-1504.

60. Perkins, C., P. Lichty, and A.W. Weimer, Determination of aerosol kinetics of thermal ZnO dissociation by thermogravimetry. Chemical Engineering Science, 2007. 62(21): p. 5952- 5962.

61. Berman, A. and M. Epstein, The kinetics of hydrogen production in the oxidation of liquid zinc with water vapor. International Journal of Hydrogen Energy, 2000. 25(10): p. 957-967.

62. Yang, W., et al., Quantum Chemical Calculations on the Reaction of Zinc and Water in Gas Phase. Combustion Science and Technology, 2013. 186(1): p. 24-33.

63. Melchior, T., N. Piatkowski, and A. Steinfeld, H2 production by steam-quenching of Zn vapor in a hot-wall aerosol flow reactor. Chemical Engineering Science, 2009. 64(5): p. 1095-1101.

64. Abu Hamed, T., J.H. Davidson, and M. Stolzenburg, Hydrolysis of Evaporated Zn in a Hot Wall Flow Reactor. Journal of Solar Energy Engineering, 2008. 130(4): p. 041010-041010.

65. Chambon, M., S. Abanades, and G. Flamant, Solar thermal reduction of ZnO and SnO2: Characterization of the recombination reaction with O2. Chemical Engineering Science, 2010. 65(11): p. 3671-3680.

66. Levêque, G., et al., Characterization of Two-Step Tin-Based Redox System for Thermochemical Fuel Production from Solar-Driven CO2 and H2O Splitting Cycle. Industrial & Engineering Chemistry Research, 2014. 53(14): p. 5668-5677.

67. Vishnevetsky, I. and M. Epstein, Tin as a Possible Candidate for Solar Thermochemical Redox Process for Hydrogen Production. Journal of Solar Energy Engineering, 2009. 131(2): p. 021007-021007.

68. Chambon, M., S. Abanades, and G. Flamant, Kinetic investigation of hydrogen generation from hydrolysis of SnO and Zn solar nanopowders. International Journal of Hydrogen Energy, 2009. 34(13): p. 5326-5336.

69. Abanades, S., CO2 and H2O reduction by solar thermochemical looping using SnO2/SnO redox reactions: Thermogravimetric analysis. International Journal of Hydrogen Energy, 2012. 37(10): p. 8223-8231.

70. Steinfeld, A., S. Sanders, and R. Palumbo, Design Aspects of Solar Thermochemical Engineering—A Case Study: Two-Step Water-Splitting Cycle Using the Fe3O4/FeO Redox System. Solar Energy, 1999. 65(1): p. 43-53.

71. Lichty, P., et al., Atomic layer deposited thin film metal oxides for fuel production in a solar cavity reactor. International Journal of Hydrogen Energy, 2012. 37(22): p. 16888-16894.

72. Coker, E.N., et al., Oxygen transport and isotopic exchange in iron oxide/YSZ thermochemically-active materials via splitting of C(18O)2 at high temperature studied by thermogravimetric analysis and secondary ion mass spectrometry. Journal of Materials Chemistry, 2012. 22(14): p. 6726-6732.

73. Charvin, P., et al., Two-step water splitting thermochemical cycle based on iron oxide redox pair for solar hydrogen production. Energy, 2007. 32(7): p. 1124-1133.

74. Allen, K.M., et al., Study of a Magnetically Stabilized Porous Structure for Thermochemical Water Splitting via TGA, High-Temperature-XRD, and SEM Analyses. Industrial & Engineering Chemistry Research, 2013. 52(10): p. 3683-3692.

75. Kaneko, H., et al., Decomposition of Zn-ferrite for O2 generation by concentrated solar radiation. Solar Energy, 2004. 76(1–3): p. 317-322.

76. Ehrensberger, K., et al., Comparative experimental investigations of the water-splitting reaction with iron oxide Fe1−yO and iron manganese oxides (Fe1−xMnx)1−yO. Solid State Ionics, 1995. 78(1–2): p. 151-160.

77. Allendorf, M.D., et al., Two-Step Water Splitting Using Mixed-Metal Ferrites: Thermodynamic Analysis and Characterization of Synthesized Materials. Energy & Fuels, 2008. 22(6): p. 4115-4124.

78. Bhosale, R., et al., H2 generation from two-step thermochemical water-splitting reaction using sol-gel derived SnxFeyOz. Journal of Renewable and Sustainable Energy, 2011. 3(6): p. -.

79. Kodama, T., et al., Thermochemical hydrogen production by a redox system of ZrO2- supported Co(II)-ferrite. Solar Energy, 2005. 78(5): p. 623-631.

80. Tamaura, Y., et al., Production of solar hydrogen by a novel, 2-step, water-splitting thermochemical cycle. Energy, 1995. 20(4): p. 325-330.

81. Abanades, S. and G. Flamant, Thermochemical hydrogen production from a two-step solar- driven water-splitting cycle based on cerium oxides. Solar Energy, 2006. 80(12): p. 1611- 1623.

82. Chueh, W.C. and S.M. Haile, A thermochemical study of ceria: exploiting an old material for new modes of energy conversion and CO2 mitigation. Philosophical Transactions of the Royal Society A: Mathematical, Physical and Engineering Sciences, 2010. 368(1923): p. 3269-3294.

83. Panlener, R.J., R.N. Blumenthal, and J.E. Garnier, A Thermodynamic Study of Nonstoichiometric Cerium Dioxide. Journal of Physics and Chemistry of Solids, 1975. 36: p. 1213-1222.

84. Bulfin, B., et al., Analytical Model of CeO2 Oxidation and Reduction. The Journal of Physical Chemistry C, 2013. 117(46): p. 24129-24137.

85. Abanades, S., et al., Investigation of reactive cerium-based oxides for H2 production by thermochemical two-step water-splitting. Journal of Materials Science, 2010. 45(15): p. 4163-4173.

86. Meng, Q.-L., et al., Reactivity of CeO2-based ceramics for solar hydrogen production via a two-step water-splitting cycle with concentrated solar energy. International Journal of Hydrogen Energy, 2011. 36(21): p. 13435-13441.

87. Meng, Q.-L., et al., Solar hydrogen production using Ce1−xLixO2−δ solid solutions via a thermochemical, two-step water-splitting cycle. Journal of Solid State Chemistry, 2012. 194(0): p. 343-351.

88. Kaneko, H., et al., Cerium ion redox system in CeO2–xFe2O3 solid solution at high temperatures (1,273–1,673 K) in the two-step water-splitting reaction for solar H2 generation. Journal of Materials Science, 2008. 43(9): p. 3153-3161.

89. Chueh, W.C. and S.M. Haile, Ceria as a Thermochemical Reaction Medium for Selectively Generating Syngas or Methane from H2O and CO2. Chemsuschem, 2009. 2(8): p. 735-739.

90. Jiang, Q., et al., Thermochemical CO2 splitting reaction with CexM1−xO2−δ (M=Ti4+, Sn4+, Hf4+, Zr4+, La3+, Y3+ and Sm3+) solid solutions. Solar Energy, 2014. 99(0): p. 55- 66.

91. Ramos-Fernandez, E.V., N.R. Shiju, and G. Rothenberg, Understanding the solar-driven reduction of CO2 on doped ceria. RSC Advances, 2014. 4(32): p. 16456-16463.

92. Le Gal, A., et al., Reactivity of Doped Ceria-Based Mixed Oxides for Solar Thermochemical Hydrogen Generation via Two-Step Water-Splitting Cycles. Energy & Fuels, 2013. 27(10): p. 6068-6078.

93. Scheffe, J.R., et al., Synthesis, Characterization, and Thermochemical Redox Performance of 4+ 4+ 3+ Hf , Zr , and Sc Doped Ceria for Splitting CO2. The Journal of Physical Chemistry C, 2013. 117(46): p. 24104-24114.

94. Le Gal, A. and S. Abanades, Dopant Incorporation in Ceria for Enhanced Water-Splitting Activity during Solar Thermochemical Hydrogen Generation. The Journal of Physical Chemistry C, 2012. 116(25): p. 13516-13523.

95. Le Gal, A., S. Abanades, and G. Flamant, CO2 and H2O Splitting for Thermochemical Production of Solar Fuels Using Nonstoichiometric Ceria and Ceria/Zirconia Solid Solutions. Energy & Fuels, 2011. 25(10): p. 4836-4845.

96. Scaranto, J. and H. Idriss, The Effect of Uranium Cations on the Redox Properties of CeO2 Within the Context of Hydrogen Production from Water. Topics in Catalysis, 2014: p. 1-6.

97. Kang, M., et al., CO2 splitting via two step thermochemical reactions over doped ceria/zirconia solid solutions. RSC Advances, 2013. 3(41): p. 18878-18885.

98. Kang, M., et al., Enhanced thermochemical CO2 splitting over Mg- and Ca-doped ceria/zirconia solid solutions. RSC Advances, 2014. 4(11): p. 5583-5590.

99. Dasari, H.P., et al., Hydrogen production from water-splitting reaction based on RE-doped ceria–zirconia solid-solutions. International Journal of Hydrogen Energy, 2013. 38(14): p. 6097-6103.

100. Al-Shankiti, I., et al., Solar Thermal Hydrogen Production from Water over Modified CeO2 Materials. Topics in Catalysis, 2013. 56(12): p. 1129-1138.

101. Al-Salik, Y., I. Al-Shankiti, and H. Idriss, Core level spectroscopy of oxidized and reduced CexU1−xO2 materials. Journal of Electron Spectroscopy and Related Phenomena, 2014. 194: p. 66-73.

102. Muhich, C.L., et al., Predicting the solar thermochemical water splitting ability and reaction mechanism of metal oxides: a case study of the hercynite family of water splitting cycles. Energy & Environmental Science, 2015. 8(12): p. 3687-3699.

103. Scheffe, J.R., J.H. Li, and A.W. Weimer, A spinel ferrite/hercynite water-splitting redox cycle. International Journal of Hydrogen Energy, 2010. 35(8): p. 3333-3340.

104. Arifin, D., et al., CoFe2O4 on a porous Al2O3 nanostructure for solar thermochemical CO2 splitting. Energy & Environmental Science, 2012. 5(11): p. 9438-9443.

105. Muhich, C.L., et al., Extracting Kinetic Information from Complex Gas–Solid Reaction Data. Industrial & Engineering Chemistry Research, 2014.

106. Nalbandian, L., A. Evdou, and V. Zaspalis, La1−xSrxMO3 (M = Mn, Fe) perovskites as materials for thermochemical hydrogen production in conventional and membrane reactors. International Journal of Hydrogen Energy, 2009. 34(17): p. 7162-7172.

107. Evdou, A., V. Zaspalis, and L. Nalbandian, La1−xSrxFeO3−δ perovskites as redox materials for application in a membrane reactor for simultaneous production of pure hydrogen and synthesis gas. Fuel, 2010. 89(6): p. 1265-1273.

108. Scheffe, J.R., D. Weibel, and A. Steinfeld, Lanthanum–Strontium–Manganese Perovskites as Redox Materials for Solar Thermochemical Splitting of H2O and CO2. Energy & Fuels, 2013. 27(8): p. 4250-4257.

109. McDaniel, A.H., et al., Sr- and Mn-doped LaAlO3-delta for solar thermochemical H-2 and CO production. Energy & Environmental Science, 2013. 6(8): p. 2424-2428.

110. McDaniel, A.H., et al., Nonstoichiometric Perovskite Oxides for Solar Thermochemical H2 and CO Production. Energy Procedia, 2014. 49(0): p. 2009-2018.

111. Yang, C.-K., et al., Thermodynamic and kinetic assessments of strontium-doped lanthanum manganite perovskites for two-step thermochemical water splitting. Journal of Materials Chemistry A, 2014. 2(33): p. 13612-13623.

112. Deml, A.M., et al., Oxide enthalpy of formation and band gap energy as accurate descriptors of oxygen vacancy formation energetics. Energy & Environmental Science, 2014. 7(6): p. 1996-2004.

113. Demont, A., S. Abanades, and E. Beche, Investigation of Perovskite Structures as Oxygen-Exchange Redox Materials for Hydrogen Production from Thermochemical Two- Step Water-Splitting Cycles. The Journal of Physical Chemistry C, 2014. 118(24): p. 12682- 12692.

114. Jiang, Q., et al., Thermochemical CO2 splitting reaction with supported LaxA1−xFeyB1−yO3 (A=Sr, Ce, B=Co, Mn; 0

115. Ermanoski, I., J.E. Miller, and M.D. Allendorf, Efficiency maximization in solar- thermochemical fuel production: challenging the concept of isothermal water splitting. Physical Chemistry Chemical Physics, 2014. 16(18): p. 8418-8427.

116. Lange, M., et al., T–S diagram efficiency analysis of two-step thermochemical cycles for solar water splitting under various process conditions. Energy, 2014. 67: p. 298-308.

117. Ermanoski, I., Cascading pressure thermal reduction for efficient solar fuel production. International Journal of Hydrogen Energy, 2014. 39(25): p. 13114-13117.

118. Dyer, P.N., et al., Ion transport membrane technology for oxygen separation and syngas production. Solid State Ionics, 2000. 134(1): p. 21-33.

119. Ermanoski, I., N.P. Siegel, and E.B. Stechel, A new reactor concept for efficient solar- thermochemical fuel production. Journal of Solar Energy Engineering, 2013. 135(3): p. 031002.

120. Houaijia, A., et al., Analysis and improvement of a high-efficiency solar cavity reactor design for a two-step thermochemical cycle for solar hydrogen production from water. Solar Energy, 2013. 97: p. 26-38.

121. Agrafiotis, C., et al., Solar water splitting for hydrogen production with monolithic reactors. Solar Energy, 2005. 79(4): p. 409-421.

122. Roeb, M., et al., Operational strategy of a two-step thermochemical process for solar hydrogen production. International Journal of Hydrogen Energy, 2009. 34(10): p. 4537-4545.

123. Diver, R.B., et al., Solar Thermochemical Water-Splitting Ferrite-Cycle Heat Engines. Journal of Solar Energy Engineering, 2008. 130(4): p. 041001.

124. Martinek, J., R. Viger, and A.W. Weimer, Transient simulation of a tubular packed bed solar receiver for hydrogen generation via metal oxide thermochemical cycles. Solar Energy, 2014. 105(0): p. 613-631.

125. Chambon, M., S. Abanades, and G. Flamant, Thermal dissociation of compressed ZnO and SnO2 powders in a moving-front solar thermochemical reactor. AIChE Journal, 2011. 57(8): p. 2264-2273.

126. Miller, J.E., et al., Sunshine to Petrol: Solar thermochemistry for liquid fuels. Abstracts of Papers of the American Chemical Society, 2011. 241.

127. Lichty, P., et al., Rapid High Temperature Solar Thermal Biomass Gasification in a Prototype Cavity Reactor. Journal of Solar Energy Engineering, 2010. 132(1): p. 011012- 011012.

128. Martinek, J. and A.W. Weimer, Design considerations for a multiple tube solar reactor. Solar Energy, 2013. 90: p. 68-83.

129. Melchior, T., et al., A cavity-receiver containing a tubular absorber for high-temperature thermochemical processing using concentrated solar energy. International Journal of Thermal Sciences, 2008. 47(11): p. 1496-1503.

130. Martinek, J. and A.W. Weimer, Evaluation of finite volume solutions for radiative heat transfer in a closed cavity solar receiver for high temperature solar thermal processes. International Journal of Heat and Mass Transfer, 2013. 58(1–2): p. 585-596.

131. Müller, R., P. Haeberling, and R.D. Palumbo, Further advances toward the development of a direct heating solar thermal chemical reactor for the thermal dissociation of ZnO(s). Solar Energy, 2006. 80(5): p. 500-511.

132. Schunk, L.O., et al., A Receiver-Reactor for the Solar Thermal Dissociation of Zinc Oxide. Journal of Solar Energy Engineering, 2008. 130(2): p. 021009-021009.

133. Abanades, S., P. Charvin, and G. Flamant, Design and simulation of a solar chemical reactor for the thermal reduction of metal oxides: Case study of zinc oxide dissociation. Chemical Engineering Science, 2007. 62(22): p. 6323-6333.

134. Mö ller, S. and R. Palumbo, The Development of a Solar Chemical Reactor for the Direct Thermal Dissociation of Zinc Oxide. Journal of Solar Energy Engineering, 2000. 123(2): p. 83-90.

135. Koepf, E., et al., A novel beam-down, gravity-fed, solar thermochemical receiver/reactor for direct solid particle decomposition: Design, modeling, and experimentation. International Journal of Hydrogen Energy, 2012. 37(22): p. 16871-16887.

136. Koepf, E.E., et al., Experimental Investigation of Vortex Flow in a Two-Chamber Solar Thermochemical Reactor. Journal of Fluids Engineering, 2013. 135(11): p. 111103-111103.

137. Scheffe, J.R., M. Welte, and A. Steinfeld, Thermal reduction of ceria within an aerosol reactor for H2O and CO2 splitting. Industrial & Engineering Chemistry Research, 2014.

138. Gokon, N., et al., New Solar Water-Splitting Reactor With Ferrite Particles in an Internally Circulating Fluidized Bed. Journal of Solar Energy Engineering, 2009. 131(1): p. 011007-011007.

139. Ermanoski, I., N.P. Siegel, and E.B. Stechel, A New Reactor Concept for Efficient Solar- Thermochemical Fuel Production. Journal of Solar Energy Engineering, 2013. 135(3): p. 031002-031002-10.

140. Perkins, C., P.R. Lichty, and A.W. Weimer, Thermal ZnO dissociation in a rapid aerosol reactor as part of a solar hydrogen production cycle. International Journal of Hydrogen Energy, 2008. 33(2): p. 499-510.

141. Kunii, D. and O. Levenspiel, Fluidization Engineering. Vol. 2. 1991: Butterworth- Heinemann Boston.

CHAPTER II

EFFICIENCY MODELING

2.1. Abstract

A comprehensive solar-to-hydrogen (STH) efficiency model, which includes the effect of oxidation kinetics, is developed for two-step solar thermochemical redox water splitting processing in which active materials flow through separate reduction and oxidation reactors. Two active redox materials are considered and compared in order to assess the impact of the rate of redox and the hydrogen productivity per cycle on STH efficiency. Reported oxidation rates for reduced cerium oxide (fast kinetics/lower H2 productivity/cycle) and a ferrite/zirconia composite

(slow kinetics/higher H2 productivity/cycle) are used in the model in order to make a realistic comparison. Generally, the efficiency at thermodynamic equilibrium is higher for the ferrite/zirconia composite than ceria. Interactions between material specific parameters are compared, such as the combination of heat capacity and flow rate on sensible heating loads.

Additionally, the sensitivity of oxidation kinetics on the overall cycle efficiency is illustrated.

Model results show that kinetics can have a drastic effect on STH efficiency. Near-isothermal redox processing is more optimal for materials with slower kinetics, especially with moderate to high gas heat recuperation. The kinetic effects are negligible for those active materials having fast oxidation rates, i.e. ceria, which benefit from a larger temperature difference (thermodynamic driving force) between the reduction and oxidation steps. These kinetic effects lead to different optimal operating conditions when oxidation kinetics are included in the analysis as compared to prior models when only thermodynamic equilibrium is considered.

Additional calculations were performed using cerium(IV) oxide and an inert sweep gas was considered as the O2 removal method. Gas and solid heat recuperation effectiveness values were varied between 0 and 100% in order to determine the limits of the effect of these parameters.

The temperature at which the inert gas is separated from oxygen for an open-loop and recycled system was varied. The hydrogen and water separation temperature was also varied and the effect on STH efficiency quantified.

Gas heat recuperation is critical for high efficiency cycles, especially at conditions that lead to high steam and inert gas flowrates. A key area for future study is identified to be the development of ceramic heat exchangers for high temperature gas-gas heat exchange. Solid heat recuperation is more important at lower oxidation temperatures that favor temperature-swing redox processing, and the relative impact of this heat recuperation is muted if the heat can be used elsewhere in the system. A high separation temperature for the recycled inert gas has been shown to be beneficial, especially for cases of lower gas heat recuperation and increased inert gas flowrates. A higher water/hydrogen separation temperature is beneficial for most gas heat recuperation effectiveness values, though the overall impact on optimal system efficiency is relatively small for the values considered.

Three methods for achieving low oxygen partial pressures for reduction are compared, and the effect of vacuum pump efficiency and inert gas/oxygen separation efficiency are quantified.

Currently available vacuum pump technologies have very low thermodynamic efficiencies at low pressures and are unlikely to provide efficient hydrogen production relative to other oxygen partial pressure lowering technologies. Using currently available pumps arranged in a cascade pressure reduction configuration increases the effective pump efficiency significantly, but by less than an order of magnitude and, therefore, still results in low STH efficiencies for the system. If vacuum

pumps could operate at a low pressure with an efficiency of ~10% or better, vacuum pumping

(including cascade pressure reduction) has the potential to operate very efficiently for solar thermochemical hydrogen production. A novel recycled inert gas sweep with high temperature separation is suggested and STH efficiency values vary significantly depending on the inert gas flowrate required, and will be reactor and reaction rate dependent. However, the use of an inert gas is likely able to take advantage of greater extents of reduction at very low oxygen partial pressures and produce high STH values if the inert gas/oxygen separation is ~10% efficient.

2.2. Introduction

Since the overall goal of thermochemical hydrogen production is to efficiently produce hydrogen from solar energy, the whole process must be analyzed to determine the effects of various parameters on the overall efficiency. Many researchers have attempted to calculate a theoretical

STH efficiency, taking into account various heat sources, sinks, recycles, and different redox materials, and some have used entropy analysis [1, 2]. Even though the two-step thermochemical cycle is conceptually simple, there are many factors in such a process than can be varied. Important factors that must be considered are the choice of the design conditions (temperature and pressure) for both the reduction and oxidation reactions, the method of oxygen removal from the reduction reaction, the heating of steam for the oxidation reaction, separation of the hydrogen product, and heat recuperation within the process [3].

Solar thermochemical water splitting efficiency calculations often use ceria as the active material [4-8]. This is useful as ceria is currently one of the most promising materials available for solar thermochemical water splitting. However, the trends found for ceria may not necessarily hold for other materials. Some groups have attempted to incorporate other materials into efficiency models, including some [9-11] that have used iron oxide in efficiency models. Importantly, Diver

et al. considered iron oxide, but assumed complete conversion for both the reduction and oxidation reaction as well as using only the stoichiometrically required amount of steam [9], which are likely not realistic in an actual system. Singh et al. also considered iron oxide, but only considered energy requirements for reduction rather than the process as a whole [11].

In addition to a constrained materials assessment, many efficiency models have focused on a cycle operating at thermodynamic equilibrium. This is seen as an upper bound on efficiency, as slower reaction rates only limit the achievable efficiency. Additionally, the achievable reaction rates are material and reactor design-dependent, making it difficult to use these analyses to explore general trends. Furler et al. have published the actual efficiency achieved by their experimental reactors [12]. Publication of experimental efficiency values is useful and necessary for the field as a whole, but not as useful for exploring a wide design space and investigating the impact of reaction rates of the system efficiency. Venstrom et al. have used an efficiency model using experimental data and extrapolated what efficiencies could be gained at a larger scale system [13]. Other groups have also discussed the impact of kinetics lowering the system efficiency, both chemical as well as heat and mass transport limitations [14-18].

Heat recuperation is important for any process, but especially so in these types of processes due to the high temperatures (and thus high-grade heat) involved, typically >1,273 K. However, different modes of operation can lead to different types of heat recuperation being more or less important. Muhich et al. [19] and Bader et al. [5] highlighted isothermal operation (where both the reduction and oxidation temperatures are the same) as potentially beneficial for efficient operation.

Ermanoski et al. assessed the concept of isothermal water splitting and found that it can be problematic due to the large amount of excess steam required to re-oxidize the material [20]. They assert that the heat required to heat this excess steam cannot be recovered at temperatures greater

than 1,000°C due to material limitations in potential heat exchangers, thus limiting the overall gas- gas recuperation effectiveness at higher oxidation temperatures [20]. However, this temperature is not a hard limit, as ceramic heat recuperation techniques have been suggested for various reactor concepts that could aid with gas heat recovery at high temperatures [21].

One way to avoid the necessity of very high gas heat recuperation due to high water/steam flowrates is to raise the temperature at which the H2 product is separated from the excess water/steam. When the separation temperature difference between the oxidation temperature and the separation temperature is smaller, the high temperature heat that must be recuperated between the gaseous streams is also smaller. This has been suggested by a previous analysis [1], but the impact of raising this temperature has never been quantified.

In addition to gaseous heat recuperation, there is a significant quantity of high temperature heat in the reactive solid streams as they cycle between the reduction and oxidation temperatures.

Therefore, effective solid-solid heat recuperation has been identified as critical for efficiency operation for quite some time [9]. Operating isothermally would eliminate the need for this solid heat recuperation because the reduction and oxidation temperatures are the same [5, 19], but it does have a separate set of heat recuperation challenges, namely an increased need for high temperature gas heat recuperation. Others have reported that solid-solid heat recuperation can be difficult, but have also illustrated how moderate to high solid heat recuperation can allow for efficient operation at higher reduction pressures [6, 7, 9, 15, 20].

The removal of oxygen from the reduction zone is critical for efficient hydrogen production in two-step metal oxide solar thermochemical redox cycles. In general, two methods have been suggested for removing the oxygen: vacuum pumping and an inert gas sweep [3, 22, 23]. Both of

these methods physically transport the oxygen and lower the oxygen partial pressure to below ambient in order to increase the extent of reduction of the active redox material.

Some previous studies have suggested that using an inert sweep gas to achieve low O2 partial pressures for the reduction reaction can lead to higher system efficiencies than using pumping [5, 6, 24]. In a separate work, we propose a modified approach to the purification of a recycled inert gas using a high temperature recycle [25]. However, many of these works assume an ideal counter-flow arrangement of the inert gas in contact with the reactive solids, which achieves a minimum flowrate which can be optimistically low in some cases and un-realistically equal to zero in other cases [5, 6, 24, 26]. Additionally, the inert gas must either be produced from ambient air or purified from the O2 released in the reduction reaction, and this introduces a new energy cost. Previous works have found the energy required to produce N2 from ambient air to be significant [6, 8, 26, 27] while others have found the energy required to purify a recycled inert gas stream to be negligible [5, 24]. Typically, these N2/O2 separation processes rely on either cryogenic separation or near-ambient temperature pressure swing absorption.

Some studies have suggested that in a process at scale, vacuum reduction leads to higher efficiencies than using an inert gas sweep [11, 27]. However, this type of large scale vacuum process is difficult to implement. Ermanoski [28] showed that a single vacuum pump is unlikely to achieve large scale O2 removal at the low pressures required for solar thermal water splitting due to the huge volumetric flowrates and associated the associated gas velocities; therefore, a cascade pressure reduction system was suggested. In such a system, several chambers are used in which a pump pulls a vacuum on each individual chamber and each chamber has a sequentially lower pressure. Brendelberger and Sattler consider cascade systems with one, two, and three chambers, and found that while the system efficiency increase was significant, it suffered from

diminishing returns even on the third chamber, and that the total system efficiency increase was

<5% [29]. Additionally, others have suggested that previous analyses over-estimate the efficiency at which vacuum pumps operate, leading to a drastic increase in pump work required [8, 30].

Recent studies have also examined the possibility of highly efficient inert gas sweeping.

Several gas flow configurations and production methods have been proposed. Some prior works

[8, 27] only considered a mixed gas system instead of a counter-flow arrangement, and the much higher inert gas flowrates calculated in this way tend to show that vacuum reduction is more efficient, even with limited pump efficiency values [8]. Other works compared counter-flow to perfect mixing [24] and parallel-flow [6, 18] arrangements, and showed that the counter-flow was more efficient, but suggested that the final flow arrangement would be an intermediate state.

Brendelberger et al. [26] examined limitations on the counter-flow arrangements and developed models to approximate this middle ground. Past work has suggested producing nitrogen continuously from surrounding air, but the energy requirements for this method of operation are very high [27]. Others have suggested recycling inert gas, rather than producing it in an open-loop fashion [5, 6]. Still others have suggested using both an inert sweep gas and vacuum pumping simultaneously [24]. In addition to the direct study of inter gas sweeping, other works have compared the energy required for inert gas and vacuum [17, 24].

In this work, we detail a thermodynamic model for the calculation of overall STH efficiency. Specifically, we expand on previous work by offering new insights from a material besides the commonly-used ceria, including showing what factors become more or less important with a different material. In addition, this work includes oxidation kinetics for both materials in a general way rather than using experimental data directly to examine the impact of kinetics on both system efficiency and the identification of ideal operation points.

Furthermore, this work offers new insights into gas and solid heat recuperation for solar thermochemical water splitting, particularly as they relate to a recycled inert sweep gas. We quantify the impact and relative importance of gas and solid heat recuperation using such a sweep gas configuration, and highlight specific conditions in which these become more or less important.

Additionally, this work explores the effect of increasing the inert gas flowrate, and shows the effect that this has on the system efficiency and how it increases or decreases the relative importance of other parameters. This work also explores the temperatures at which both the N2/O2 (for reduction) and H2/H2O (for oxidation) separation occurs, and shows the conditions at which an increase in these temperature will have beneficial effects on system efficiency.

Finally, the thermodynamic model is used explore the effect of O2 removal methods on the overall solar-to-hydrogen (STH) efficiency. Specifically, we expand on previous studies that consider the vacuum pump efficiency by considering a wide range of vacuum pump efficiency values, including but not limited to currently available pump technologies. We also quantify the impact of cascade pressure reduction on STH efficiencies. Lastly, we propose the use of a recycled inert sweep gas with a high temperature purification step and illustrate its impact on system efficiency, along with a direct comparison of all of these possibilities.

2.3. Methods

The STH efficiency is calculated using a system normalized to producing 1 mole of H2.

This model is meant to illustrate trends within a solar thermochemical water splitting system by examining the influence of different parameters on system losses that are un-avoidable and necessary to operation of the system. A real system will have a number of other trade-offs and energy losses, meaning that a real system will likely have lower efficiency values than those reported here. Below, the thermodynamic model will be described in detail. Oxidation kinetics

will then be included and the method for applying them to the thermodynamic model will be described. Finally, details will be given for all model derivations, inputs, and calculations in order to provide a guide for future work.

2.3.1. Efficiency Calculations for Thermodynamic Equilibrium

The model consists of temporal heat fluxes into and out of the system and heat flows within the system that are normalized to the production of a single mole of hydrogen. The system models ceria and a ferrite/zirconia composite as the active materials, and a description of the thermodynamics of these materials is included. The model assumes that a recycled inert sweep gas is used to maintain a low O2 partial pressure for the reduction reaction and that excess steam is used for oxidation. Heat recuperation for both of these gaseous streams is included, as well as solid heat recuperation for the active materials. Equations for all of these heat flows will be given explicitly, as well as any constants used.

2.3.1.1. Overall System Description

We consider an overall process, as shown in Figure 2.1, in which concentrated solar energy drives a two-step metal oxide thermal reduction and oxidation reaction cycle. A continuous process is assumed in which reactive material flows around a process to avoid batch operation [3, 27, 31].

Filled block arrows denote thermal temporal fluxes, while empty block arrows represent heat flows normalized per mole of H2. The thermal source of each of the intra-system thermal requirements

(empty block arrows) is assumed to either come from an intra-system heat benefit or the thermal flux available for thermochemical reaction (Q̇ TC), while the thermal sink for each of these requirements is the associated unit operation or stream that the empty arrow indicates. Similarly, empty block arrows for the intra-system heat benefits point from the heat source out into the

system, where the heat sinks for these benefits are assumed to be either other intra-system heat requirements or heat rejected from the system (Q̇ REJECT).

Figure 2.1: Model process schematic, not to scale. Line arrows indicate mass flows, while block arrows indicate energy flows. A block arrow directed into a block is a heat requirement, a block arrow directed out of a block is a heat benefit. A solid filled block arrow denotes an energy flux per unit time, and empty block arrow denotes an energy requirement per mole of hydrogen. The grey cross-hatched block arrows denote energy flow per mole of hydrogen that is rejected. The empty block arrows sum to equal the QMOL term. 2.3.1.2. Reduction and Oxidation Chemistry

The ceria active material operates via an O-vacancy mechanism, as shown in Equations

(2.19) and (2.20). The amount of metal oxide required to generate a mole of H2 is given by the difference in non-stoichiometry (δ) between the reduction and oxidation thermodynamic states (Δδ

= δRED - δOX). The notations MyOx-δOX and MyOx-δRED denote a generic non-stoichiometric metal oxide in the oxidized state and reduced state, respectively.

1 1 1 푀 푂 → 푀 푂 + 푂 (2.19) 훥훿 푦 푥−훿푂푋 훥훿 푦 푥−훿푅퐸퐷 2 2

1 1 푀 푂 + 퐻 푂 → 푀 푂 + 퐻 (2.20) 훥훿 푦 푥−훿푅퐸퐷 2 훥훿 푦 푥−훿푂푋 2 However, not all solar thermochemical water splitting materials operate via an O-vacancy reaction mechanism, such as ferrites. For these metal oxides, analogous expressions can be written using the overall conversion of the reactants, χ. This is based on a generic metal oxide, MyOx, undergoing a stoichiometric reduction and oxidation cycle, such as one shown in Equation (2.21).

푦 훥휒 푀 푂 → (1 − 휒 )푀 푂 + 휒 푀푂 + 푂 (2.21) 푦 푥 푅퐸퐷 푦 푥 푅퐸퐷 푥 − 푦 2 2 Based on this general equation, the reduction and oxidation cycle can be re-written in an analogous way to Equations (2.19) and (2.20), where the oxidation reaction does not necessarily fully re-oxidize the material to MyOx, but rather to some mixture of MyOx and MO. These expressions are given in Equations (2.22) and (2.23). For the remainder of the discussion, Δδ will be used, but the analogous Δχ expression could also be used.

1 − 휒 휒 푦 1 − 휒 휒 푦 1 푂푋 푀 푂 + 푂푋 푀푂 → 푅퐸퐷 푀 푂 + 푅퐸퐷 푀푂 + 푂 (2.22) 훥휒 푦 푥 훥휒 푥 − 푦 훥휒 푦 푥 훥휒 푥 − 푦 2 2

1 − 휒 휒 푦 푅퐸퐷 푀 푂 + 푅퐸퐷 푀푂 + 퐻 푂 훥휒 푦 푥 훥휒 푥 − 푦 2 1 − 휒 휒 푦 (2.23) → 푂푋 푀 푂 + 푂푋 푀푂 + 퐻 훥휒 푦 푥 훥휒 푥 − 푦 2 To calculate the Δδ value, the thermodynamic extent of reduction is determined for both the reduced and oxidized state of the material at the reduction and oxidation temperature and pressure conditions of interest. For ceria, the model from Bulfin et al. [8] is used, which utilizes data from Panlener et al. [32]. For the ferrite zirconia mixture, calculations were performed using the FactSage thermodynamic prediction software[33] for a 20%-wt. CoFe2O4 mixed with ZrO2; the results of these thermodynamic predictions are shown in the Supplemental Material of Ref

[34]. To calculate the extent of reduction of the material in the reduced state, the reduction

temperature is used in conjunction with the reduction pressure as the partial pressure of oxygen.

To calculate the extent of reduction of the oxidized material, the oxidation temperature is used along with the equilibrium partial pressure of oxygen at the oxidation temperature and overall pressure (1 atm) due to water thermolysis (pOX). The water thermolysis equilibrium values were also calculated using the FactSage thermodynamic prediction software. Details and results of the

FactSage calculations for oxygen partial pressure based on water thermolysis can be found in the

Supplemental Material of Ref [34].

2.3.1.3. Overall System Efficiency

The overall STH efficiency is defined as the ratio of the chemical energy in hydrogen

(using the higher heating value, HHV) to the amount of earth-incident solar energy required to produce that hydrogen by the proposed process. This relationship is shown in Equation (2.24), where ηSTH is the STH efficiency, ṅ H2 is the molar flow rate of hydrogen, HHVH2 is the higher heating value of hydrogen (286.61 kJ/mol = 142.18 MJ/kg [35]), and Q̇ SOLAR is the amount of solar energy per unit time incident on the concentrating mirrors.

푛̇ 퐻2퐻퐻푉퐻2 휂푆푇퐻 = (2.24) 푄̇푆푂퐿퐴푅 The molar flow rate of hydrogen is calculated from the thermal power available for the thermochemical reaction, Q̇ TC, and the amount of thermal energy needed to produce a single mole of hydrogen, QMOL (see Equation (2.25)). This analysis assumes a single mole of H2, and this will be discussed further below.

푄̇ 푇퐶 푛̇ 퐻2 = (2.25) 푄푀푂퐿

2.3.1.4. Solar Energy Fluxes

The relationship between the solar energy, Q̇ SOLAR, and thermal energy available for hydrogen production, Q̇ TC, depends on the heat absorbed by the receiver/reactor, the losses associated with re-radiation, and the efficiency of the solar field which collects and concentrates the incident sunlight; these values are calculated using Equations (2.26), (2.28), and (2.30). These energy flows and losses depend heavily on the exact receiver/reactor design and materials of construction. Because this report is meant to be a general analysis, these losses are simplified, as detailed below. The thermal energy available for the thermochemical reaction, Q̇ TC, is calculated from the effective thermal absorptivity of the solar receiver/reactor α (taken to be 0.95), the thermal energy flux at the receiver (Q̇ A), the amount of energy flux lost due to re-radiation (Q̇ RERAD), and convection (Q̇ CONV) as shown in Equation 8. The heat not absorbed by the receiver (Q̇ NA) is rejected from the system, and is calculated in Equation (2.27). Conduction losses are not considered in this analysis and neither are convective or radiative losses for the non-aperture receiver area, as these are typically small for a well-insulated receiver [4, 36] and are typically excluded in these types of analyses [4-6, 20]. These assumptions enable the calculation of an upper limit on system efficiency without assuming a particular receiver/reactor configuration.

푄̇ 푇퐶 = 훼 ⋅ 푄̇퐴 − 푄̇푅퐸푅퐴퐷 − 푄̇퐶푂푁푉 (2.26)

푄̇푁퐴 = (1 − 훼)푄̇퐴 (2.27) The solar thermal energy at the receiver aperture (Q̇ A) is determined by the solar flux incident on the heliostats and the optical efficiency of the collecting reflectors (ηoptic). This optical efficiency is taken to be 89.3%, based on a single reflection from a reflective surface with an ideal reflectivity of 94% and an average cleanliness factor of 95% [37]. It should be noted that only optical losses are considered here, as these are both generally applicable and unavoidable in concentrated solar thermal systems. By contrast, losses from blocking and shading, cosine effects,

atmospheric attenuation, and mirror misalignment are not considered here. Many of these losses depend greatly on specific field/dish arrangements, geographical location, time of day, time of year, and mirror design. Thus, this analysis is an upper limit on efficiency for the exploration of various trends, rather than an estimation of specific values achievable in a real system. The equation for determining Q̇ A is shown in Equation (2.28). The corresponding amount of solar energy lost due to optical inefficiencies (Q̇ FIELD) is the amount of energy not reflected with the optical efficiency ηoptic, and is shown in Equation (2.29).

푄̇퐴 = 휂표푝푡𝑖푐푄̇푆푂퐿퐴푅 (2.28)

푄̇퐹퐼퐸퐿퐷 = (1 − 휂표푝푡𝑖푐)푄̇푆푂퐿퐴푅 (2.29) The amount of solar energy hitting the solar collectors depends on the solar flux at the

2 earth’s surface (ζsun = 1,000 W/m ) and the total reflective area of the heliostat field or other reflective area, such as a parabolic dish, (Afield) as shown in Equation (2.30). This reflective area of the field can be thought of as an effective reflective area, since mirrors may block and shade one another as well as a variety of other losses that occur when collecting and concentrating solar thermal energy, as outlined above. However, these losses are not explicitly accounted for in this analysis, and so care must be used if these results are compared to other analyses that do take these losses into account explicitly.

푄̇푆푂퐿퐴푅 = 휁푠푢푛퐴푓𝑖푒푙푑 (2.30) The reflective area of the collectors is directly related to the flow rate of product hydrogen, but in this analysis has no direct effect on the final efficiency. This is because the current analysis is based on the production of a single mole of hydrogen in the QMOL term described below. The solar energy flow rate terms (Q̇ SOLAR, Q̇ A, etc.) scale the reflective area directly with the flow rate of hydrogen because Equation (2.25) can be re-arranged to obtain an expression for the ratio of

ṅ H2/Afield, which is a constant. In an actual system this does not hold true and cosine, attenuation, and alignment losses increase with increasing size of the heliostat field.

The losses due to re-radiation at the receiver (Q̇ RERAD) are determined by Equation (2.31), which includes the thermal emissivity of the receiver (ϵ, assumed to be 0.95), the Stefan-

-8 -2 -4 Boltzmann constant (σ = 5.67 × 10 W m K ), the temperature of the reactor wall (Twall), and the aperture area (Aaperture). This simplified and optimistic estimation of radiative losses based on a simple cavity aperture area is meant to be a general estimation rather than based on a calculation of a specific system. The area of the aperture is given by the reflective area of the solar field (Afield) and the concentration ratio (C = Afield/Aaperture). Here the concentration ratio is assumed to be 3,000.

The temperature at the outer wall of the receiver (Twall) is assumed to be 20 K above the reduction temperature.

̇ 4 푄푅퐸푅퐴퐷 = 휖휎퐴푎푝푒푟푡푢푟푒푇푤푎푙푙 (2.31)

Convective losses (Q̇ CONV) are determined by a lumped heat transfer parameter (heff), the area of the cavity receiver aperture (Aaperture), and the difference between the temperature at the receiver wall (Twall) and the ambient temperature (Tamb = 298 K). A more rigorous calculation for convective losses would need to take specific cavity interior dimension into account in addition to the dimensions of the aperture, but this simplified form allows for convective losses to be approximated and to scale with both temperature and aperture size. The effective heat transfer coefficient is estimated from Falcone [36], includes both natural and forced convection, and is scaled to relate to the aperture area rather than the interior absorber area. The value of heff is estimated to be 48.2 W m-2 K-1.

푄̇퐶푂푁푉 = ℎ푒푓푓퐴푎푝푒푟푡푢푟푒(푇푤푎푙푙 − 푇푎푚푏) (2.32)

2.3.1.5. Thermal Energy Required to Produce Hydrogen

The energy required to produce a single mole of hydrogen (QMOL) is determined by the thermal reduction endotherm of the solid material (QHTR = ΔHRED), the energy needed to heat the material from the oxidation temperature to the reduction temperature (QSH), the energy required to heat the inert gas from the inert gas separation temperature to the reduction temperature (QIH), and other auxiliary heating requirements (QAUX), as shown in Equation (2.33). The values for ΔHRED and CP for both ceria and the ferrite/zirconia are shown in Table 2.1.

푄푀푂퐿 = 푄퐻푇푅 + 푄푆퐻 + 푄퐼퐻 + 푄퐴푈푋 (2.33)

Table 2.1: Reduction endotherms and heat capacity for reactive solid materials. The references for the ceria values are shown in the table. The values for the ferrite/zirconia composite were estimated using FactSage. -1 -1 Material ΔHRED [kJ/(1/2 mol O2)] CP [J mol K ] CeO2 480 [38] 80 [20] 20 %-wt CoFe2O4 on ZrO2 309 187.9

The energy required to heat the solid from the oxidation to the reduction temperature (QSH)

solid is given in Equation (2.34). This incorporates the heat capacity of the solid (CP ), the difference between the reduction and oxidation temperatures (ΔT), the solid-solid heat recuperation effectiveness (εSS), and the quantity of material necessary to produce one mole of H2, which is accounted for by the extent of reaction (Δδ) from Equation (2.19), and the fraction of solid material that is thermochemically active (FR), The fraction of solid material that is thermochemically active

(FR) denotes how much of the solid material that must be heated is active rather than an inert support or binder material, which does not contribute towards the thermochemical reaction but still must be heated.

푠표푙𝑖푑 퐶푃 푄푆퐻 = 훥푇(1 − 휀푆푆) (2.34) 퐹푅훥훿

2.3.1.6. Auxiliary Heating Requirements and Benefits

The auxiliary heating requirement includes terms deleterious to the overall efficiency including: water heating requirements (QH2O), heat equivalent of pump work (QPUMP), heat equivalent of mechanical work needed for reactor operation (QMECH), heat equivalent of separations work needed (QSEP); and terms beneficial to the overall efficiency including: oxidation exotherm (QHOX), the un-recuperated heat in the solid flow as it cools (QSC), the un-recuperated heat of the inert gas (if any) as it cools (QIC), and the sensible heat from the oxygen stream from the reduction reaction (QO2). The overall auxiliary heating equation is calculated by Equation

(2.35).

푄퐴푈푋 = (푄퐻2푂 + 푄푃푈푀푃 + 푄푀퐸퐶퐻 + 푄푆퐸푃) − (푄퐻푂푋 + 푄푆퐶 + 푄퐼퐶 + 푄푂2) (2.35) This is an optimistic simplification, as heat recuperation is not always realistic even when possible. There is no specific determination of what heating benefit will heat what specific heating requirement in the equation above; rather, the heat requirements and benefits that are possibly able to interact are compared. This analysis considers what is possible to achieve without assuming a particular system design in order to examine more general trends. QAUX is set to a value of zero if

REJ the initial calculated QAUX value is negative, and this excess heat (QAUX ) is rejected from the system. It should be noted that QAUX can never be negative, as explained by Ermanoski et al.[20].

This is because the heating benefits (QHOX, QSC, and QO2) will reject heat at a temperature near the oxidation temperature, and therefore cannot contribute to the heating requirements at the reduction temperature (QSH, QIH, and QHTR). The oxidation exotherm is given by subtracting the heat of combustion of hydrogen (ΔHc,H2 = 286 kJ/mol) from the reduction exotherm, QHOX = ΔHRED -

ΔHc,H2. The other terms are described in more detail below.

The water heating requirements (QH2O) are calculated in Equation (2.36) and include two

heat terms: water heating (QH2O ), and heat recuperation from the steam and hydrogen mixture as it

cool heat cools (QH2O ). The water heating term (QH2O ) is the energy necessary to heat the calculated amount of water required for oxidation from a liquid phase at ambient temperature to steam at the

heat oxidation temperature. The steam/hydrogen cooling term (QH2O ) is the sensible heat released from the cooling of the one mole of hydrogen formed and the excess water from the oxidation temperature to ambient. This term is then discounted by the heat recuperation effectiveness (εGG).

These expressions incorporate the specific heat capacity at constant pressure (CP) of liquid water, steam, and hydrogen; the molecular weight (MW) of water; and the latent heat of vaporization of water to steam (ΔHvap). The heat capacities are all taken from the NIST Chemistry WebBook [39].

The amount of water required for reaction is given by nwh, which is a ratio of the number of moles of water to 1 mole of hydrogen after oxidation is complete; this is explained below. The un-

REJ recuperated heat of the water/steam mixture (QH2O ) which is rejected to ambient as it cools is equal in magnitude to QH2O, and is shown explicitly in Equation (2.37). Specifics of how the water/hydrogen heating and cooling terms are calculated are given in the Supplemental Material of Ref [34].

ℎ푒푎푡 푐표표푙 푄퐻2푂 = 푄퐻2푂 − 휀퐺퐺푄퐻2푂 (2.36)

푄푅퐸퐽 = 푄ℎ푒푎푡 − 휀 푄푐표표푙 (2.37) 퐻2푂 퐻2푂 퐺퐺 퐻2푂 We assume that the steam and solids are in a well-mixed, counter-flow arrangement that reaches thermodynamic equilibrium at each infinitesimal point along the length of contact. The ratio of excess steam to generated hydrogen (nwh) is calculated by re-arranging the equilibrium

WS constant of the net water splitting reaction (Keq ), as shown in Equation (2.38). This includes the standard enthalpy of water splitting (ΔHWS° = 250.8 kJ/mol) and the standard entropy of water

-1 -1 splitting (ΔSWS° = 57.35 J mol K ) both on a steam basis [20]. The partial pressure of oxygen

(pAO) is then found by calculating the equilibrium partial pressure of oxygen of the reduced material at the oxidation temperature. The value of nwh typically varies greatly depending on the operating conditions for reduction and oxidation being used. An example of the range of nwh values for an example calculation is given in the Supplemental Information of Ref [34].

1/2 ∘ ∘ ∘ 푛 푛 푝 푊푆 훥퐺푊푆 훥퐻푊푆 − 푇푂푋훥푆푊푆 퐻2 푂2 √ 퐴푂 퐾푒푞 = 푒푥푝 (− ) = 푒푥푝 (− ) = = (2.38) 푅푇푂푋 푅푇푂푋 푛퐻2푂 푛푤ℎ The mechanical heat requirements, shown in Equation (2.39), are the heat-equivalent of the work required to move the solids within the reaction system. It should be noted that this term is particularly dependent on the reactor design used, and so the form of this term can vary widely.

For this work, it is assumed that the solid materials are raised some distance h for each cycle, taken to be 10 meters in this analysis. This term is inflated by ηMECH = 10%, which converts useful mechanical work to the heat equivalent. This lumped efficiency term includes thermal-to- electricity conversion, electric-to-pump/motor efficiency, and an estimation of the conversion of pump/motor work-to-useful work done on the solid materials (such as frictional losses for a mechanically driven system or pressure drop losses for a fluidized bed-type system). This mechanical work involves a conversion to heat equivalent, and so the heat rejected in this

REJ conversion (QMECH ) is calculated in Equation (2.40).

푀푊푠표푙𝑖푑𝑔ℎ 푄푀퐸퐶퐻 = (2.39) 퐹푅훥훿휂푀퐸퐶퐻 푅퐸퐽 (2.40) 푄푀퐸퐶퐻 = (1 − 휂푀퐸퐶퐻)푄푀퐸퐶퐻 The overall separations work is a combination of the separations work for water/hydrogen

O2 H2 (after oxidation) and inert/oxygen (after reduction), QSEP = QSEP + QSEP . The separations work

H2 for the water/hydrogen separation (QSEP ) is based on an estimation of the heat-equivalent of the

work required to separate the hydrogen product from the excess water, and is shown in Equation

(2.42). This is based on the 2nd Law of Thermodynamics (ΔQ = T ΔS), which is then inflated by

H2 an assumed separation efficiency, ηSEP = 15%. The entropy of mixing (ΔSmix) for each stream

(feed into separator, product stream out of separator, and effluent stream out of separator) is given in Equation (2.41), where x is the molar fraction of hydrogen in the stream and n is the molar flow rate of that stream. Finally, it is assumed that the final product stream has a purity of 99.9% hydrogen, while the effluent stream has an equilibrium-determined amount of hydrogen dissolved

-5 in it (1.4 x 10 mole H2 per mole H2O at 298 K, from the NIST Chemistry WebBook [40]). The

H2 thermodynamic separation is done at the water/hydrogen separation temperature, TSEP , here assumed to be 373 K (100°C). This separation work involves a conversion to heat equivalent, and

REJ so the heat rejected in this conversion (QSEPH2 ) is calculated in Equation (2.43).

훥푆푚𝑖푥 = −푁퐴푘퐵푛[푥 푙푛(푥) + (1 − 푥) 푙푛(1 − 푥)] (2.41)

푇퐻2 푄퐻2 = 푆퐸푃 (훥푆푝푟표푑푢푐푡 + 훥푆푒푓푓푙푢푒푛푡 − 훥푆푓푒푒푑) (2.42) 푆퐸푃 퐻2 푚𝑖푥 푚𝑖푥 푚𝑖푥 휂푆퐸푃

푅퐸퐽 퐻2 퐻2 푄푆퐸푃퐻2 = (1 − 휂푠푒푝)푄푆퐸푃 (2.43) 2.3.1.7. Recycled Inert Gas Terms

This analysis assumes that a counter-flow recycled inert gas sweep is used for the reduction reaction. A similar method as Ref [5] is used to calculate the amount of inert gas (assumed to be nitrogen) required. We assume a counter-flow inert gas/solids arrangement, and we assume that each infinitesimal point along the solid/gas contact reaches local thermodynamic equilibrium. The amount of inert gas required is calculated as a ratio of moles of inert gas per mole of oxygen (nio), as shown in Equation (2.44). Here, pAR is the oxygen partial pressure in the inert sweep gas stream after reduction, PRED is the total pressure of the reduction zone (PRED = 1 atm = 101,325 Pa for the

inert gas sweep case), and pRED is the oxygen partial pressure of the solid stream after reduction.

This equation compares the amount of oxygen produced by determining the oxygen content of the inert gas stream before and after reduction, then calculating how much inert gas is needed to achieve that level of dilution. pAR is obtained by calculating the equilibrium oxygen partial pressure above the reactive solids that are at the reduction temperature (TRED) and the oxidized stoichiometric state (δOX). It should be noted that the validity of this method of calculating inert gas flow (as well as other inert gas assumptions) will be discussed in subsequent analyses [25, 41].

An example of the range of nio values for an example calculation is given in the Supplemental

Information of Ref [34].

−1 푝퐴푅 푝푅퐸퐷 푛𝑖표 = ( − ) (2.44) 푃푅퐸퐷 − 푝퐴푅 푃푅퐸퐷 − 푝푅퐸퐷 O2 The separation work of the inert gas and oxygen (QSEP ) is shown in Equation (2.45),

O2 where TSEP is the oxygen/inert gas separation temperature (assumed to be 1,223 K = 950°C),

O2 ηSEP is the oxygen/inert separation efficiency (assumed to be 15%), and ΔSmix is the entropy of mixing of each sub-stream (general equation shown in Equation (2.41)). The mole fractions are calculated using the oxygen partial pressures that were used to calculate nio (Equation (2.44)). An oxygen product purity after separation of 100% was assumed for this closed system, and the oxygen partial pressure for reduction (pRED) was used to determine the final effluent oxygen concentration. This separation work involves a conversion to heat equivalent, and so the heat

REJ rejected in this conversion (QSEPO2 ) is calculated in Equation (2.46).

푇푂2 푄푂2 = 푆퐸푃 (훥푆푝푟표푑푢푐푡 + 훥푆푒푓푓푙푢푒푛푡 − 훥푆푓푒푒푑) (2.45) 푆퐸푃 푂2 푚𝑖푥 푚𝑖푥 푚𝑖푥 휂푆퐸푃 푅퐸퐽 푂2 푂2 (2.46) 푄푆퐸푃푂2 = (1 − 휂푆퐸푃)푄푆퐸푃

In addition to the energy required to purify the recycled inert gas, the inert gas must be heated to the reduction temperature. This sensible heating requirement is calculated in Equation

(2.47), where εGG is the gas heat recuperation effectiveness, nio is the inert-to-oxygen molar ratio,

O2 TSEP is the oxygen/inert separation temperature (here assumed to be 950°C = 1,223 K), TRED is

N2 the reduction temperature, and CP is the specific heat capacity of nitrogen gas, which is the assumed inert gas. The factor of ½ applied to the nio ratio is to account for that fact that ½ moles of oxygen are produced per 1 mole of hydrogen. The specific heat capacity of nitrogen is obtained from the NIST Chemistry WebBook [39]. This term is not included in the auxiliary heating requirements (QAUX), but rather in the QMOL term. This is because the heat required is typically

(though not necessarily) at or above the oxidation temperature (TOX), and so will not benefit from any potential heat benefit terms in QAUX.

푛 푇푅퐸퐷 𝑖표 푁2 ′ ′ 푄퐼퐻 = (1 − 휀퐺퐺) ∫ 퐶푃 (푇 ) 푑푇 (2.47) 2 푂2 푇푆퐸푃 This quantity of inert gas used in the process is normally considered to be the amount required to achieve a level of dilution of the oxygen, but not to provide any sort of convective transport of the oxygen. As such, it is to be considered a minimum. This is a logical choice as this analysis are meant to show thermodynamic limits on the system. Previous analyses [6, 26] have pointed out that for certain conditions, a flowrate of 0 moles of inert gas is calculated to be required and that this condition is not physically realizable. These authors suggested that the true amount of inert gas flow would likely lie somewhere in between a true counter-flow and a true parallel- flow arrangement. Here, we attempt to quantify the impact of a high temperature separation recycle using more realistic gas flowrates. The minimum nio is calculated using Equation (2.44), but we then subsequently alter it in two different ways to illustrate how a more realistic system might behave.

First, the ratio is multiplied by a scaling factor ranging from 1 to 1,000; this factor is chosen arbitrarily, as the true scaling factor will depend heavily on reactor configuration and other design choices. The second way in which the inert gas flowrate is altered is to assume that some minimum inert flow is required to provide convective transport of the oxygen away from the solids. As such, a lower limit is set on the value of nio. As with the scaling multiplier, this lower limit is chosen arbitrarily with values from nio = 0 to 1,000. When a value of nio is calculated (using Equation

(2.44)) that is lower than this bound, the value of nio is set equal to the value of the lower bound itself for the purposes of heat requirement calculations. However, we do not alter the thermodynamic state of the material based on this change. Thus, the heating requirements are impacted by a higher-than-calculated inert gas flowrate, but the reactive solids do not experience a corresponding increase in reduction. This is not how a physical system would necessarily behave, and is meant to illustrate the impact of different inert gas flow rates.

2.3.1.8. Vacuum Pump Case

Vacuum pump work (shown in Equation (2.48)) incorporates the heat-equivalent of the work necessary for pumping oxygen. This was assumed to be zero in a previous analysis which considered only inert sweeping [34], but is considered here because of the changing reduction pressure. The heat equivalent is calculated for pumping one half mole of oxygen from the reduction pressure (pRED) to ambient pressure (Pamb = 1 atm = 101,325 Pa). This is done at ambient temperature (Tamb). Here, the reduction oxygen partial pressure (pRED) is equal to the total pressure of the reduction zone (PRED). An efficiency factor (ηPUMP) is used to convert from pump work to heat equivalent. This factor is calculated based on the reduction pressure being considered, and an interpolation of the pump efficiency data from Bulfin et al. [8]. Details of this calculation, as well as a plot showing the pump efficiency ηPUMP used is shown in the Supplemental Material of Ref

[25]. If inert gas is to be used (as described in Section 2.3.1.7), then the pump work (QPUMP) is set to zero.

푛푂2푅푇푎푚푏 푃푎푚푏 푄푃푈푀푃 = 푙푛 ( ) (2.48) 휂푃푈푀푃 푝푅퐸퐷 Since pump work involves a conversion from pump work to heat equivalent, there is some

REJ amount of heat rejected in that conversion. This rejected heat for the pump work (QPUMP ) is calculated in Equation (2.49), and contributes to the total rejected heat per mole of hydrogen

MOL (QREJ ).

푅퐸퐽 푄푃푈푀푃 = (1 − 휂푃푈푀푃)푄푃푈푀푃 (2.49) 2.3.1.9. Cascade Pressure Reduction Case

In addition to the single-pump work term shown above, we also consider a cascade pressure reduction method. In cascade pressure reduction, each pump is assumed to have an identical volumetric flow rate. A schematic of this type of cascade pumping is shown in Figure 2.2.

Figure 2.2: A schematic (not to scale) of a cascade pressure reduction system for NC chambers. Each pressure reduction chamber has an equally sized vacuum pump attached to it, which pumps an equal volumetric flow rate of oxygen at an efficiency based on the pressure of that stage of the cascade. The molar flow rate of oxygen for each chamber is based on the temperature and pressure of that chamber. It is assumed the reactive solids reach thermodynamic equilibrium in each chamber. Chamber 0 is a chamber with no pump for when the equilibrium partial pressure of oxygen is >1 atm.

An effective pump efficiency of the system is calculated by dividing the work required to pump the oxygen in a single, perfectly efficient pump by the total work required by the cascade pump series. This is done in order to directly compare pump efficiency values for this and the

eff above single-pump systems. The final expression is shown in Equation (2.50). Here, ηPUMP is the effective pump efficiency, ṅ O2 is the total molar flow rate of oxygen per unit time, Pamb is the ambient pressure, pRED is the final reduction pressure, NC is the number of chambers considered,

i th i ṅ O2 is the molar flow rate per unit time of oxygen from the i chamber, ηPUMP is the pump

th th efficiency at the pressure of the i chamber, and pi is the pressure of the i chamber. The cascade pump system and the sizes (volumetric flow rates) of the pumps are solved iteratively, using equations from Ermanoski [28]. After the effective pumping efficiency is calculated, the pump work (QPUMP) is calculated using Equation (2.48). Details on how this effective pump efficiency was calculated are given in the Supplemental Material of Ref [25].

−1 푁퐶 푃 푛̇ 𝑖 푃 푒푓푓 푎푚푏 푂2 푎푚푏 (2.50) 휂푃푈푀푃 = 푛̇ 푂2 푙푛 ( ) [∑ 𝑖 푙푛 ( )] 푝푟푒푑 휂 푝𝑖 𝑖=1 푃푈푀푃 Additionally, the effective pump efficiency for cascade pressure reduction is also considered for the limiting case of an infinite number of reduction chambers. The resulting equation is shown in Equation (2.51), and this is consistent with an independently derived expression for the upper limit of vacuum pump efficiency [30]. Here, ṅ S is the solids flow rate per unit time, δRED is the extent of non-stoichiometry of the fully reduced material, and δ0 is the extent of non-stoichiometry of the material as it enters the infinite cascade pressure reduction system.

Additional details are available in the Supplemental Material of Ref [25].

푃 −1 푙푛 ( 푎푚푏 ) 2푛̇ 푃 훿푅퐸퐷 푝 (훿′) 휂푒푓푓,𝑖푛푓 = 푂2 푙푛 ( 푎푚푏) [∫ 𝑖 푑훿′] (2.51) 푃푈푀푃 푛̇ 푝 𝑖 ′ 푆 푅퐸퐷 훿0 휂푃푈푀푃(훿 )

2.3.1.10. Heat Benefits

There are also a number of heat benefits within the system. In addition to the oxidation exotherm (QHOX) as a heat benefit to QAUX, the un-recuperated heat in the solid reactive material as it cools down (QSC) is also a thermal credit. This credit is equal in magnitude to the un- recuperated heat required to heat the solids (Equation (2.34)); the only difference is that QSC is the un-recuperated heat of the solids as they cool, whereas QSH is the un-recuperated heat of the solids as they are heated. The equation for QSC is shown in Equation (2.52).

푠표푙𝑖푑 퐶푃 푄푆퐶 = 훥푇(1 − 휀푆푆) (2.52) 퐹푅훥훿 There is also a sensible heat benefit from the inert gas as it cools from the reduction temperature to the separation temperature. This heat is especially useful, as all of this heat is at

O2 temperatures somewhere between the relatively high oxygen separation temperature (TSEP ) and the even higher reduction temperature (TRED). This heat benefit is equal in magnitude to the inert gas sensible heat requirement (Equation (2.47)), and is shown in Equation (2.53) for completeness.

This term is included in QAUX because, while the inert/oxygen separation temperature is high, it can be at or below the oxidation temperature.

푇푅퐸퐷 푛𝑖표 푁2 ′ ′ 푄퐼퐶 = (1 − 휀퐺퐺) ∫ 퐶푃 (푇 ) 푑푇 (2.53) 2 푂2 푇푆퐸푃 Finally, the sensible heat credit from the oxygen is shown in Equation (2.54). This includes

O2 the specific heat capacity of oxygen (CP ), the amount of oxygen (nO2 = 0.5 moles), and the difference between the reduction and ambient temperatures (Tred and Tamb). The entire temperature

from reduction to ambient is considered, because the 0.5 moles of oxygen are cooled over that entire range, even though there is a separation step somewhere within that range. Sensible heat could be recovered both above and below the separation temperature. The heat capacity of oxygen is given by the NIST Chemistry WebBook [39].

푇푅퐸퐷 푂2 ′ ′ 푄푂2 = 푛푂2 ∫ 퐶푃 (푇 ) 푑푇 (2.54) 푇푎푚푏 2.3.1.11. Heat Rejection

The system boundary shown in Figure 2.1 includes a number of rejected heat fluxes. Most of these fluxes originate from the solar field and receiver, as described above, but there is also the

MOL heat rejected from the thermochemical process that produces one mole of H2: Q̇ REJ . This heat rejection flux is calculated in Equation (2.55), and includes heat rejection from the near-ambient

REJ REJ steam cooling (QH2O ), heat not able to be used by auxiliary terms (QAUX ), as well as rejected

REJ heat from the conversion from heat to useful work for mechanical motion (QMECH ), pump work

REJ REJ REJ REJ (QPUMP ), and separation (QSEP = QSEPO2 + QSEPH2 ). Thus, the total heat rejected from the system (Q̇ REJECT) is calculated in Equation (2.56) and consists of all of the heat fluxes from both the solar field, solar receiver, and thermochemical process. The total heat input to the system (Q̇ IN,

Equation (2.57)) and the total heat output from the system (Q̇ OUT, Equation (2.58)) differ by only

0.0004% of Q̇ IN. This small discrepancy in the model is due to the difference between temperature and pressure conditions of the H2O input/reactant stream and the H2 and O2 output/product streams, and the mismatch associated with the standard conditions assumed in the HHV value.

푄̇ 푀푂퐿 = 푛̇ (푄푅퐸퐽 + 푄푅퐸퐽 + 푄푅퐸퐽 + 푄푅퐸퐽 + 푄푅퐸퐽) (2.55) 푅퐸퐽 퐻2 퐻2푂 퐴푈푋 푀퐸퐶퐻 푃푈푀푃 푆퐸푃 ̇ ̇ ̇ ̇ ̇ ̇ 푀푂퐿 푄푅퐸퐽퐸퐶푇 = 푄퐹퐼퐸퐿퐷 + 푄푁퐴 + 푄푅퐸푅퐴퐷 + 푄퐶푂푁푉 + 푄푅퐸퐽 (2.56)

푄̇퐼푁 = 푄̇푆푂퐿퐴푅 (2.57) ̇ ̇ (2.58) 푄푂푈푇 = 푛̇ 퐻2 퐻퐻푉퐻2 + 푄푅퐸퐽퐸퐶푇

2.3.2. Kinetic Limitations

Kinetic limitations are taken into account by determining the relative extents of reaction at the reaction rates at a given oxidation temperature and the highest possible oxidation temperature, i.e. the isothermal operation temperature. This is done by calculating the time (tcritical) required to reach full conversion (> 0.999) of the oxidation reaction at the highest oxidation temperature considered (ΔT = 0) using existing kinetic models of solar thermochemical water splitting oxidation. For each oxidation temperature lower than the maximum temperature (ΔT > 0), the reaction conversion is found for the same oxidation time (tcritical). This extent of conversion is then multiplied by the thermodynamic efficiency to calculate the kinetic-corrected efficiency. We use the kinetic model of cerium oxide and ferrite/zirconia as determined by Arifin [42] and Scheffe et al.[43], respectively. Details of these expressions, as well as an example calculation, are given in the Supplemental Material of Ref [34].

2.3.3. Calculation and Analysis

Using the expressions described above, a computer code was written and executed using the MATLAB computing language and software (MATLAB Release 2015a, The MathWorks, Inc.,

Natick, Massachusetts, United States). A single value was evaluated for both the reduction temperature (TRED) and pressure (pRED), as many others [5, 7, 20, 44, 45] have considered the effects of these parameters on efficiency, and generally find that increasing the reduction temperature increases STH efficiency. A range of oxidation temperature (TOX) were considered, and are directly related to TRED by the temperature difference between reduction and oxidation (ΔT

= TRED – TOX). Because TRED is constant, as ΔT increases, TOX decreases (TOX = TRED – ΔT). A

TRED of 1,600 K with ΔT values of 0 to 250 K (step size of 10 K) were considered for the ferrite/zirconia composite in order to avoid issues with the thermodynamic predictions for the

ferrite at higher temperatures, such as slag formation and volatilization. The same temperatures are considered for ceria in order to make more direct comparisons between the two materials, but for other analyses where ceria alone is considered, ΔT values from 0 to 800 K (step size of 10 K) where used.

A reduction partial pressure for oxygen of 0.1 Pa was typically used, though partial pressures from 0.1 to 100,000 Pa were used for other analyses. Both the reduction and oxidation reactions are assumed to take place at 1 atmosphere (101,325 Pa), except when vacuum is used for reduction, then the absolute pressure matches the oxygen partial pressure. Only a single value of

FR = 1 (100% active material) is considered, as the addition of inert solid material will decrease efficiency in a predictable way. Solid-solid heat recuperation efficiencies (εSS) and gas-gas heat recuperation efficiencies (εGG) from 0-100% were included to illustrate bounds on the possible effects of these parameters; additional analysis of the impact of these factors on the system efficiency is given in Ref [41]. A full set of example calculations are given in of the Supplemental

Material of Refs [25, 34, 41] listing conditions, constants, and calculated values for all major terms in the efficiency calculations. Additionally, while the main focus of this work is on the thermochemical process rather than the solar field and receiver performance, a summary of various efficiency terms to show the relative performance of the field, receiver, and thermochemical parts of the plant are shown in the Supplemental Material of Ref [34].

Solid heat recuperation can occur between the active material as it is heated and cooled as it cycles between the reduction and oxidation temperatures; the fraction of the amount of this heat that is recuperated in a useful way is described by the solid-solid heat recuperation effectiveness

(εSS). Gas heat recuperation can occur between both the steam before and after it enters the oxidation reactor as well as the inert gas as it cycles between the reduction reactor and the

separator. Similar to the solids, the fraction of this heat that is recovered is described by the gas- gas heat recuperation effectiveness (εGG). Values for εSS and εGG from 0-100% were included to illustrate bounds on the possible effects of these parameters, and a discussion of the implications of this full range is given in the Supplemental Material of Ref [41]. Water/H2 separation temperatures from 373 K to 1,000 K were assessed; the lowest value is the condensation temperature of steam and the highest value is the oxidation temperature at the highest value of ΔT.

2.4. Results

System efficiency values for both ceria and ferrite/zirconia active redox materials are shown here. These values include comparisons to work by others on similar efficiency models.

Differences between the ceria and ferrite/zirconia efficiency are shown. The effect of oxidation kinetics are then illustrated and shown for both materials. Next, the effect of various limitations on the idealized inert gas flowrate are illustrated, and the impact of the high temperature recycle separation will be explored. The impact of raising the water/hydrogen separation temperature will be described, as this is closely related to the gas heat recuperation effectiveness. System efficiency values will then be presented for three methods of achieving low oxygen partial pressures. Direct vacuum pumping will be considered first, followed by cascade pressure reduction, and then an inert gas case will be presented. For each case, the sensitivity of critical assumptions on system efficiency will also be shown.

2.4.1. Thermodynamic Efficiency of Ceria and Ferrite/Zirconia Cycles

Efficiency values for both ceria and the ferrite/zirconia composite are shown in Figure 2.3a and b respectively for various values of gas heat recuperation effectiveness.

a) b)

Figure 2.3: Solar-to-hydrogen efficiency for a) ceria and b) ferrite/zirconia active redox materials at various levels of gas heat recuperation (εGG) and a reduction temperature of 1,600 K, a reduction pressure of 0.1 Pa, and a solid heat recuperation effectiveness of 50%. Generally, the efficiency values for ceria are consistent with previous efficiency analyses using ceria [5-7, 20]. They show an increasing efficiency with increasing ΔT up to an optimum, and then the efficiency value decreases. However, the efficiency of the ferrite/zirconia system behaves quite differently. The STH efficiency increases monotonically up to ΔT = 250 K.

a) b)

Figure 2.4: Heat loads and benefits for a) ceria and b) ferrite/zirconia active redox materials with gas heat recuperation of 90%, a reduction temperature of 1,600 K, a reduction pressure of 0.1 Pa, and a solid heat recuperation effectiveness of 50%. The reason for this discrepancy between the shape of the efficiency curves between ceria and ferrite/zirconia stem from the various heat requirements as shown in Figure 2.4. The main

difference in the shapes of the efficiency curves is the behavior of the QAUX term. The slope of the efficiency curve changes when this term goes to zero. As can be seen in Figure 2.4, the QAUX for ceria starts at zero due to the high value of the QIC heat benefit relative to the other terms in QAUX.

When this QIC benefit value decreases significantly, QAUX becomes >0, leading to the slope change at a ΔT ≈ 25 K. However, the QH2O term continues to decrease as the difference between the reduction and oxidation temperatures (ΔT) increases, leading to QAUX going to zero again. This yields the slope change at the ηSTH maximum. By contrast, the ferrite/zirconia composite has a

QAUX term that starts much higher and does not approach zero until higher values of ΔT. This is due to a number of reasons, including the fact that the QH2O term is much larger than the other terms in QAUX for the ferrite/zirconia composite. It should also be noted that the magnitudes of

QAUX and associated terms are much smaller for ferrite/zirconia than for ceria; this is why the ηSTH for the ferrite/zirconia composite is so much larger than ceria.

Figure 2.4 shows that the solids cooling heat benefit (QSC) is much larger for ceria than for ferrite/zirconia, despite the fact that the heat capacity for ferrite/zirconia (187.9 J mol-1 K-1) is much higher than ceria (80 J mol-1 K-1). This occurs because the normalized solids flow rate (1/Δδ) for the ferrite/zirconia is much smaller than for ceria (see Figure 2.5). Because a smaller solids flow rate is needed, the solids cooling benefit (and the solids heating requirement) is correspondingly smaller for ferrites than for ceria.

Figure 2.5: Values for Δδ for ceria and ferrite/zirconia at a reduction temperature of 1,600 K and a reduction pressure of 0.1 Pa. 2.4.2. Illustration of Kinetics

The kinetic effects of oxidation on the overall STH efficiencies are taken into account by discounting the total efficiency values by the oxidation reaction rate for ceria and ferrite/zirconia as outlined in Section 2.3.1.11. These results are shown in Figure 2.6, and show that the impact of oxidation kinetics on ceria is minimal but significant for ferrite/zirconia. Even for the largest value of ΔT considered here (250 K), which correspond to low values of oxidation temperature, the overall impact on ceria efficiency is <1%. This is not a trivial efficiency loss, but for the vast majority of ΔT values likely to be used, the oxidation kinetics have little noticeable impact on ceria efficiency. However, for ferrite/zirconia the impact is large. At high values of ΔT, the kinetics effect on efficiency is >50% of the thermodynamic efficiency value.

a) b)

Figure 2.6: Efficiency values with efficiency discounts for various values of εGG for a) ceria and b) ferrite/zirconia both with a reduction temperature of 1,600 K, a reduction pressure of 0.1 Pa, and a solid heat recuperation effectiveness of 50%. Slow kinetics may be overcome by increased residence time or solid flowrates in the reactor, but these increases have an impact on reactor performance and cost. Here, we attempt to quantify the effects of the oxidation kinetics on the thermodynamic efficiency to illustrate the impact of lower oxidation temperatures. While we are aware that an operating reactor would not lose productivity as simply as described, it does illustrate the general effects of different reaction rates on efficiency. An additional illustration of the time needed to overcome kinetic limitations at different oxidation temperatures is shown in the Supplemental Material of Ref [34].

2.4.3. Inert Sweep Gas Flow Rate

The impact of the inert gas flow rate in previous works [5, 6, 24, 26] has been examined by considering a perfect counter-flow inert/solid arrangement (as has been done here) or by considering a perfect parallel flow arrangement. Here we examine this effect by using the counter- flow assumption (as this is much more likely to be used in a real system) but we increase the inert gas flowrate above the minimum.

2.4.3.1. Effect of Non-Minimum Inert Gas Flowrates on System Efficiency

Here we show the results for two types of increases: 1) a direct multiplication of the minimum inert gas flowrate by an arbitrary value, and 2) a lower limit on the inert gas flow, where the calculated flowrate is used if it is above this limit or the flowrate is set equal to the limit if the calculated value is less than the limit. These results are shown in Figure 2.7 and Figure 2.8, respectively. The resulting inert gas flowrate profiles are shown in the Supplemental Material of

Ref [41]. Of particular interest in this and subsequent results is the point at which ηSTH reaches a maximum for a single set of conditions (each line on the plot) and the ΔT value at which this occurs is denoted as ΔTopt.

Figure 2.7: System efficiency (ηSTH) for inert gas sweep system with various multipliers of the minimum inert flowrate. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5. The multiplication constant inflation (direct manipulation) of the minimum inert gas flowrate follows the same general behavior as the minimum flow (nio mult = 1), but exacerbates the deleterious effect of the inert gas flow rate where ΔT < ΔTopt. Specifically, the energy requirements of inert gas heating and separation increase, which lowers the overall efficiency. The lowering of system efficiency is greatest at points at which the minimum inert gas flowrate is large

(e.g., low ΔT values). However, at conditions where the inert gas flowrate is already low or zero

(e.g., high ΔT values) energy requirements are unaffected or negligibly unaffected by the direct multiplication. This results in the overall efficiency of the system decreasing with increasing inflation factors and results in larger ΔTopt. The larger ΔTopt with increasing inflation factors arises from higher oxidation extents achievable with larger temperature swings, and therefore less material must be cycled, and therefore less inert gas must be used. At very high ΔT the inert gas flowrate, both inflated and minimal, becomes zero, which is physically unrealizable. This is addressed below by imposing a minimum inert gas flow for all calculations as shown in Figure

2.8. The resulting inert gas flowrates are shown in the Supplemental Material of Ref [25].

Figure 2.8: System efficiency (ηSTH) for inert gas sweep system with various lower limits on the minimum inert gas flowrate. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5. The imposition of a minimum inert gas flow has the opposite effect on the system efficiency than a direct multiplication of the minimum. Rather than altering the shape of the curve at ΔT < ΔTopt while being consistent at ΔT > ΔTopt, the results are consistent at ΔT < ΔTopt, and differ at ΔT > ΔTopt. This is because the direct multiplication of the minimum inert gas flowrate has the greatest impact when the minimum flow is already high, and less impact when the

minimum flow is low (or zero). However, the minimum flow limit has no impact when the minimum inert gas flowrate is high, but a substantial impact when the inert gas flowrate is small.

This is because some of the heat requirements that tend towards zero when the inert gas flowrate goes to zero are maintained at significant values.

2.4.3.2. Impact of Gas Heat Recuperation on Non-Minimum Inert Gas Flowrates

The gas heat recuperation does not appear to have a large impact on optimal system efficiency for the optimistic inert gas flowrate case, as seen in Figure 2.9a (additional values given in the Supplemental Material of Ref [41]). However, when the inert gas flowrate is not allowed to get down the unrealistic value of zero, the gas heat recuperation effectiveness (εGG) has a much larger impact on system efficiency. This is shown clearly in Figure 2.9b, in which the nio value was not allowed to go below a value of 100. As expected, the differences in system efficiency

(ηSTH) are much more pronounced at different values of gas heat recuperation. This behavior occurs because with greater heat recuperation, less solar thermal energy is spent heating the inert sweep gas.

a) b)

Figure 2.9: System efficiency (ηSTH) for inert gas sweep system with various gas heat recuperation effectiveness values (εGG) with a a) the ideal counter-flow minimum and b) a lower limit of 100 on the inert gas flowrate. All of these values are calculated for ceria at a reduction

temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, and a solid heat recuperation effectiveness (εSS) of 0.5. 2.4.4. Inert Sweep Gas/Oxygen Separation

There are two ways to provide the necessary inert sweep gas, continual production from air in an open loop fashion, or recycle. If recycling is adopted, the temperature at which the gas is purified is a critical design point. To assess these designs, the effect of the previously examined

O2 high separation temperature TSEP = 1,223 K (950°C) is compared to low temperature separation at 90 K (-183°C), the temperature at which cryogenic N2/O2 separation can occur. For each of these temperatures, the separation work is calculated for the recycle as described above, and for an open loop system in which a feed of 21 mol% O2 and 79 mol% N2 is used. For the open loop system, the sensible energy to heat the feed stream from ambient temperature to the high temperature case is also calculated, which is then discounted using εGG by the pure O2 stream from the separator (details are provided in the Supplemental Material of Ref [41]). In all four of these cases, the separation work is calculated and added to the sensible heat required. The results of this for all ΔT and pRED values considered is shown in the Supplemental Material of Ref [41], and an example result for pRED of 0.1 Pa is shown in Figure 2.10.

Figure 2.10: Combined separation and sensible heat required for inert/O2 separation and heating, based on a high separation temperature of 1,223 K and a low temperature of 90 K. All calculations are done for ceria at a reduction temperature of 1,800 K, a reduction partial pressure of 0.1 Pa, and a gas heat recuperation effectiveness of 0.9.

These results show the energy required for those specific aspects of the system, and since the results are normalized to a single mole of H2, more energy will lead to a lower efficiency. The two low temperature separation cases in Figure 2.10 require much more sensible energy for heating the inert gas (QIH), since they require a much larger temperature change than the high temperature

O2 separation cases. The separation work (QSEP ) is directly proportional to the separation

O2 temperature (TSEP ), meaning that the separation work will be higher for the high temperature separation cases. The separation work also depends on the difference in O2 purity of the input and output streams, which is much higher for the open loop cases. Finally, the open loop, high temperature separation case must also heat the ambient air to the separation temperature (QOLH).

This explains why the high temperature open loop case in Figure 2.10 shows the highest combined amount of energy required; the open loop oxygen purity difference and high separation temperature both increase the separation work, and these aspects also necessitate extra heating of the ambient air before separation. However, the high temperature recycle case is much lower than both of the open loop cases, due to the smaller amount of separation work done in the recycle case and the lack of ambient air to be heated. These same trends hold for all εGG values, and for all the

ΔT and pRED values considered. The low separation temperature gives the highest efficiency for the idealized value of εGG = 1, when the sensible heating terms are zero, but for all other εGG values,

O2 the high TSEP value does as well as or better than lower values. These results are shown in the

Supplemental Material of Ref [41].

2.4.5. Impact of Water/Hydrogen Separation Temperature

The temperature at which the H2 product is separated from excess water has been investigated by varying the separation temperature from 373 K to 1,000 K, and this shown in

Figure 2.11. For perfect gas heat recovery and a low reduction partial pressure (Figure 2.11a), the

H2 maximum efficiency decreases and ΔTopt increases as the water/H2 separation temperature (TSEP ) increases. However, at εGG < 1, the maximum efficiency increases for higher separation temperatures. For the range of separation temperatures considered, the change in the maximum

ηSTH and ΔTopt is small compared to the effects of gas and solid heat recuperation. A higher reduction partial pressure (Figure 2.11b) shows an increased change in ηSTH and ΔTopt at lower εGG values, and no discernable effect at εGG = 1. For reference, the nwh values used in these calculations are given in the Supplemental Material of Ref [41].

a) b)

H2 Figure 2.11: Solar-to-hydrogen efficiency for ceria with TSEP = 373 K – 1,000 K for a reduction partial pressure of a) 0.1 Pa and b) 1,000 Pa. All cases have a reduction temperature of 1,800 K and a solid heat recuperation effectiveness of 50%. H2 The inconsistent effect of TSEP on efficiency arises from varying dominance of the auxiliary terms, which are shown in the Supplemental Material of Ref [41]. When εGG < 1, the energy required to sensibly heat the water and inert gas stream (QH2O and QIH) have a large impact on system efficiency, as would be expected. This leads to a change in QAUX, and thus a ηSTH increase and change in ΔTopt. However, when εGG = 1 the value of QH2O is essentially unchanged

H2 by varying Tsep due to the perfect heat recuperation. There is a very small change due to the heat capacity difference between the 1 mole of H2O that is heated and becomes 1 mole of H2 which is

H2 cooled, but this has a negligible effect of system efficiency. Instead, QSEP increases with

H2 H2 increasing Tsep (directly proportional to QSEP ) leading to an efficiency decrease. This increase

H2 in QSEP does occur for the cases when εGG < 1, but the associated decrease in QH2O outweighs

H2 the increase in QSEP . Similar results to those in Figure 2.11 occur for when the inert gas flowrate is changed; these results are shown in the Supplemental Material of Ref [41].

2.4.6. Efficiency of Vacuum Pumping

The STH efficiency is shown in Figure 2.12 for the complete range of gas heat recuperation values for O2 removal by vacuum pumping at efficiencies of currently available pumps.

QAUX and QMOL terms, which has a detrimental impact on system efficiency. Specific energy terms for these conditions showing this can be seen in the Supplemental Material of Ref [25].

The pump efficiency is greatly affected by the reduction pressure, and so the impact of different reduction pressures on system efficiency is shown in Figure 2.13. The STH efficiency is very low at very low reduction pressures (< 10 Pa) due to both exponentially increasing pump

work (the ln(Pamb/pred) term in Equation (2.48)), and the exponentially decreasing pump efficiency

(the ηPUMP term in Equation (2.48)). The STH efficiencies are also low at higher reduction pressures (10,000 Pa), in this case due to the substantially lower reduction extent of the material.

A lower extent of reduction (smaller value of δRED) leads to a larger solids flowrate which increases sensible heating and mechanical work requirements, as well as a larger amount of excess steam required for oxidation, which leads to significantly higher sensible heating requirements. As such, the optimum reduction pressure for this type of system is somewhere near a more moderate (1,000

Pa) reduction pressure.

These analyses primarily consider the point at which the system efficiency (ηSTH) reaches the highest point; the ΔT value at which this is reached is denoted ΔTopt. For example, in Figure

2.13 the case for the 1,000 Pa reduction pressure has a lower system efficiency than the case for

100 Pa reduction pressure for ΔT values <150 K. However, the pRED = 1,000 Pa case reaches a maximum ηSTH ≈ 0.12 at ΔTopt ≈ 350 K, whereas the pRED = 100 Pa case only reaches ηSTH ≈ 0.05 at ΔTopt = 800 K. There are a number of reasons for the different shape of the various lines in

Figure 2.13; the model describes a complex non-linear system, but a few major trends are worth identifying. First, some of the cases for the higher reduction pressures (pRED ≥ 100 Pa) show a discontinuity at low ΔT values; this is due to the fact that at such a high partial pressure of oxygen, the thermodynamic cycle is simply not feasible with such a small temperature swing, no matter how much water is used for oxidation. This is described in more detail in previous works [5, 20].

The other major trend is the fact that two of the ηSTH lines in Figure 2.13 (pRED = 1,000 and pRED

= 10,000 Pa) first increases, reaches a peak, and then decreases whereas other lines monotonically increase ηSTH with increasing ΔT. This is due to the QAUX term reaching a zero value, and which is when the heat benefits outweigh the auxiliary heating requirements; this is described further in

other works in this series [20, 34]. It should be noted that for these (and subsequent) analyses, the reduction pressures being considered are order-of-magnitude estimates, and are not optimized.

Figure 2.13: ηSTH values for various ΔT values and various reduction pressures (pRED). All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5. In addition to comparing real pump data for vacuum reduction with other methods of reduction below, we also examined the effect of the pump efficiency itself. This is done not because of any preference to any specific pumping technology, but rather to see the full range of vacuum pump efficiency values and the impact that these will have on the overall STH efficiency.

Singh et al. [11] have analyzed the effect of vacuum pump efficiencies before, but these results are included here both as a direct comparison to other results in this work, as well as the fact that these results show the effect of different vacuum pump efficiency values on the entire system, rather than just the reduction reaction terms. The results of these were done for the highest (10,000

Pa) and lowest (0.1 Pa) reduction pressures considered, and the results are shown in Figure 2.14.

a) b)

Figure 2.14: ηSTH values for various vacuum pump efficiency values for 1) pRED = 0.1 Pa and b) pRED = 10,000 Pa. All calculations were done for a reduction temperature of 1,800 K, a gas heat recuperation effectiveness of 0.9, and a solid heat recuperation effectiveness of 0.5. Unsurprisingly, order-of-magnitude increases to vacuum pump efficiency at low efficiency values have the greatest positive impact on STH efficiencies, relative to the smaller increases at higher pump efficiency values. If vacuum pumps could operate at a low pressure with an efficiency of ~10% or better, vacuum pumping has the potential to facilitate highly efficient solar thermochemical hydrogen production. Efficiency gains above this level have more moderate impacts on system efficiency: the difference in optimum STH efficiency values between ηPUMP =

0.1 and ηPUMP = 1.0 is <5%. This is because after a critical level of pump efficiency, the vacuum pump work required becomes small relative to other heating requirements. For higher reduction pressures, these impacts are generally the same, but much more muted; for example, at pRED =

10,000 Pa the only significant change in optimal ηSTH values occurred between ηpump = 0.001 and

0.01. At ηpump values ≥ 0.01, the ηSTH values are relatively unchanged. This is because the thermodynamic pumping work for compressing the oxygen is less, meaning that the pump has less of an impact on the final value of the QPUMP term. The STH efficiency values are much lower (ηSTH

< 10%) for the high pressure case because of the lower extent of reduction achieved.

2.4.7. Efficiency of Cascade Pressure Reduction

The STH efficiency results for a cascade pressure reduction system with five chambers using currently available pump efficiencies are shown in Figure 2.15 for a variety of gas heat recuperation values. As can be seen, the system efficiencies are still quite low. This is because effective pump efficiency value for five chambers is still very low (~0.00025%). Thus, the pump work dominates the other heat requirements for the rest of the system. This is shown explicitly in the Supplemental Material of Ref [34].

Figure 2.15: System efficiency (ηSTH) for a cascade pressure reduction system with 5 chambers. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 1 Pa, and a solid heat recuperation effectiveness (εSS) of 0.5. The STH efficiency for various reduction pressures is shown in Figure 2.16. As can be seen, the same general trends are followed that were shown for a single vacuum chamber system

(Figure 2.13). At very low pressures, the STH efficiency is low due to the very high amounts of pump work required. At very high pressures, the active material is not reduced enough to be operated efficiently. However, at a moderate pressure (~100 Pa), the system will reach an optimum. This exact value is somewhat different for a system with five chambers than it is for a system with one chamber, due to the change in effective pump efficiency. It should be noted, though, that the five chamber case has higher STH efficiency values than the single chamber case.

Figure 2.16: System efficiency (ηSTH) for a cascade pressure reduction system with five chambers. All of these values are calculated for ceria at a reduction temperature (TRED) of 1800 K, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5. Effective pump efficiencies for various numbers of chambers are shown in Figure 2.17(a).

As can be seen, the effective pump efficiency increases moderately for systems with more chambers, but not necessarily by orders of magnitude. This is because different chambers are pumped at different pressures and thus at different efficiencies. However, even for the limiting case of a system with an infinite number of chambers, the effective pump efficiency is still relatively low (between 0.00025% and 0.00045%). This still very low effective vacuum pump efficiency at a low reduction pressure leads to still very low system efficiency values, which are shown in Figure 2.17(b).

a) b)

eff Figure 2.17: a) effective pump efficiency (ηPUMP ) and b) solar to hydrogen efficiency (ηSTH) values for cascade pressure reduction systems with various numbers of chambers. All the

calculations here are done for ceria at a reduction temperature (TRED) of 1,800 K and a final reduction pressure (pRED) of 1 Pa. Though the improvement at very low reduction pressures is not very high due to the still low effective vacuum pump efficiency values, Figure 2.16 shows that more moderate pressures have higher system efficiency values. Additionally, the same moderate pressures have the highest improvement to effective pump efficiency (shown in the Supplemental Material of Ref [25]).

These two effects can therefore be combined to see the maximum system improvement possible on the system efficiency (within the values considered here), and this is shown in Figure 2.18 for pRED = 100 Pa. The largest increases in system efficiency occur for 2 and 5 cascade chambers, while additional chambers above this value have much less of an effect on the maximum ηSTH.

This reflects the improvement to the effective pump efficiency on the reduction pressure that is most sensitive to this change, as discussed above. Conversely, the system efficiency is almost entirely unchanged for higher reduction pressures due to the negligible impact that vacuum pump work has on system efficiency, even for an infinite number of cascade chambers (shown in the

Supplemental Material of Ref [25]). All of the system efficiency values calculated for pressures and cascade configurations calculated are given in the Supplemental Material of Ref [25] for comparison.

a) b)

eff Figure 2.18: a) effective pump efficiency (ηPUMP ) and b) solar to hydrogen efficiency (ηSTH) values for cascade pressure reduction systems with various numbers of chambers. All the calculations here are done for ceria at a reduction temperature (TRED) of 1,800 K and a final reduction pressure (pRED) of 100 Pa. 2.4.8. Efficiency of Inert Sweep Gas Reduction

The system efficiency using inert gas reduction for various gas heat recuperation values is shown in Figure 2.19, and the specific heat requirements are shown the Supplemental Material of

Ref [25]. The decreasing value of inert gas heating requirements (QIH) and inert gas/oxygen

O2 separation work (QSEP ) with increasing ΔT are due to the decreasing inert gas flowrate (nio) required, and lead to the increasing ηSTH at low values of ΔT. At higher values of ΔT (>200 K), the inert gas flowrate has decreased to zero, meaning that further increases to system efficiency from decreasing inert gas requirements are not possible; instead, the system efficiency decreases due to the increasing sensible heating of the reactive solids (QSH).

Figure 2.19: System efficiency (ηSTH) for inert gas sweep system. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, and a solid heat recuperation effectiveness (εSS) of 0.5. The use of an inert gas sweep is especially useful to achieve very low oxygen partial pressures, which in turn gives high STH efficiency values. The system efficiency tends to decrease at higher oxygen partial pressures for reduction. This is because as the solid achieves a greater extent of reduction, it is oxidized more easily. This means that less excess steam is required to

oxidize the material, and less overall active redox material is needed to produce the same amount of hydrogen. As the reduction oxygen partial pressure increases, the inert gas requirements decreases, but the benefits from greater extents of reduction also decrease. This is shown in Figure

2.20.

Figure 2.20: System efficiency (ηSTH) for inert gas sweep system with different reduction pressures. All of these values are calculated for ceria at a reduction temperature (TRED) of 1,800 K, a reduction pressure (pRED) of 0.1 Pa, a gas heat recuperation effectiveness (εGG) of 0.9, and a solid heat recuperation effectiveness (εSS) of 0.5. The STH efficiency shown above depends greatly on the inert gas flowrate, as this impacts the inert gas sensible heating and cooling terms as well as the inert/oxygen separation work. As such, the inert gas flowrates calculated for the full range of reduction oxygen partial pressures considered are shown in Figure 2.21. The flowrates for the very low (≤ 10 Pa) oxygen partial pressures are similar, but the behavior changes for reduction partial pressures of 100 Pa and above.

This is because at very low ΔT values, the reduction and oxygen reactions simply do not proceed because the reduced material is no longer able to split water. This is consistent with previous studies which have shown that ceria is unable to undergo a reduction/oxidation cycle at low ΔT values and high oxygen partial pressures [5, 20]. The exact ΔT value at which the cycle is thermodynamically viable requires an infinite inert gas flowrate because Δδ = 0, meaning that there will be an infinite solids flowrate, an infinite oxygen flowrate, and thus an infinite inert gas

flowrate. The nio plots that include very large (>1,600) nio values are truncated in Figure 2.21 for ease of viewing. The effect of non-minimum inert gas flowrates (such as those calculated here) are considered in a separate analysis [41].

Figure 2.21: Inert gas flowrate ratio (nio) for use in an inert gas sweep for reduction at various reduction oxygen partial pressures. Calculations were done for ceria with a reduction temperature of 1,800 K. The effect on overall STH efficiency for a wide range of inert/oxygen separation efficiency values (0.001 to 1.0) are shown in Figure 2.22. The result of a lower oxygen separation efficiency is lower STH efficiency at low ΔT values due to the increase of oxygen/inert separation work. STH efficiency values at higher ΔT values are similar, due to the fact that the inert gas flowrate goes to zero (which is not realistic). However, the results in Figure 2.22 do indicate that an oxygen/inert separation efficiency of ~10% will give very high STH efficiencies for this active material and set of conditions. This is because the optimal efficiency and ΔT value are not at the same point as the lower efficiency curves; that is, it is not when nio ≈ 0. Thus, this shows that if an efficiency value of ~10% can be achieved, the inert gas sweep option is likely to achieve the STH efficiencies necessary for commercialization of this process (for this set of conditions). Additionally, these results show that improving the inert/oxygen separation efficiency much above 10% will likely

have little additional impact on overall STH efficiency. It should also be noted that separation efficiencies below 10% also give high STH efficiencies, but these high STH efficiency values occur when nio is at or near zero. Thus, the actual STH efficiency values at lower inert/O2 separation efficiencies will depend on the case-specific inert gas ratio used.

Figure 2.22: ηSTH values for ceria at various inert/oxygen separation efficiency values. All calculations were done at a reduction temperature of 1,800 K, a reduction oxygen partial pressure of 0.1 Pa, a gas heat recuperation effectiveness of 0.9, and a solid heat recuperation effectiveness of 0.5. 2.5. Discussion of Model Assumptions

Thermodynamic modeling is useful to illustrate trends and gain an understanding about trade-offs in a potential solar thermochemical system, even if the actual performance numbers will be different for a real system. Many previous analyses have examined the process impacts of ceria, though few studies have included a direct comparison of ceria with another active material. We will discuss the implications and tradeoffs within the system based on a new material. Then, the impact of material-specific oxidation kinetics for the two active materials, ceria and ferrite/zirconia, will be discussed. Next, the practicality of achieving the minimum inert gas flowrate for a perfectly counter-flow configuration will be discussed. Then the implications of a high inert/oxygen separation temperature will be highlighted. The possibility of achieving different

levels of gas and solid heat recuperation will next be discussed, and the impact on the system efficiency will be emphasized. The feasibility of increasing the water/hydrogen separation temperature will then be briefly discussed. Limitations on vacuum pump efficiency improvements will be discussed next, both in the context of direct and cascade pressure reduction. Additionally, the critical assumptions for the recycled inert gas sweep will be discussed, including the flow rates and separation efficiency.

2.5.1. Material Specific Effects on Thermodynamic Efficiency

Section 2.4.1 shows the thermodynamic efficiency of both ceria and a ferrite zirconia composite, which illustrates interactions between material-specific factors and the resulting impact on efficiency. By way of example, Scheffe et al. have suggested that a higher heat capacity active material may have a lower efficiency [46]. The results here show that the impact of sensibly heating the reactive solids for a material with a higher heat capacity might actually have a smaller impact on system efficiency if those same materials have a higher hydrogen productivity per mass of solids. Additionally, the results here illustrate the fact that different materials can have wildly different optimal operating conditions, and that these differences come from many different aspects of the process. There is no one parameter driving the optimal thermodynamic efficiency or set of optimal operating conditions; it is the result of a complex non-linear system that is difficult to predict a priori. This shows that this type of modeling is generally useful, especially for the impact of new assumptions and materials on the system.

Even though the efficiency values using the ferrite/zirconia composite are generally much higher than for ceria, this is not necessarily an endorsement or suggestion that this material is an ideal candidate for solar thermochemical water splitting. There is no experimental validation of the thermodynamic predictions of this material done here, and thus no sort of justification that this

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material could actually perform as described. Additionally, this material may suffer from kinetic limitations; this is described more below. Furthermore, it should be stated clearly that the authors do not necessarily consider either of these materials to be ideal candidates for STWS in an optimal process. These materials were used due to the availability of thermodynamic and kinetic information. There are many novel materials which may provide much better performance [47-50] but for which little or no thermodynamic and kinetic data are available for a variety of operational conditions. As with all thermodynamic models of this type, this comparison is still useful to illustrate trends and identify specific tradeoffs within the system.

2.5.2. Impact of Reaction Rates on Process Design

Section 2.4.1 illustrates the impact of kinetics on efficiency. Only the oxidation reaction rates are considered, not the reduction kinetics. The analysis assumes that the same reaction time will be used for all oxidation temperatures, when in a well-designed system the space-time would increase for the oxidation reactor to overcome these affects (albeit at some cost). However, this analysis illustrates the impact of reaction rates on process design when selecting the optimal operating conditions. Thermodynamic equilibrium is used both here and in other works to estimate material productivity and cycle efficiency, and these calculations can predict fairly large temperature swings between reduction and oxidation. The current attempt to include kinetics does quantify the drastic effect that reaction rates can have. Near-isothermal behavior thus becomes more optimal for materials with slower kinetics, and especially with moderate to high gas heat recuperation. The difference in ΔTopt for thermodynamic equilibrium vs those calculated with the kinetic discounting is drastic for the ferrite/zirconia. Additionally, this analysis may be optimistic for ceria, as others have found ceria oxidation to be kinetically limited in some conditions [18].

Furthermore, this analysis points to the benefits of a flexible reactor design which can

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accommodate different conditions for reduction and oxidation as well as different flow rates between the two reaction zones. This would allow for reaction rates to be addressed in each reaction zone independently rather than running one or both reactions to less conversion.

2.5.3. Practicality of Counter-Flow Inert Gas Sweep

A major assumption of this work is the perfect counter-flow arrangement of inert gas and reactive solids. This leads to a much more promising operating efficiency than if perfect gas mixing were assumed in the reduction reaction, and this has also been shown in previous analyses [6, 24,

26]. However, this type of perfect flow assumption is unlikely to be fully realized in practice [18,

26]. Regardless of how it is implemented, mass transport effects will inevitably require additional inert gas to make up for the loss in purity. Thus, we have shown two examples of efficiency calculations done with excess inert gas in order to examine the effects, and others have examined this flowrate also [6, 26]. Instead, this work focuses on a concept which has not been previously considered for the reduction sweep gas of a high temperature separation recycle stream. Like most of these calculations, the results presented here are only for a specific set of conditions and materials; the final efficiency will ultimately depend on detailed materials properties and reactor designs. However, this does point to the usefulness of a reactor design that can minimize back- mixing in an inert gas sweep reduction reactor. Designs can be taken from catalytic cracking and chemical looping reactors, which utilize gas-solid contact for reaction cycles [51].

2.5.4. Implications for High Inert/Oxygen Separation Temperature

There are multiple established and mature technologies for separating O2 from air or another inert gas, but these technologies typically operate at ambient or sub-ambient temperatures.

This can be problematic for high temperature systems, as the inert gas must be sensibly heated up to the reduction temperature. Heat recuperation is certainly possible, even to fairly high levels, but

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the sensible energy load is still significant. Thermodynamically, more separation work is expected at high temperature, but this is offset by the benefit gained from less sensible heating. Some other authors [5, 24] have suggested recycling of the inert gas in the past, and the results here are consistent with their arguments that this is more efficient than continual inert production. The results here show that an additional benefit can be gained by performing the recycle purification at high temperature. Indeed, this benefit is even more pronounced when the inert gas flowrate is higher, such as when the idealized minimum is not achievable. Section 2.4.4 above and the

Supplemental Material of Ref [41] show that the system efficiency at ΔTopt is as good as or better for higher separation temperatures for any εGG < 1. This difference is more pronounced for the case in which a minimum nio value of 100 is imposed. This suggests that the high temperature separation is more beneficial for systems that have lower gas heat recuperation or higher inert gas flowrates.

Indeed, results in Supplemental Material of Ref [41] indicate that even higher system efficiency

O2 values can be achievable if the TSEP is raised above 1,223 K. High temperature electrochemical cells have been operated at 1,673 K [52], and so it is not unreasonable to consider further increasing the temperature for O2 separation. However, the best case will ultimately depend on the attainable values of heat recuperation and separation efficiency, and especially of the economics of each potential option.

2.5.5. Practicality of Gas and Solid Heat Recuperation

Gas heat recuperation is especially important for isothermal (ΔT = 0) or near-isothermal operation, in which the reduction and oxidation temperatures are the same or very close in value, while solid heat recuperation is more important for larger temperature swings. Small temperature swings require more excess steam for oxidation [20] and more inert gas for reduction than operation with larger temperature swings [5, 6]. As the gas heat recuperation increases, less energy

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is required to heat inert gas and excess steam, and the energy required to heat the solids from the oxidation to the reduction temperature becomes more significant. Thus, the energetic cost of heating the solids as they move between the reduction and oxidation temperatures (increasing ΔT) can outweigh the benefit gained by increasing the Δδ. It is difficult to know what a reasonable value is for the gas heat recuperation effectiveness. An infinite amount of heat transfer area would be required for the heat exchanger effectiveness to be 100%. Thus, the results that show a εGG =

1.0 are meant to illustrate a bound only, not to show what is realistically achievable. While an effectiveness of 100% is not physically realizable, it is also un-realistic practically to limit heat effectiveness values to significantly below 100% based on temperature limitations of metal-based heat exchangers. Ceramic materials can operate at temperatures >1,273 K, even though these materials are not ideal for heat exchange. There are no physical or thermodynamic limitations to the highly effective recuperation of heat from these gaseous streams. Some different concepts have been suggested [21, 53], and some commercial products already transfer heat at high (>1273 K) temperatures (e.g., [54]). The trade-off between highly effective heat recuperation and cost will likely be decided on a case-by-case basis, and will certainly change as the price of heat exchanger materials changes.

This analysis assumes that in any case where QAUX > 0, solid heat recuperation does not matter because the un-recuperated sensible heat in the solids is utilized for other heating loads elsewhere in the system. This can easily occur at lower values of ΔT, where the higher steam flow cools the solids and the steam is concomitantly heated to the desired oxidation temperature. The un-recuperated sensible heat in the solids is also easily usable in systems where the material has a low oxidation enthalpy change and therefore cannot sufficiently heat steam by the exothermic oxidation reaction, meaning that the system has even more use for the un-recuperated solids heat.

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However, once one moves away from heating steam, it becomes less clear exactly how this excess heat will be utilized. This argument is also applicable to the un-recuperated sensible heat form the recycled inert gas stream (QIC), as this heat is also at a usefully high temperature. It could be converted into electricity to counter some of the other energy demands within the system, but this will not necessarily be highly efficiency. The heat somehow needs to be collected (probably through some intermediate which will introduce a loss not considered here), converted (at some efficiency loss) to electricity, then transformed (again at an efficiency loss) to mechanical work.

The latter two factors are accounted for by the lumped efficiency terms (ηpump, ηmech, etc.) in the equations above, but this implicitly assumes that the heat collection can occur in the first place. In the current model, many heat loads and benefits are lumped together without specifying exactly how the heat will be recuperated and used in the system. More detailed, reactor- and process- specific analyses are required to show that this type of optimistic heat recovery is achievable.

2.5.6. Implications for High Hydrogen/Water Separation Temperatures

Raising the temperature at which the hydrogen product is separated from excess steam

H2 (TSEP ) has been suggested before as a way to increase the efficiency of the process[1]. This makes sense in that the large amount of excess steam would not need to be heated and cooled repeatedly with imperfect gas heat recuperation. Just raising the separation temperature to above 373 K so that the separation occurs without having to condense (and then re-vaporize) the excess steam should have a significant impact, as the heat of vaporization alone is 45% of the total energy required to heat liquid water at 295 K to steam at 1,473 K. This effect is shown in Section 2.4.3 for non-perfect gas heat recuperation, as the efficiency of the process goes up when the excess steam does not need to re-condense repeatedly.

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H2 However, the impact of TSEP on the maximum STH efficiency and the optimum ΔT value

H2 is not very large. The impact of a higher TSEP is highest when nwh is large (low ΔT values), meaning that this gaseous separation could be especially useful for near-isothermal operation. The practicality of raising the separation temperature must also be considered. Condensation is conceptually simple and well established in chemical processes, though it comes with its own set of challenges such as multi-phase operation and high heats of vaporization/condensation. Other hydrogen/water separation methods like a membrane or electrochemical cell could be implemented, but these would require additional energy to operate. This is accounted for in the

H2 current model by the QSEP term, which finds the thermodynamic energy required to separate the streams and inflates it by some efficiency. This could be the energy required for operating a condenser, membrane, or cell. However, current non-condensation technologies are limited to temperatures of 773 K [55, 56], less than the maximum of the current analysis. New technologies may provide a path to a higher separation temperature, but these other methods must be weighed against the energetic requirements to operate them.

2.5.7. Feasibility of High Vacuum Pump Efficiency

It is difficult to determine what a realistic value for a vacuum pump efficiency will be. Two factors affect this, solar to electric efficiency and electric to pump work efficiency. As solar-to- electric efficiency values rise due to improvements with photovoltaics, concentrated solar thermal power cycles, or other methods, the overall solar-to-pump work efficiency will rise. However, even the most optimistic performance targets (which have not been demonstrated) for concentrated solar power are for a solar-to-electric efficiency of 50% [57], not that much higher than the currently assumed value of 40%. As such, it is of more interest to consider possible improvements to the electric-to-work efficiency for vacuum pumps.

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Current technologies for near-atmospheric compressors are between ~70-80% efficient [6,

58]. However, vacuum pumps face challenges because of the inherently low mass flows associated with vacuum pumping, which yield energy efficiencies of a single-digit percent or less [59]. This means that pump and compressor efficiency values that can be assumed from other technologies can not necessarily be applied to low pressures. That said, we know of no fundamental limit on vacuum pump efficiency for very low pressures. As such, we do not limit the possibility of higher vacuum pump efficiency values, and an increase to this value will certainly have an impact on

STH efficiency; this is shown in Section 2.4.6 and reflected in many other works [8, 20, 24, 26].

It is likely that future advancements in vacuum pump technologies will increase vacuum pump efficiency values, possibly driven by economic market forces based on an increase in demand for high-efficiency, high-vacuum pumps. However, vacuum pumps have fewer physical gas molecules to pump, making them inherently less efficient at lower pressures. Thus, it is not likely that these increases will lead to improvements in vacuum pump efficiency of multiple orders of magnitude which would be necessary to match the efficiencies of efficient inert gas use.

2.5.8. Practicality of Cascade Pressure Reduction

Some type of multi-pump cascading pressure system may be helpful for any type of large- scale solar thermochemical hydrogen production system that uses vacuum-driven reduction. The large scale and low pressures give rise to enormous volumetric flow rates, which a single pump is unlikely to be able to handle. However, while the gains in effective pump efficiency using multiple cascading pumps are significant, they are less than a single order of magnitude, similar to the results of Brendelberger and Sattler [29]. Even for an infinite number of chambers, the increase in effective pump efficiency does not have a large impact on system efficiency for low reduction pressures; this is consistent with the findings by Jarrett et al. [30]. However, this increase in

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effective pump efficiency does change the optimum system efficiency and optimal operating conditions at higher reduction pressures, showing that this does have an effect on overall system efficiency for some conditions.

The current analysis assumed that each vacuum pump in the cascading system would have the same volumetric flowrate. This is not necessarily the most efficient way to operate, and other strategies were not explicitly explored here. However, we do not foresee a drastic change in effective pump efficiency based on changing the volumetric flowrate distribution. This is because increasing the flowrate of pumps in the beginning of the cascade (at higher pressures) so that they do more of the pump work at a higher pump efficiency will mean that the chambers for those pumps will operate at a lower pressure, thus negating the gain in pump efficiency.

Another aspect not considered here is the economic impact of additional vacuum pumps.

Obviously, there are a minimum number of pumps required to achieve a certain pressure at a certain scale, because pumps can realistically only be so large. However, increasing the number of pumps beyond this minimum number will have significant economic impacts on any system. These economic impacts should be justified by cost savings elsewhere, or substantially higher efficiencies such that the long term lower operation costs off-set the higher initial costs. However, a moderate increase in efficiency that can be gained even with large numbers of pumps would likely not make this tradeoff economically worthwhile.

Finally, there is the practical challenge of implementing a system in which a series of distinct chambers hold very different pressures while also allowing solids flow between them. This is especially true for some of the very low pressures considered here. Any sort of moving packed bed between chambers will have a number of engineering challenges to overcome both to controllably move the particles between chambers and to maintain large pressure differences

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between chambers, especially since the entire configuration will operate at the very highest temperature in the thermochemical system (reduction temperature). Additionally, this type of cascade faces challenges similar to those for the single pressure vacuum case, in which the largest pressure differential occurs between the final chamber (NC) and the subsequent oxidation step. All of these large pressure differences will be difficult to implement with a continuously flowing solid material, especially at the extremely high temperatures under which the system operates. These challenges will be difficult to overcome, and it is not clear that the gain in system efficiency shown here would justify this difficult and expensive system.

2.5.9. Practicality of Inert Gas Sweep Reduction

There are a number of challenges to operate a process at any sort of scale at both very high temperatures and at very high or low pressures as discussed above. However, there are also challenges with removing O2 using an inert sweep gas; it requires a large amount of sensible heating and cooling for each cycle, it requires separation for each cycle, and a reactor leak would require the inert gas to be re-purified or replaced entirely. There are also practical issues with achieving the thermodynamic productivity shown here. Some issues are related to reaction rates and kinetics; these are discussed elsewhere [34]. Additionally, there are other issues with processing of the inert gas to achieve the very low oxygen partial pressures required for efficiency reduction. Two important issues are the ability of an oxygen separation technology to rapidly achieve the inert gas purity assumed in this work, as well as the ability of the inert gas and solid flows to achieve (or at least approach) the counter-flow arrangement assumed here. Brendelberger et al. has found that the latter of these issues may be difficult [26]. These will need to be explored and demonstrated experimentally.

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The inert/oxygen separation energy calculated in this work did not make any assumptions on a specific technology, but rather relied on different entropies of mixing for the different streams into and out of the separator. Thus, it does not make any assumptions about the physical ability of a real separator to achieve a particular inert gas purity. Additionally, it does not make any assumptions about the time it would take a separator to achieve a given purity. This will have a large impact on the practicality and system efficiency of an inert gas sweep system. There are currently some commercial products that have demonstrated oxygen purities at temperatures of interest [60, 61], but more information is needed about the limitations of these technologies and the energetic efficiencies they are able to achieve. Further development is likely required to improve upon existing technologies to make this type of operation truly preferable.

2.6. Conclusions

A thermodynamic model has been developed for investigating various aspects of a solar thermochemical water splitting process. This model builds upon previous models described in the literature, and has expanded upon the prior state of the art by including a new material and kinetic effects. It uses a thermodynamic cycle of solid reactive material flowing between reduction and oxidation chambers. The model calculates flow rates, energy requirements, and energy benefits to produce one mole of hydrogen and uses these values to calculate the solar flux required to generate that amount of heat. The solar-to-hydrogen efficiency is then calculated for a variety of conditions and for two different materials.

In general, the ferrite/zirconia composite considered here displayed higher thermodynamic efficiency values than ceria, primarily due to the increased hydrogen production capacity of the material. Generally, oxidation kinetic limitations to ceria are minimal except at very high ΔT values. The kinetic limitations for ferrite/zirconia are larger, causing large decreases in efficiency

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at even moderate ΔT values. This means that even for conditions that appear to be optimal for a particular material when considering only thermodynamic end-points, a very different set of operating conditions may be more efficient when kinetics are considered.

Heat recuperation is important for any efficient solar thermochemical water splitting process. Reasonable values for both gas and solid heat recuperation are not known in general, and will depend on case-specific factors such as materials availability and cost. In general, the optimal

STH efficiency increases with both gas and solid heat recuperation. The optimal ΔT to operate the process increases with increasing solid heat recuperation but decreases with increasing gas heat recuperation. The impact of gas heat recuperation on the optimal STH efficiency is somewhat muted, based on the current assumptions of a minimum inert gas flow rate and a low oxygen partial pressure; the impact will increase if either of these assumptions is relaxed. The impact of solid heat recuperation on the optimal STH efficiency, on the other hand, is somewhat inflated, especially since this optimal condition typically exists at or near when the inert gas flow rate goes to zero. A zero flowrate is not physically realistic; a higher inert gas flowrate will shift the balance of importance more towards gas heat recuperation rather than solids heat recuperation.

A high separation temperature for the recycled inert gas has been shown to be beneficial, especially for cases of lower gas heat recuperation and increased inert gas flowrates. The thermodynamic increase in separation work is offset by greatly reduced sensible energy requirements to reach the very high reduction temperature. Furthermore, the high separation temperature means that the un-recuperated sensible heat from the inert gas stream is still at a useful temperature, and can be used elsewhere in the system. However, this high temperature separation is only beneficial for a recycled inert gas system, not an open loop continuous production of N2 from ambient air.

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The impact of a higher water/hydrogen separation temperature was examined as a way to decrease the importance of high temperature gas heat recuperation, even though it is not immediately clear how this would be implemented in practice. An increase in this separation temperature has the largest impact on efficiency for smaller ΔT values and moderate values of gas heat recuperation, though the increase to the maximum ηSTH and shift of ΔTopt to smaller values are both fairly small. This relatively small increase with the current configuration is unlikely to be worth the increased system complexity and parasitic losses.

In addition to specific insights about particular conditions or parameters, more general lessons can also be gained from this type of efficiency analysis. Change in a particular parameter can shift the ideal operating conditions drastically, meaning that a flexible reactor design is important. A reactor that allows for different operating temperatures, pressures, and flow rates between the reduction and oxidation reactions would allow for the system to always operate at the most efficient conditions, instead of being tied to a particular method or mode of operation.

Materials research should be directed towards developing high temperature refractory steam/steam heat exchangers which can operate above 1273K, thus providing for an increase in εGG; this becomes more important if a non-ideal minimum inert gas flowrate is considered.

Different methods for achieving a low oxygen partial pressure have been compared, and the impact that each of these methods has on the system efficiency as a whole. Lower vacuum pump efficiencies are incorporated into the model, and suggest that currently available vacuum pumping results in low STH efficiencies for thermodynamic pump efficiencies of <10%.

Meanwhile, a novel recycled inert gas sweep with a high temperature separation step maintains relatively high STH efficiencies over a range of gas purification efficiencies and flow rates.

Thermodynamic separation efficiency values of around 10% are also likely to result in high system

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efficiency capability. Ultimately, the most efficient mode of operation will depend on the efficiency at which the inert gas and oxygen can be separated (in the inert gas case) and the efficiency at which the vacuum pumps can operate (in the vacuum case). Further, the efficiency of a cascade pressure reduction has been examined, which uses multiple vacuum pumps operating at different pressures to achieve high levels of vacuum. It was found that cascade pressure reduction does increase the effective pump efficiency relative to a single vacuum pump, but this increase

(typically a maximum of 5X) is not large enough to cause the system efficiency to overtake the inert gas case. These results suggest that future research should focus on the development of high temperature ionic transport membranes, or other technologies with similar effects, for efficiently separating O2 from an O2-containing inert gas sweep stream where the inert gas is recycled to the reduction reactor without having been substantially cooled. Alternately, additional efforts to achieve a thermodynamic vacuum pump efficiency of 10% would drastically increase the possibility of that approach to thermochemical reduction.

2.7. References

1. Lange, M., et al., T–S diagram efficiency analysis of two-step thermochemical cycles for solar water splitting under various process conditions. Energy, 2014. 67: p. 298-308.

2. Lange, M., et al., Entropy Analysis of Solar Two-Step Thermochemical Cycles for Water and Carbon Dioxide Splitting. Entropy, 2016. 18(1): p. 24.

3. Muhich, C.L., et al., A review and perspective of efficient hydrogen generation via solar thermal water splitting. Wiley Interdisciplinary Reviews: Energy and Environment, 2016. 5(3): p. 261-287.

4. Siegel, N.P., et al., Factors Affecting the Efficiency of Solar Driven Metal Oxide Thermochemical Cycles. Industrial & Engineering Chemistry Research, 2013. 52(9): p. 3276- 3286.

5. Bader, R., et al., Thermodynamic Analysis of Isothermal Redox Cycling of Ceria for Solar Fuel Production. Energy & Fuels, 2013. 27(9): p. 5533-5544.

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6. Krenzke, P.T. and J.H. Davidson, On the Efficiency of Solar H2 and CO Production via the Thermochemical Cerium Oxide Redox Cycle: The Option of Inert-Swept Reduction. Energy & Fuels, 2015. 29(2): p. 1045-1054.

7. Lapp, J., J.H. Davidson, and W. Lipiński, Efficiency of two-step solar thermochemical nonstoichiometric redox cycles with heat recovery. Energy, 2012. 37(1): p. 591-600.

8. Bulfin, B., et al., Thermodynamics of CeO2 Thermochemical Fuel Production. Energy & Fuels, 2015. 29(2): p. 1001-1009.

9. Diver, R.B., et al., Solar Thermochemical Water-Splitting Ferrite-Cycle Heat Engines. Journal of Solar Energy Engineering, 2008. 130(4): p. 041001-041001-8.

10. Mallapragada, D.S. and R. Agrawal, Limiting and achievable efficiencies for solar thermal hydrogen production. International Journal of Hydrogen Energy, 2014. 39(1): p. 62-75.

11. Singh, A.K., et al., Thermal Reduction of Iron Oxide under Reduced Pressure and Implications on Thermal Conversion Efficiency for Solar Thermochemical Fuel Production. Industrial & Engineering Chemistry Research, 2015. 54(26): p. 6793-6803.

12. Furler, P., et al., Thermochemical CO2 splitting via redox cycling of ceria reticulated foam structures with dual-scale porosities. Physical Chemistry Chemical Physics, 2014. 16(22): p. 10503-10511.

13. Venstrom, L.J., et al., Efficient Splitting of CO2 in an Isothermal Redox Cycle Based on Ceria. Energy & Fuels, 2014. 28(4): p. 2732-2742.

14. Keene, D.J., J.H. Davidson, and W. Lipiński, A Model of Transient Heat and Mass Transfer in a Heterogeneous Medium of Ceria Undergoing Nonstoichiometric Reduction. Journal of Heat Transfer, 2013. 135(5): p. 052701-052701-9.

15. Yuan, C., et al., A new solar fuels reactor concept based on a liquid metal heat transfer fluid: Reactor design and efficiency estimation. Solar Energy, 2015. 122: p. 547-561.

16. Lange, M., et al., Efficiency assessment of a two-step thermochemical water-splitting process based on a dynamic process model. International Journal of Hydrogen Energy, 2015. 40(36): p. 12108-12119.

17. Levêque, G. and S. Abanades, Investigation of thermal and carbothermal reduction of volatile oxides (ZnO, SnO2, GeO2, and MgO) via solar-driven vacuum thermogravimetry for thermochemical production of solar fuels. Thermochimica Acta, 2015. 605: p. 86-94.

18. Venstrom, L.J., et al., Applicability of an Equilibrium Model To Predict the Conversion of CO2 to CO via the Reduction and Oxidation of a Fixed Bed of Cerium Dioxide. Energy & Fuels, 2015. 29(12): p. 8168-8177.

19. Muhich, C.L., et al., Efficient Generation of H2 by Splitting Water with an Isothermal Redox Cycle. Science, 2013. 341(6145): p. 540-542.

114

20. Ermanoski, I., J.E. Miller, and M.D. Allendorf, Efficiency maximization in solar- thermochemical fuel production: challenging the concept of isothermal water splitting. Physical Chemistry Chemical Physics, 2014. 16(18): p. 8418-8427.

21. Bala Chandran, R., R.M. De Smith, and J.H. Davidson, Model of an integrated solar thermochemical reactor/reticulated ceramic foam heat exchanger for gas-phase heat recovery. International Journal of Heat and Mass Transfer, 2015. 81: p. 404-414.

22. Smestad, G.P. and A. Steinfeld, Review: Photochemical and Thermochemical Production of Solar Fuels from H2O and CO2 Using Metal Oxide Catalysts. Industrial & Engineering Chemistry Research, 2012. 51(37): p. 11828-11840.

23. Agrafiotis, C., M. Roeb, and C. Sattler, A review on solar thermal syngas production via redox pair-based water/carbon dioxide splitting thermochemical cycles. Renewable and Sustainable Energy Reviews, 2015. 42: p. 254-285.

24. Lin, M. and S. Haussener, Solar fuel processing efficiency for ceria redox cycling using alternative oxygen partial pressure reduction methods. Energy, 2015. 88: p. 667-679.

25. Ehrhart, B.D., et al., System efficiency for two-step metal oxide solar thermochemical hydrogen production – Part 3: Various methods for achieving low oxygen partial pressures in the reduction reaction. International Journal of Hydrogen Energy, 2016. 41(44): p. 19904- 19914.

26. Brendelberger, S., et al., Counter flow sweep gas demand for the ceria redox cycle. Solar Energy, 2015. 122: p. 1011-1022.

27. Ermanoski, I., N.P. Siegel, and E.B. Stechel, A New Reactor Concept for Efficient Solar- Thermochemical Fuel Production. Journal of Solar Energy Engineering, 2013. 135(3): p. 031002-031002-10.

28. Ermanoski, I., Cascading pressure thermal reduction for efficient solar fuel production. International Journal of Hydrogen Energy, 2014. 39(25): p. 13114-13117.

29. Brendelberger, S. and C. Sattler, Concept analysis of an indirect particle-based redox process for solar-driven H2O/CO2 splitting. Solar Energy, 2015. 113: p. 158-170.

30. Jarrett, C., et al., Critical limitations on the efficiency of two-step thermochemical cycles. Solar Energy, 2016. 123: p. 57-73.

31. Xu, R. and T.F. Wiesner, Conceptual design of a two-step solar hydrogen thermochemical cycle with thermal storage in a reaction intermediate. International Journal of Hydrogen Energy, 2014. 39(24): p. 12457-12471.

32. Panlener, R.J., R.N. Blumenthal, and J.E. Garnier, A Thermodynamic Study of Nonstoichiometric Cerium Dioxide. Journal of Physics and Chemistry of Solids, 1975. 36: p. 1213-1222.

115

33. Bale, C.W., et al., FactSage thermochemical software and databases — recent developments. Calphad, 2009. 33(2): p. 295-311.

34. Ehrhart, B.D., et al., System efficiency for two-step metal oxide solar thermochemical hydrogen production – Part 1: Thermodynamic model and impact of oxidation kinetics. International Journal of Hydrogen Energy, 2016. 41(44): p. 19881-19893.

35. Hydrogen Analysis Resource Center: Lower and Higher Heating Values of Hydrogen and Fuels. Available from: http://hydrogen.pnl.gov/hydrogen-data/lower-and-higher-heating- values-hydrogen-and-other-fuels.

36. Falcone, P.K., Handbook for solar central receiver design. 1986, Sandia National Laboratories: Report No. SAND86-8009 Livermore, CA.

37. Ehrhart, B.D. and D.D. Gill, Evaluation of Annual Efficiencies of High Temperature Central Receiver Concentrated Solar Power Plants With Thermal Energy Storage. 2013, Sandia National Laboratories: Report No. SAND2013-5493 Albuquerque, NM.

38. Bulfin, B., et al., Analytical Model of CeO2 Oxidation and Reduction. The Journal of Physical Chemistry C, 2013. 117(46): p. 24129-24137.

39. Domalski, E.S. and E.D. Hearing, Condensed Phase Heat Capacity Data, in NIST Chemistry WebBook, Standard Reference Database Number 69, P.J. Linstrom and W.G. Mallard, Editors, National Institute of Standards and Technology: Gaithersburg, MD.

40. Sander, R., Henry's Law Constants, in NIST Chemistry WebBook, Standard Reference Database Number 69, P.J. Linstrom and W.G. Mallard, Editors, National Institute of Standards and Technology: Gaithersburg, MD.

41. Ehrhart, B.D., et al., System efficiency for two-step metal oxide solar thermochemical hydrogen production – Part 2: Impact of gas heat recuperation and separation temperatures. International Journal of Hydrogen Energy, 2016. 41(44): p. 19894-19903.

42. Arifin, D., Study of redox reactions to split water and carbon dioxide, in Chemical and Biological Engineering. 2013, University of Colorado: Boulder.

43. Scheffe, J.R., et al., Kinetics and mechanism of solar-thermochemical H2 production by oxidation of a cobalt ferrite-zirconia composite. Energy & Environmental Science, 2013. 6(3): p. 963-973.

44. Chueh, W.C. and S.M. Haile, A thermochemical study of ceria: exploiting an old material for new modes of energy conversion and CO2 mitigation. Philosophical transactions. Series A, Mathematical, physical, and engineering sciences, 2010. 368(1923): p. 3269-94.

45. Scheffe, J.R. and A. Steinfeld, Thermodynamic Analysis of Cerium-Based Oxides for Solar Thermochemical Fuel Production. Energy & Fuels, 2012. 26(3): p. 1928-1936.

116

46. Scheffe, J.R., D. Weibel, and A. Steinfeld, Lanthanum–Strontium–Manganese Perovskites as Redox Materials for Solar Thermochemical Splitting of H2O and CO2. Energy & Fuels, 2013. 27(8): p. 4250-4257.

47. McDaniel, A.H., et al., Sr- and Mn-doped LaAlO3-δ for solar thermochemical H2 and CO production. Energy & Environmental Science, 2013. 6: p. 2424-2428.

48. Takacs, M., J.R. Scheffe, and A. Steinfeld, Oxygen nonstoichiometry and thermodynamic characterization of Zr doped ceria in the 1573-1773 K temperature range. Physical Chemistry Chemical Physics, 2015. 17(12): p. 7813-7822.

49. Al-Shankiti, I., et al., Solar Thermal Hydrogen Production from Water over Modified CeO2 Materials. Topics in Catalysis, 2013. 56(12): p. 1129-1138.

50. Rao, C.N.R., et al., Solar photochemical and thermochemical splitting of water. Philosophical Transactions of the Royal Society of London A: Mathematical, Physical and Engineering Sciences, 2016. 374(2061).

51. Fan, L.-S., Chemical Looping systems for Fossil Energy Conversions. 2010, Hoboken, NJ: American Institute of Chemical Engineers (AIChE) and John Wiley & Sons, Inc.

52. Alioshin, Y., et al., High Temperature Electrolysis of CO2 for Fuel Production. Journal of The Electrochemical Society, 2016. 163(2): p. F79-F87.

53. Banerjee, A., R. Bala Chandran, and J.H. Davidson, Experimental investigation of a reticulated porous alumina heat exchanger for high temperature gas heat recovery. Applied Thermal Engineering, 2015. 75: p. 889-895.

54. Heat Transfer International. High Temperature Ceramic Heat Exchangers. [cited 2015 July 14]; Available from: http://www.heatxfer.com/products/htche.html.

55. Yun, S. and S. Ted Oyama, Correlations in palladium membranes for hydrogen separation: A review. Journal of Membrane Science, 2011. 375(1–2): p. 28-45.

56. Kanezashi, M. and M. Asaeda, Hydrogen permeation characteristics and stability of Ni- doped silica membranes in steam at high temperature. Journal of Membrane Science, 2006. 271(1–2): p. 86-93.

57. SunShot Concentrating Solar Power (CSP) R&D Funding Opportunity Announcement, Department of Energy, DE-FOA-0000595 CFDA Number 81.087. 2011: Golden, CO.

58. Smith, J.M., H.C. Van Ness, and M.M. Abbott, Introduction to chemical engineering thermodynamics. 7th ed. 2005: McGraw-Hill.

59. Hablanian, M.H., High-Vacuum Technology: A Practical Guide, Second Edition. 1997: Taylor & Francis.

117

60. Dyer, P.N., et al., Ion transport membrane technology for oxygen separation and syngas production. Solid State Ionics, 2000. 134(1–2): p. 21-33.

61. Meixner, D.L., et al., Electrochemical Oxygen Separation Using Solid Electrolyte Ion Transport Membranes. Journal of the Electrochemical Society, 2002. 149(9): p. D132-D136.

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CHAPTER III

SPRAY DRYING ACTIVE PARTICLES

3.1. Abstract

Particles are an important aspect of many proposed solar thermochemical water splitting reactor concepts. These particles must maintain physical integrity and chemical performance at very high temperatures while moving around a system. Spray drying is used to produce iron aluminate (hercynite) particles using pH-modification of a charge stabilized sol as a precursor.

Nanoparticle suspensions are mixed and the pH is modified in order to induce partial flocculation.

This gives larger particles that are more spherical and structurally robust. This relatively simple technique shows promise for a new material and application.

High-temperature thermochemical energy storage shows promise in aiding concentrating solar power plants in meeting variable, grid-scale electricity demand. Manganese oxide-based mixed metal oxide particles have been designed and tested for thermochemical energy storage.

Particles are designed for high energy storage capacity, flowability, and physical and chemical stability. We evaluate the effects of Al2O3, Fe2O3, and ZrO2 in Mn2O3-based spray-dried particles in a TGA between 650°C and 1,200°C over six consecutive redox cycles. Results are compared with thermodynamic predictions from 400-1,400°C under oxidizing and reducing atmospheres. A mixture of 2:1 Fe2O3:Mn2O3 formed iron manganese oxide spinel (FeMn2O4) on calcination, and demonstrated the highest thermochemical activity despite particle agglomeration and deformation.

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3.2. Introduction

Metal oxides can be very useful for a variety of high temperature thermochemistry applications, but only if these materials can be made into a form useful in a realistic reactor. The background of active particles for solar thermochemical water splitting and solar thermochemical energy storage will be briefly reviewed, and spray drying will be highlighted as a useful technique for these types of materials and applications.

3.2.1. Active Particles for Solar Thermochemical Water Splitting

Many current efforts explore the use of particle reactors for solar thermochemical water splitting; these are described in Chapter 1. These particles are produced by sol-gel methods [1], thin film deposition [2], and solid state synthesis [3] among many others. These particles have been used for their high surface area relative to monolith structures. Characterization and experiments have been performed on flowing or falling particles [4-7] as well as fixed bed particles

[8, 9]. In addition to the reactivity of the particles, solid particles have also been examined for transporting and exchanging sensible energy [8, 10-13]. These applications show the importance of flowable and reactivity particles that need to withstand high temperatures.

3.2.2. Thermochemical Energy Storage

Thermochemical energy storage is very similar to thermochemical waster splitting, but instead of producing H2 the chemical cycle stores and releases heat via the enthalpy of reaction.

High-temperature thermochemical energy storage is a promising approach for efficient and cost- effective storage of concentrated solar energy for dispatchable solar-thermal power generation.

Solid metal oxide materials are considered as high-temperature thermochemical energy storage media as they often reduce at the high temperatures reachable with point-focusing solar concentrating systems and have the potential to reach high energy storage densities [14-16].

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Manganese oxide is of particular interest as it is inexpensive, non-corrosive, and relatively hazard- free. It reversibly reduces from Mn2O3 to MnO in the temperature range of 500 to 1,500°C in an oxygen-lean environment. Particles can then be redox cycled in a thermochemical storage system involving two fluidized-bed reactors, one for solar-driven high-temperature endothermic reduction to “charge” the particles and one for non-solar lower-temperature exothermic oxidation to

“discharge” the particles. The goal is to develop easily flowable particles that exhibit high storage capacity, physical robustness, and chemical activity over thousands of redox cycles.

Manganese oxides have previously shown promise as redox materials for solar thermochemical water splitting [17-19] and chemical looping combustion [20, 21]. A cycle utilizing sodium hydroxide with manganese oxide over three or more reactions has also been studied [22-26]. In these cycles, various oxidation states of manganese oxide have been matched to temperatures and co-reactants to utilize the various oxidation states of manganese oxide.

Manganese oxide has also been studied for thermochemical energy storage. Wong et al. studied the thermodynamics of this material for thermochemical energy storage [27]. Carrillo et al. tested manganese oxide over many cycles and found it to be highly chemically stable and very promising as a thermochemical storage material [28, 29].

As MnO2 is heated to higher temperatures, it undergoes several reduction steps, shown in

Equation (2.21). Similarly, as hot MnO is cooled down to room temperature, it reoxidizes via several oxidation steps back to MnO2. The temperatures at which these transitions occur depend on the gas environment in which the reaction is conducted; indeed, MnO2 is typically not observed upon reoxidation, as a very high oxygen partial pressure is required [30].

푀푛푂2 ↔ 푀푛2푂3 ↔ 훼 − 푀푛3푂4 ↔ 훽 − 푀푛3푂4 ↔ 푀푛푂 (2.59)

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A number of secondary oxides have been examined for use in thermochemical energy storage with manganese oxide. Some of these improve re-oxidation performance [27] while others rely on a multi-doped manganese perovskite [31, 32]. Carrillo et al. found that while mixed Co/Mn oxides performed worse than the corresponding pure oxides [28], Fe-doping of the Mn oxide improved energy storage density and reaction stability [33]. Azimi et al. showed active particles with a 2:1 Fe:Mn molar ratio in a high-temperature redox cycle optimized the activity of particles and also their retention of fluidizable, spherical morphology over multiple cycles. Particles made with lower Fe:Mn ratios did not release oxygen when temperature cycled, and particles made with higher Fe:Mn ratios experienced sintering and defluidization at the temperatures tested [34].

However, it was not clear why this behavior occurred. Later, Azimi et al. examined the addition of aluminum oxide to Fe/Mn materials and found a trade-off between reactivity and physical robustness [21].

3.2.3. Spray Drying

Spray drying is a commonly used method of making high-performance, fluidizable particles for chemical looping combustion applications [35, 36], and spray dried have been shown to withstand long exposure to high temperature cycling [37]. The morphology of spray dried particles impacts flowablity and friability [38]. Spherical particles are easier to fluidize [39] and are likely to survive longer in a physical and thermally stressful environment than particles of other shapes. Colloidal metal oxide dispersions are often used in the manufacture of spray-dried particles to create better particle cohesion during the spraying process and to improve the particle robustness under high-temperature cycling conditions by acting as a structural support for the active material in the particle [40]. In addition, some metal oxide dopants (including Fe2O3, ZrO2, CuO, ZnO, and

TiO2) have been found to increase the typically slow kinetics of the reoxidation step Mn3O4 →

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Mn2O3 [30]. On the other hand, structural supports are typically chemically inactive. Hence, adding an inactive second metal oxide to the particles reduces their mass-specific oxygen exchange capacity and energy storage density.

There are many factors to consider with spray drying, each of which impacts the final particles produced. The temperature and gas flow rate inside the drying chamber can change drying rates which can affect how droplets dry and thus can affect the final particle properties. However, many groups have found that the composition, rheological properties, and method of preparing the solution or slurry that is fed into the spray dryer has a much greater impact on a number of properties of the final particles [40]. Lukasiewicz pointed out that since a lot of suspensions to be spray dried comprise micrometer (or less) sized particles, the slurry can be considered a colloidal dispersion, meaning that DLVO theory can be used to describe the stability of the suspension [41].

This theory states that a stable colloid balances attractive van der Waals forces with repulsive electrostatic forces. Walker, et al. proposed a model for how changes to colloidal stability of the slurry can be used to modify the morphology of spray dried particles [42]. Specifically, when the pH of the suspension is modified so that the zeta potential is near zero, the repulsive electrostatic forces are minimized which leads to less stable colloids and particle agglomeration and flocculation [42]. An increase in yield stress upon flocculation will prevent particle flocs from re- arranging as the droplet dries in the spray drier. Particles formed by this method are less dense than particles sprayed from stable slurries, but form solid spheres as a network of flocs.

Conversely, particles in deflocculated slurries can re-arrange and pack much more densely at the droplet surface as it dries, leading to a crust and then a hollow or “donut” shaped particle [42].

This behavior has been shown for many different materials, including alumina [41-46], titania

[44], zirconia [46-51], and silica [52] among others. Many of these reports [41-46, 48, 49] use

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organic binders, which are necessary to hold the particle together after drying but before sintering

[41]. In addition to yield stress [42, 45] and zeta potential measurements [45], single droplet tests

[48, 49] and sedimentation tests [45, 46, 48, 49] are used to quantify the flocculation of the suspension.

3.2.4. Specific Novelty of This Work

In this work, we evaluate the effect of three secondary metal oxides (Al2O3, ZrO2, and

Fe2O3) on the redox behavior and sintering temperatures of manganese-oxide-based materials.

Colloidal suspensions of the secondary metal oxides are combined with Mn2O3 nanopowder to manufacture spray-dried particles. Particles are tested in a thermogravimetric analyzer (TGA) over six consecutive redox cycles with oxidation temperature of 650°C and reduction temperatures of

950, 1,050, and 1,200°C, to evaluate the impact of the different secondary metal oxides on chemical activity and robustness of the particles. The particles are characterized after spray-drying, and after calcining, with respect to their shape, crystal structure, specific surface area, and particle size distribution, using scanning electron microscopy (SEM), X-ray diffraction (XRD), Brunauer–

Emmett–Teller (BET) surface area, and laser diffraction. This characterization helps to explain various performance trends and trade-offs for future development of these particles.

In this work, we implement a previously used method to produce spray dried particles with a different combination of ceramic starting particles. Specifically, we use a known stable colloid suspension with partial flocculation driven by pH modification in order to produce solid spherical particles of a mixed metal oxide. Particle morphology is shown by SEM. Zeta potential and yield stress measurements are used to show partial flocculation, which does not exhibit expected sedimentation behavior. This new formulation and effects will be helpful for the future scalable production of solar thermochemical water splitting particles.

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3.3. Methods

A number of slurries are prepared, typically based on nanoparticle colloids in water. These slurries are characterized and spray dried to form aggregates, which will then be calcined at high temperature to form solid particles. These particles are characterized for morphology, structure, composition, and reactivity.

3.3.1. Preparation of Spray Dried Particles

Hercynite slurries are prepared by combining purchased commercial suspensions of aluminum oxide (NYACOL Nano Technology, AL20DW, Colloidal Alumina Sol) and iron oxide

(Aldrich, 720704, Iron(III) oxide, dispersion). The relative amounts of each of these suspensions were calculated to achieve the 1:2 Fe:Al molar ratio in FeAl2O4 (hercynite). The appropriate amount of each suspension was weighed out and combined in a beaker, which was then stirred for

20 minutes on a stir plate at 400 rpm.

Five candidate manganese oxide materials are evaluated, and are summarized in Table 3.1.

A slurry of Mn2O3 nanopowder (US Research Nanomaterials Mn2O3 Nanopowder), DI water, and in most cases a colloidal secondary metal oxide (Nyacol ZrO2(AC) Acetate stabilized colloidal zirconia, Nyacol AL20DW Colloidal alumina, or Aldrich Iron (III) oxide dispersion) is prepared as the precursor solution. Slurry suspensions are mixed on a stir plate for 20 minutes at 400 rpm as-prepared.

Table 3.1. Overview of composition and sample ID for candidate materials.

Description Sample ID

Mn2O3 spray dried with no secondary metal oxide NS

Mn2O3 spray dried with 30wt% Al2O3 Al30

Mn2O3 spray dried with 30wt% ZrO2 Zr30

Mn2O3 with 30wt% Fe2O3 Fe30

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Mn2O3 and 67wt% Fe2O3 Fe67

1M NaOH (2.5M for hercynite slurries) or 1M HCl were added dropwise to the slurry to affect the pH of the suspension, which was measured with a digital pH meter. After each drop, the suspension was allowed to settle and stir for 2 minutes before the pH was measured again. This process continued until the desired pH was achieved for each sample. Water was added as needed to make slurry able to spray. Prepared slurries are 20-30 wt% solids, and the powders and colloidal dispersions used all have particle diameters < 100 m.

Spray-dried metal oxide particles are made using a Büchi B-290 Mini Spray Dryer. The slurry is fed into the nozzle of the spray dryer via a peristaltic pump where it is nebulized with compressed air to form droplets. The liquid droplets are entrained in a flow of hot air (~200 C) in the drying chamber. Particles form as the droplets dry, and are collected and separated from ultra- fines using a cyclone.

After spray drying, particles are calcined in a box furnace under ambient atmosphere at

1,200 ºC for 8 hours. This was done to form the thermodynamically favorable mixed metal oxide phases, and to sinter the particles prior to analysis.

3.3.2. Particle Characterization

Particles are imaged in a JEOL JSM-6480LV Scanning Electron Microscope. The BET surface area is obtained using a Micrometrics Gemini V Surface Area and Pore Size Analyzer.

Particle size distribution is determined using a Malvern Mastersizer 2000 with a Hydro 2000SM wet dispersion unit. Crystal structure of the spray-dried materials after calcining is determined using a Bruker D2 Phaser X-Ray Diffraction Desktop System with Cu Kα radiation (λ = 1.5418

Å). Particle composition was determined by inductively coupled plasma optical electron spectroscopy (ICP-OES) using an ARL 3410+ inductively coupled optical emission spectrometer.

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For the ICP-OES, samples were digested prior to analysis using a mixture of hydrochloric, hydrofluoric, and nitric acid at 95°C for two hours, after which boric acid was added and held for

15 minutes, samples were cooled, diluted, and analyzed.

Thermogravimetric analyses are performed with a Netzsch STA 449 F1 Jupiter. Samples of ~20 mg calcined metal oxide particles are placed in an alumina-lined platinum crucible. Prior to redox cycling, samples are heated to 105°C for 30 minutes under argon flow to remove any adsorbed water. After the drying step, samples are redox cycled six times. Each candidate material was cycled at three different reduction temperatures (900, 1,050, or 1,200°C) and an oxidation temperature of 650°C. One redox cycle consists of a 10°C/min ramp to the reduction temperature, a 15-minute hold at the reduction temperature, and a 10°C/min ramp down to the oxidation temperature followed by a 45-minute hold at the oxidation temperature. Every part of the redox cycle is run under argon except for the 45-minute hold at the oxidation temperature, which is run under air.

Thermodynamic equilibrium calculations are performed to predict the chemical and phase composition of manganese oxide-based materials as a function of secondary metal oxide, temperature, and gas environment. These calculations indicate the temperatures at which reaction steps shown in Equation (2.21) occur, and the temperature ranges in which manganese oxide and the secondary metal oxide form mixed metal oxide phases. In addition, the results yield the upper temperature limit imposed by the formation of a slag phase (liquid solution phase), which results in sintering of the particles. The parameters used in the calculations are summarized in Table 3.2.

Table 3.2. Conditions used in the FactSage thermodynamic equilibrium calculations of manganese oxide-based mixed metal oxides. Pressure 1 atm Temperature 400-1,400°C Secondary metal oxides ZrO2, Al2O3, Fe2O3 Initial solid material 1 mol of mixed metal oxide (Mn2O3 + secondary)

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Initial gas compositions Inert gas: 100 mol N2, Air: 79 mol N2, 21 mol O2 Possible gaseous products 18 Possible liquid products 6 Possible solid products 18

Thermodynamic prediction calculations are performed using the FactSage Gibbs free energy minimization software package [53]. This software uses a thermodynamic database of enthalpy and entropy values, which are combined to calculate the Gibbs free energy for a variety of materials at the specific conditions of interest. The total Gibbs free energy of the system is then calculated subject to mass constraints and the amounts of various possible materials are varied to minimize the free energy [54]. The FactPS and FTOxid databases within FactSage are used in this study. Maximum theoretical mass changes during reduction and oxidation were calculated for the candidate materials using reaction stoichiometry.

Zeta potential measurements were performed using a Malvern Zetasizer Nano ZSP using a disposable folded capillary cell. The suspensions of interest were optically opaque, and so they were centrifuged at 4,300 xg for 7 hours. After this, the supernatant was removed and measured for zeta potential. The viscosity and shear stress of slurries was determined using a Discovery

Hybrid Rheometer (TA Instruments DHR-3). Samples were pipetted between the bottom of a temperature controlled Peltier plate and a 60 mm cone tool (12 μm truncation, 0:28:52 angle).

Measurements were performed in triplicate at shear rates from 50 to 500 1/s and a temperature of

25°C. Measurements were also made with a 20 mm parallel plate at 25°C (300-micron gap) and samples were protected from evaporation effects by a thin mineral oil ring.

3.4. Results

Spray dried particle morphology and size is critical for use in high temperature reactors that depend on the motion of active particles. Composition and crystal structure is also critical for thermochemical cycle chemistry. The spray dried particles have been characterized for these

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factors, and additional characterization has been done on the slurries that were used to spray these particles in order to achieve a better understanding of what caused various particle properties.

3.4.1. Manganese Oxide Particle Characterization

SEM images of the spray-dried candidate materials before calcining are shown in Figure

3.1. Spray-drying a pure manganese oxide slurry with no added secondary colloidal suspension produces few spherical particles, and instead yields mostly flakes (Figure 3.1a). Including 30wt% of Al2O3, ZrO2, or Fe2O3 as added colloidal suspensions in the spray drying slurry results in a significant fraction of spherical particles (Figure 3.1b-d), along with needle-like and oblong formations. Including 67wt% of Fe2O3 as a colloidal suspension leads to the formation of a very large fraction of polydisperse spherical particles with nearly no formation of needles, oblong formations, or any other non-spherical shapes. Addition of colloidal suspensions to the manganese oxide slurries appears to have helped spherical particles form; this is consistent with prior work in which a colloidal boehmite binder gave spherical particles with higher mechanical strength than alumina nanoparticle slurries [55]. Fe67 was the most affected by pH tuning, becoming very viscous around pH 7-8. None of the other slurries appeared to be affected by pH tuning, with no visible coagulation of the slurry over a pH range of 4-11. Using pH alone to control the dispersion of a mixed-oxide suspension has been previously identified as difficult [55], but additional effort is required to determine if such control is possible for these solutions, such as zeta potential measurements.

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a) 5 μm b) 5 μm

c) 5 μm d) 5 μm

e) 5 μm

Figure 3.1. SEM images of spray-dried particles before calcine: a) NS, b) Al30, c) Zr30, d) Fe30, e) Fe67. White scale bars represent 5 m lengths. Liquid entrainment necessary for particle size analysis would likely have contributed to the degradation of the pre-calcined spray dried particles at least partially back into their original nanoparticle components, and so therefore the particle size distribution of spray dried particles before calcining was not measured. However, from a qualitative visual assessment of the SEM images in Figure 3.1, it appears that most of the distinct particles produced using the laboratory- scale spray dryer have a diameter of roughly 1-10m.

The selection of the secondary oxide used in the spray-drying process strongly influences how well particles retain their original shape and resist sintering after 8 hours of calcining at

1,200°C. Comparison of the morphologies in Figure 3.1 and Figure 3.2 shows that NS (a), Fe30

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(d), and Fe67 (e) lose spherical particle shape entirely during calcining. Al30 (b) and Zr30 (c) retain some spherical particles. This could be attributable to some materials (Al30 and Zr30) maintaining separate support phases that help to main this morphology, whereas materials that combine to form a single mixed-metal phase (NS, Fe30, Fe67; this is discussed below) are not able to maintain the spherical particle morphology.

a) 5 μm b) 5 μm

c) 5 μm d)) 5 μm

e) 5 μm

Figure 3.2. SEM images of spray-dried particles after calcine at 1,200°C for 8 hours: a) NS, b) Al30, c) Zr30, d) Fe30, e) Fe67. White scale bars represent 5 m lengths. Measurement of BET surface area of the particles before and after calcining reveals significant reduction in particle surface area after calcining (Table 3.3). Although the degree of interparticle agglomeration due to calcining cannot be quantified without particle size distribution data for the pre-calcined materials, measurement of the particle size distribution of the candidate

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materials after calcining reveals much larger average particle sizes, and much broader distributions than would have been expected for the pre-calcined materials based on SEM images. Figure 3.3 and Table 3.3 show the volumetric particle size distribution for calcined candidate materials. The diameter range of NS particles is considerably smaller than the rest of the materials, with 50% of particles remaining < 50 m diameter. It is not clear, however, whether this is because the pure manganese oxide is less inclined to sinter than a mixed metal oxide, or whether this is because the

NS samples suffered from poor particle formation in the spray dryer and subsequently started out with much smaller, nano-scale powder instead of intact spray dried particles. Zr30, Fe30, and Fe67 had wide particle size distributions. The two materials containing iron oxide were skewed toward higher particle diameters, with 50% of their sample volumes comprising particles 259 and 523 μm, respectively. This could be due to formation of larger particles during spray drying with ZrO2 and

Fe2O3, or it could be an indication of a higher propensity toward sintering and agglomeration at

1,200°C.

Table 3.3. BET surface area (before and after calcining) and particle size distribution for spray dried candidate materials after calcining at 1,200°C. BET Surface Area (m2/g) Particle Size Distribution (μm) Sample ID Pre Calcine Post-Calcine D10 D50 D90 NS 10.22 0.37 5 46 229 Al30 69.87 1.28 7 73 394 Zr30 64.75 0.74 9 102 890 Fe30 35.08 0.39 17 259 854 Fe67 54.97 0.61 33 523 1250

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16 5 NS 4 14 Al30 %

Zr30 e 3

u 12 Fe30 l 2

Fe67 V 1 10 0

% 0 20 40 60 80 100

m 8 Particle Diameter (µm)

o V 6

0 0 500 1000 1500 2000 Particle Diameter (µm)

Figure 3.3. Particle size distribution of candidate materials after calcining at 1,200°C. The x-axis on the inset figure is zoomed in on range of particles from 0-100 m, and larger figure’s axis is extended over the entire particle range. The XRD spectra for the calcined candidate materials are displayed in Figure 3.4. The pure manganese oxide material, NS, shows peaks corresponding to the Mn3O4 spinel phase. XRD peaks that do not match the Mn3O4 spinel phase seem to correspond well with a mixed Mn/Na oxide phase, as shown in the figure. The thermodynamic predictions done below for the inclusion of Na predict a mixed-metal slag phase, which may present as XRD crystal peaks. Mixed metal oxide

Al30 has peaks corresponding to a mixed manganese aluminum oxide spinel (MnAl2O4) and a mixed Na/Mn oxide phase, similar to NS but somewhat different peaks. XRD spectra of Zr30 indicate that manganese oxide and zirconia do not form solid solutions at 1200 C. Zirconia in the sample consists of a mixture of monoclinic and tetragonal ZrO2. The tetragonal ZrO2 is matched with cubic ZrO2 in the figure, as the peak positions of tetragonal and cubic zirconia are very similar

[56].Manganese oxide is present in the Zr30 sample as the Mn3O4 spinel phase. Fe30 spectra correspond to MnFe2O4 spinel, and a similar mixed Na/Mn phase as above. Fe67 spectra indicate that this sample is entirely composed of MnFe2O4 spinel.

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Figure 3.4. X-ray diffraction spectra for spray-dried candidate materials after calcining at 1,200°C 3.4.2. Thermodynamic Predictions of Manganese Oxide Particles

In order to gain a more complete understanding of the effect of secondary metal oxide additives to the manganese oxide active materials, thermodynamic predictions for each material were generated using the FactSage software. Predictions of the favored metal oxide phases present under given temperature and atmospheric conditions encompassing the calcining process and subsequent redox cycling conditions (400-1,400 C under air and N2 atmospheres) are shown in

Figure 3.5. Table 3.4 presents predicted slagging temperatures for candidate materials under air and

N2 atmospheres.

Zirconia does not form a solid solution with manganese oxide up to ~1,300°C under an inert atmosphere, instead the zirconia and manganese oxide maintain separate phases. This prediction is supported by the XRD spectra for Zr30 calcined in air at 1,200°C, which shows no

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indication of the presence of a mixed metal oxide. Alumina forms a mixed bixbyite phase with manganese oxide (Mn,Al)2O3 at temperatures below 700°C and transitions to a MnAl2O4 spinel between 700 and 800°C. XRD spectra corroborate the formation of a mixed (Mn,Al)3O4 spinel phase during calcination in air, and also confirm that all added alumina is incorporated into the solid solution. The formation of this spinel phase may prove to be problematic in the cycleability of the material, unless the spinel phase is redox active. Iron oxide is predicted to form entirely mixed metal oxide phases for both Fe30 and Fe67 compositions. Increasing the Fe2O3 concentration eliminates the presence of the α-spinel (Mn,Fe)3O4 phase, and introduces a hexagonal (Mn,Fe)2O3 phase. Interestingly, although FactSage predicts the formation of a mixed

(Mn,Fe)3O4 spinel for both Fe30 and Fe67 at high temperatures, the XRD spectra for the two materials are quite different; this will be discussed below.

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Figure 3.5. Thermodynamic predictions of reduction and oxidation behavior of candidate materials in air and inert environments.

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Table 3.4. Thermodynamic predictions of slagging temperatures of candidate materials in air and inert environments.

Candidate Predicted Slagging Temperature (°C)

Material N2 Atmosphere Air Atmosphere NS 1,657 1,581 Al30 1,455 1,465 Zr30 1,457 1,499 Fe30 1,588 1,572 Fe67 1,565 1,594

3.4.3. Gravimetric Measurements of Manganese Oxide Particles

Greater repeatable mass change during redox cycling corresponds to higher oxygen exchange (whereas other effects such as material sublimation would appear irreversible), which is an indication of higher chemical storage capacity. The results from the TGA redox cycling tests are summarized in Figure 3.6. Fe67 and NS outperform the other candidate materials at all three reduction temperatures in terms of their specific mass change during cycling. Zr30 performs best at a reduction temperature of 1,050°C, and indicated a higher specific mass change than either

Al30 or Fe30. Fe30 shows virtually no redox cycling capabilities at any reduction temperature.

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Figure 3.6. Thermogravimetric analysis for candidate active materials over six redox cycles, with oxidation at 650°C and reduction at 900°C, 1,050°C, and 1,200°C. Plots at the top of each column show the temperature profile, and plots below show the mass change of the sample during redox cycling. 3.4.4. Characterization of Hercynite Suspension

The viscosities of the hercynite slurries that have been modified to a specific pH have been measured and are shown in Figure 3.7. All slurries have a higher viscosity than water at all shear rates, due to the presence of solids in the suspension. As the pH of the slurry is increased from 3 to 7, the viscosity of the slurry also increases. Conversely, as the pH continues to increase to 9, the viscosity falls. This is attributed to peak flocculation at a pH of around 7, increasing the viscosity of the slurry, whereas at a pH less than or greater than 7, the viscosity decreases.

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Figure 3.7: Measured viscosity of pH-tuned hercynite slurries at various shear rates. Final pH of slurry is given in legend, and pure water is included for comparison. 95% confidence intervals are shown for multiple measurements. The data for pH 11 is not shown due to measurement issues. Measured stress values at various shear rates are shown in Figure 3.8 for hercynite slurries modified to different pH values. While the DI water and slurries tuned to a low pH value seem to be generally Newtonian fluids (i.e., linear slope and projected intercept of zero stress at zero shear), the slurries with pH values ≥7 show non-Newtonian characteristics. This is attributable to higher degrees of flocculation in the suspension.

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Figure 3.8: Measured stress of pH-tuned hercynite slurries at various shear rates. Final pH of slurry is given in legend, and pure water is included for comparison. 95% confidence intervals are shown for multiple measurements. The data for pH 11 is not shown due to measurement issues. The yield stress of the pH-tuned hercynite slurries was calculated using the Casson equation (similar to [51]), and is shown in Figure 3.9. High yield stress values correspond to flocculation conditions, further indicating flocculation at a pH value of 7 and above.

Figure 3.9: Calculated yield stress of hercynite slurries modified to various pH values. 95% confidence intervals are shown for each calculated value. The data for pH 11 is not shown due to measurement issues. The viscosity of the boehmite sol that supplied the aluminum to the hercynite particles was also modified to various pH values and the viscosity of each was measured. This is reported in

Figure 3.10 and shows higher viscosity values for pH values of 7 and above.

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Figure 3.10: Measured viscosity of pH-tuned boehmite slurries at various shear rates. Final pH of slurry is given in legend, and pure water is included for comparison. 95% confidence intervals are shown for multiple measurements. The data for pH 5 is not shown due to measurement issues. Measured stress values at various shear rates are shown in Figure 3.11 for boehmite slurries modified to different pH values. Again, DI water and the low-pH slurries seem to be generally

Newtonian fluids while the slurries with pH values ≥7 show non-Newtonian characteristics.

Figure 3.11: Measured stress of pH-tuned boehmite slurries at various shear rates. Final pH of slurry is given in legend, and pure water is included for comparison. 95% confidence intervals are shown for multiple measurements. The data for pH 5 is not shown due to measurement issues.

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Similar to the hercynite slurries, the stress-shear results for the pH-modified boehmite slurries were also fit to the Casson equation to calculate yield stress. The calculated values are shown in Figure 3.12, and show a peak of yield stress at a pH value of ~7. This is similar the hercynite slurry yield stress shown in Figure 3.9, and suggests that the boehmite constituent drives the yield stress of the hercynite mixture.

Figure 3.12: Calculated yield stress of boehmite slurries modified to various pH values. 95% confidence intervals are shown for each calculated value. The data for pH 5 is not shown due to measurement issues. The zeta potential of the hercynite slurry is shown in Figure 3.13. As can be seen, the isoelectric point of this slurry is estimated to be between 7 and 8, based on where the zeta potential is equal to zero.

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Figure 3.13: Zeta potential of combined iron oxide and aluminum oxide slurry. Average measurement is shown with 95% confidence intervals. The zeta potential of boehmite slurry is reported in Figure 3.14 with 95% confidence intervals. The iso-electric point based on these data is measured to be ~8. This is again very similar to the hercynite slurry shown above, suggesting that the behavior of the hercynite slurry is driven by the behavior of the boehmite sol.

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Figure 3.14: Zeta potential of boehmite slurry. Average measurement is shown with 95% confidence intervals. 3.4.5. Hercynite Particle Characterization

Four formulations of hercynite slurries were prepared and spray dried. The resulting particles were imaged with SEM and the images are shown in Figure 3.15. These images show small (<10 μm diameter) particles for the low (3.7, 5.4) and high (10.1) pH values, but larger (>10

μm diameter) and much more spherical diameter particles. a) b)

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c) d)

Figure 3.15: SEM images of spray dried hercynite particles, each with 10wt% excess alumina and a pH modified to a) 3.7, b) 5.4, c) 7.4, and d) 10.1. All images were done at a magnification of 1,000X, and the scale bar represents 10 μm. 3.5. Discussion

Mixed-metal oxide systems can have a number of benefits, but can also give rise to greater complexity. Many of the results shown above show materials that do not necessarily behave in a straightforward way. The underlying reasons for the complexity seen in the characterization results for the manganese-based mixed metal oxide particles will be discussed. Then reasons for the beneficial action of pH tuning on the hercynite slurries will be discussed in light of other possible explanations of this effect.

3.5.1. Mass Change in Mixed-Metal Manganese Oxide Particles

The substantial difference in mass changes per cycle between Fe30 and Fe67 (Figure 3.6) is especially interesting due to the similar equilibrium compositions predicted (Figure 3.5). The

XRD patterns (Figure 3.4) of these two samples are also different, despite the same phase predicted at equilibrium. Taken together, we attribute these results to the Fe30 sample not reaching thermodynamic equilibrium during calcining while Fe67 does reach equilibrium. The calcination process is expected to be accompanied by a mass loss (as the metal oxides go from M2O3 to M3O4), and this is consistent with the mass loss seen in the TG results. The slight drift seen in all the

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samples in Figure 3.6 could also be attributed to further approaching thermodynamic equilibrium.

The XRD pattern for Fe30 does show the spinel-structure peaks that appear prominently in Fe67, but the peaks are much weaker in magnitude and other small peaks are also present; this is indicative of a material that has not crystallized fully to the equilibrium condition. This difference in the end states of the Fe30 and Fe67 samples would also explain the similar results obtained by

Azimi et al., in which a sample with a 1:2 molar ratio of Fe:Mn had almost no observable oxygen release whereas a sample with a 2:1 molar ratio of Fe:Mn performed the best of all the samples studied [34].

However, Carrillo et al. reported that Fe/Mn mixed oxide samples (20 and 40 mol% Fe) showed clear spinel-structure XRD peaks and repeatable mass cycling in a TGA, despite the fact that the samples were only calcined at 700°C and for 4 hours [33]. We attribute this difference to the sample preparation technique; Carrillo et al. used a Pechini method [33] while this work spray dried constituent metal oxides. A sol-gel-type method such as the Pechini method will produce a much more homogenous mixture of metal ions, whereas a solid state synthesis such as the one used here will require the diffusion of metal ions into neighboring crystallites. Thus, we attribute the dramatic differences between the Fe30 and Fe67 samples to differences in ion diffusion rates.

Gilezicz-Wolter et al. demonstrated almost 2X faster diffusion of Mn relative to Fe [57]. In the

Fe30 sample, the slower Fe would need to diffuse farther through relatively more bulk manganese oxide to reach equilibrium, whereas in the Fe67 sample the Fe needs to travel through less manganese oxide. Conversely, the faster Mn must travel further in the Fe67 sample, leading it to reach thermodynamic equilibrium more quickly. This issue can be overcome in future work by lengthening the time at high temperature during calcining if solid state synthesis is used, or by using a synthesis technique that produces a more homogenous ionic mixture before calcining.

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The mass change during reduction and oxidation for the top performing candidate materials

Fe67, NS, and Zr30 is shown in Figure 3.16. Neither Fe67 nor Zr30 exhibit a decrease in mass change over six cycles at any reduction temperature tested; however, the mass change decreases for NS during repeated reduction at 1,200°C. Fe67 outperforms the other candidate materials when reduced at both 1,050°C and 1,200°C. The mass changes for Fe67 and NS decrease considerably at a reduction temperature of 900°C, while the influence of reduction temperature on the mass change of Zr30 is not as pronounced. At a reduction temperature of 1,200°C, Fe67 has a specific mass change 50% greater than NS and 150% greater than Zr30.

It is likely that most of the samples do not display decreased apparent mass loss over six redox cycles because the particles have already undergone significant sintering and surface area loss during the calcining process. Loss of surface area before TGA testing could be responsible for the lower than theoretically predicted mass loss results (Figure 3.16). Compositional analysis using

ICP-OES of candidate materials lends further credence to this theory, and provides a possible explanation for the unexpectedly high amount of particle sintering and the lower than expected mass loss. Elemental analysis revealed that the Mn2O3 nanopowder used to make the particles contains substantial quantities of sodium (~0.5 wt%). Including this sodium contamination in thermodynamic predictions of candidate materials dramatically reduces the predicted slagging temperatures of these materials; in the case of pure manganese oxide, slagging is predicted at

<800°C in air. This will drastically reduce surface area available for reaction. Furthermore, the sodium contamination causes the formation of additional crystal phases which do not undergo the redox cycling of interest. This causes a portion of the active material in each sample to become inert, thus lowering the specific activity by contributing mass but not chemical activity. These

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findings underscore the importance of sourcing pure materials when producing spray dried active particles.

3.5 Theoretical Mass Change for NS = 3.38%

Theoretical Mass Change for Fe67 = 3.35%

3.0

2.5

Theoretical Mass Change for Zr30 = 2.20%

g 2.0

C Fe67 1200 ºC

M Fe67 1050 ºC 1.5

% NS 1050 ºC

NS 1200 ºC 1.0

Fe67 900 ºC NS 900 ºC 0.5 Zr30 1050 ºC Zr30 1200 ºC Zr30 900 ºC

0.0

1 2 3 4 5 6 Cycle Number

Figure 3.16. Mass change for top performing candidate materials over six redox cycles. Theoretical maximum mass change for complete reduction is shown by solid lines. 3.5.2. Effect of Flocculation on Droplet Formation in Hercynite Particles

Two different metal oxide nanoparticle suspensions were used to form solid, spherical spray dried particles. The particles produced from the mixed suspension were largest and most spherical when the pH of the suspension was modified to ~7. Here the zeta potential shows a value very close to zero. With near-zero net surface charge, the particles have no columbic repulsion to keep them more stable in suspension. This has been shown to produce larger and more spherical particles because of the inability of the solid particles to easily flow and re-arrange [42].

The increased viscosity and yield stress of the suspension at the IEP also indicates flocculation [42, 45]. As the repulsive forces between particles are reduced, the relative inter-

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particle attraction increases which thus increases the resistance to flow in the suspension. Prior works [42, 46] have shown that this does not necessarily make for particles with maximum density; indeed, because the flocs are unable to flow around in suspensions as easily as individual particles, they pack less densely. The dense packing at the surface of the droplet as it dries is what leads to hollow shell rather than solid spherical particles. This must be kept in mind for applications in which dense solid spheres are desired; longer calcine times at higher temperature are likely needed to fully densify these spheres.

Many prior explorations of spray drying ceramic particles have used aluminum oxide, but to our knowledge this is the first time that it has been combined with iron oxide. The results of modifying the pH are consistent with previous studies, in that suspensions modified to the IEP tend to show a higher yield stress and yield more spherical particles when spray dried [42, 45, 46,

48, 49]. However, prior investigations used solid particles dispersed in water with an organic dispersant and binder, while the current effort relies more on metal hydroxides with minimal dispersant. Some have relied on charge stabilization alone, but tended to rely on concentration or flow rates in order to produce particles with the desired morphology [52]. This can be more difficult and expensive at scale than pH modification. The iron oxide suspension used here incorporates a ethoxylated carboxylic acid, while the alumina suspension is charge stabilized only.

Some prior works have used pseudo-boehmite sols (very similar to the one used here) to obtain large, solid, and spherical particles [58-62], though these works used a less concentrated sol [58] or use spray pyrolysis instead of spray drying [61]. Furthermore, many these previous efforts tended to not perform pH modifications. The exception is Albano and Garrido [62], which used pH modification with a pseudo-boehmite sol. However, this work used the pseudo-boehmite to coat other solid particles, rather than to help form the spray-dried spheres itself.

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Some prior works [45, 46, 48, 49] indicated that a simple settling test (the measured height of a bed of sediment is compared to the liquid level) is useful for indicating flocculation. While the current work does observe other hallmarks of flocculation (measured IEP and increased yield stress), samples of the mixtures used here were allowed to settle for over two weeks with no visible sedimentation. The alumina suspension has been shown to be stable over months [61]. While settling velocities have been shown to be inversely correlated with solids concentration [63], the solids concentration of the slurry here is 20wt%, less than the 30vol% (~63wt%) for the slurries that were used for the sedimentation tests [45, 46, 48, 49]. However, the viscosity of the slurries studied here were much higher than reported in previous works, which can slow settling. Both the mixed (hercynite) suspension and the boehmite suspension alone both show isoelectric points (at pH values of 7-9 and 8, respectively) and associated peaks in yield stress. Thus, we attribute the current results to partial flocculation. Importantly, the behavior of the mixed hercynite slurry appears to be driven by the boehmite sol, meaning this type of sol could be very useful for future spray dried synthesis of any aluminum containing oxides for STWS.

3.6. Conclusions

This study shows a trade-off between spray dried particle robustness during redox cycling and particle activity. While candidate materials with added alumina and zirconia retained better individual particle structure after being heated to 1,200°C, they did not perform as well during redox cycling. Conversely, pure manganese oxide was unable to reliably form spray dried particles without the help of a colloidal suspension, and thus calcining the material produced mostly sintered agglomerates; however, these particles showed good redox activity. Manganese oxide to iron oxide in a molar ratio of 1:2 was able to form spherical spray dried particles which agglomerated and lost spherical shape during calcining, but also exhibited the best activity of any candidate material

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tested. This thermochemical performance was obtained over multiple cycles despite this apparent loss of surface area.

Zirconia appears to function as a stable solid support material at the temperatures tested, and did not form a solid solution with manganese oxide. This provides an opportunity for inclusion of zirconia as an inert support for higher activity materials. Alumina and iron oxide both formed mixed metal oxide spinels with manganese oxide after calcining. The addition of 30wt% alumina and 30wt% iron oxide was not beneficial to the redox activity of the materials. The difference in performance of the two iron-containing samples is attributed to the more rapid attainment of thermodynamic equilibrium before cycling. This in turn has been attributed to the difference in ionic diffusion rates of iron and manganese.

The candidate material Fe67, a manganese iron oxide spinel with a 1:2 Mn:Fe molar ratio, outperformed pure manganese oxide by a factor of 1.5, but does not retain spherical particle shape and undergoes sintering when heated to 1,200°C. However, this beneficial performance was shown over multiple cycles shown here that were done after agglomeration and sintering. Thus, future research should investigate the combination of this active material with robust, inert particle supports to find a balance between activity and physical stability while still maintaining higher activity than pure manganese oxide.

Spray dried hercynite particles have been successfully produced using a simple mixture and pH modification method. This type of sol preparation has been demonstrated in the past, but typically the suspension flocculated to a much higher degree. This shows that partial flocculation, especially using a common charge stabilized pseudo-boehmite sol, can be an effective method to produce solid spherical particles. The incorporation of iron oxide has not been used with this material previously, and indicates that the flocculation behavior of the boehmite sol seems to drive

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performance. This suggests that this system could be used flexibly to produce different mixed oxides based on an aluminate base.

Thermochemical reaction cycles will likely not depend on pure oxides alone, but rather on mixed or doped oxides. Thus, synthesis techniques that are able to combine multiple oxides in a scalable fashion are of critical importance for future work. The relatively straightforward mixture of two oxides thus illustrates an important effect for future scalable production. The fact that this simple synthesis technique of pH adjustment is able to be implemented successfully on a mixed oxide system illustrates a path forward for future synthesis at scale.

3.7. References

1. Bhosale, R.R., et al., Sol–gel derived CeO2–Fe2O3 nanoparticles: Synthesis, characterization and solar thermochemical application. Ceramics International, 2016. 42(6): p. 6728-6737.

2. Lichty, P., et al., Atomic layer deposited thin film metal oxides for fuel production in a solar cavity reactor. International Journal of Hydrogen Energy, 2012. 37(22): p. 16888-16894.

3. Ehrhart, B., et al., Thermochemical Cycle of a Mixed Metal Oxide for Augmentation of Thermal Energy Storage in Solid Particles. Energy Procedia, 2014. 49: p. 762-771.

4. Gokon, N., et al., New Solar Water-Splitting Reactor With Ferrite Particles in an Internally Circulating Fluidized Bed. Journal of Solar Energy Engineering, 2009. 131(1): p. 011007- 011007.

5. Koepf, E., et al., A novel beam-down, gravity-fed, solar thermochemical receiver/reactor for direct solid particle decomposition: Design, modeling, and experimentation. International Journal of Hydrogen Energy, 2012. 37(22): p. 16871-16887.

6. Ermanoski, I., N.P. Siegel, and E.B. Stechel, A New Reactor Concept for Efficient Solar- Thermochemical Fuel Production. Journal of Solar Energy Engineering, 2013. 135(3): p. 031002-031002-10.

7. Oles, A.S. and G.S. Jackson, Modeling of a concentrated-solar, falling-particle receiver for ceria reduction. Solar Energy, 2015. 122: p. 126-147.

8. Felinks, J., et al., Particle–particle heat transfer coefficient in a binary packed bed of alumina and zirconia-ceria particles. Applied Thermal Engineering, 2016. 101: p. 101-111.

9. Venstrom, L.J., et al., Applicability of an Equilibrium Model To Predict the Conversion of CO2 to CO via the Reduction and Oxidation of a Fixed Bed of Cerium Dioxide. Energy & Fuels, 2015. 29(12): p. 8168-8177.

152

10. Hildebrandt, A.F. and K.A. Rose, Receiver design considerations for solar central receiver hydrogen production. Solar Energy, 1985. 35(2): p. 199-206.

11. Imhof, A., The cyclone reactor — an atmospheric open solar reactor. Solar Energy Materials, 1991. 24(1–4): p. 733-741.

12. Felinks, J., et al., Heat recovery concept for thermochemical processes using a solid heat transfer medium. Applied Thermal Engineering, 2014. 73(1): p. 1004-1011.

13. Tregambi, C., et al., Heat transfer in directly irradiated fluidized beds. Solar Energy, 2016. 129: p. 85-100.

14. Stekli, J., L. Irwin, and R. Pitchumani, Technical Challenges and Opportunities for Concentrating Solar Power With Thermal Energy Storage. Journal of Thermal Science and Engineering Applications, 2013. 5(2): p. 021011-021011.

15. Pardo, P., et al., A review on high temperature thermochemical heat energy storage. Renewable and Sustainable Energy Reviews, 2014. 32: p. 591-610.

16. Wu, J. and X.f. Long, Research progress of solar thermochemical energy storage. International Journal of Energy Research, 2015. 39(7): p. 869-888.

17. Agrafiotis, C., M. Roeb, and C. Sattler, A review on solar thermal syngas production via redox pair-based water/carbon dioxide splitting thermochemical cycles. Renewable and Sustainable Energy Reviews, 2015. 42: p. 254-285.

18. Muhich, C.L., et al., A review and perspective of efficient hydrogen generation via solar thermal water splitting. Wiley Interdisciplinary Reviews: Energy and Environment, 2016. 5(3): p. 261-287.

19. Francis, T.M., P.R. Lichty, and A.W. Weimer, Manganese oxide dissociation kinetics for the Mn2O3 thermochemical water-splitting cycle. Part 1: Experimental. Chemical Engineering Science, 2010. 65(12): p. 3709-3717.

20. Arjmand, M., et al., Screening of Combined Mn-Fe-Si Oxygen Carriers for Chemical Looping with Oxygen Uncoupling (CLOU). Energy & Fuels, 2015. 29(3): p. 1868-1880.

21. Azimi, G., et al., Comprehensive study of Mn–Fe–Al oxygen-carriers for chemical-looping with oxygen uncoupling (CLOU). International Journal of Greenhouse Gas Control, 2015. 34: p. 12-24.

22. Perkins, C. and A.W. Weimer, Solar-thermal production of renewable hydrogen. AIChE Journal, 2009. 55(2): p. 286-293.

23. Kreider, P.B., et al., Manganese oxide based thermochemical hydrogen production cycle. International Journal of Hydrogen Energy, 2011. 36(12): p. 7028-7037.

153

24. Marugán, J., et al., Study of the hydrogen production step of the Mn2O3/MnO thermochemical cycle. International Journal of Hydrogen Energy, 2014. 39(10): p. 5274-5282.

25. Xu, B., Y. Bhawe, and M.E. Davis, Low-temperature, manganese oxide-based, thermochemical water splitting cycle. Proceedings of the National Academy of Sciences of the United States of America, 2012. 109(24): p. 9260-9264.

26. Alonso, E., et al., Kinetics of Mn2O3–Mn3O4 and Mn3O4–MnO Redox Reactions Performed under Concentrated Thermal Radiative Flux. Energy & Fuels, 2013. 27(8): p. 4884-4890.

27. Wong, B., et al., Oxide based thermochemical heat storage, in SolarPACES. 2010: Perpignan, France.

28. Carrillo, A.J., et al., Thermochemical energy storage at high temperature via redox cycles of Mn and Co oxides: Pure oxides versus mixed ones. Solar Energy Materials and Solar Cells, 2014. 123: p. 47-57.

29. Carrillo, A.J., et al., Thermochemical heat storage based on the Mn2O3/Mn3O4 redox couple: influence of the initial particle size on the morphological evolution and cyclability. Journal of Materials Chemistry A, 2014. 2(45): p. 19435-19443.

30. Wong, B. and Project Staff, Thermochemical Heat Storage for Concentrated Solar Power: Thermochemical System Reactor Design for Thermal Energy Storage. 2011, General Atomics: Report No. GA-C27137.

31. Babiniec, S.M., et al., Investigation of LaxSr1−xCoyM1−yO3−δ (M = Mn, Fe) perovskite materials as thermochemical energy storage media. Solar Energy, 2015. 118: p. 451-459.

32. Babiniec, S.M., et al., Doped calcium manganites for advanced high-temperature thermochemical energy storage. International Journal of Energy Research, 2016. 40(2): p. 280-284.

33. Carrillo, A.J., et al., Improving the Thermochemical Energy Storage Performance of the Mn2O3/Mn3O4 Redox Couple by the Incorporation of Iron. ChemSusChem, 2015. 8(11): p. 1947-1954.

34. Azimi, G., et al., Chemical-looping with oxygen uncoupling using combined Mn-Fe oxides, testing in batch fluidized bed. Energy Procedia, 2011. 4: p. 370-377.

35. Adanez, J., et al., Progress in Chemical-Looping Combustion and Reforming technologies. Progress in Energy and Combustion Science, 2012. 38(2): p. 215-282.

36. Hallberg, P., et al., Chemical Looping Combustion and Chemical Looping with Oxygen Uncoupling Experiments in a Batch Reactor Using Spray-Dried CaMn1–xMxO3−δ (M = Ti, Fe, Mg) Particles as Oxygen Carriers. Energy & Fuels, 2013. 27(3): p. 1473-1481.

154

37. Linderholm, C., T. Mattisson, and A. Lyngfelt, Long-term integrity testing of spray-dried particles in a 10-kW chemical-looping combustor using natural gas as fuel. Fuel, 2009. 88(11): p. 2083-2096.

38. Walton, D.E. and C.J. Mumford, Spray Dried Products—Characterization of Particle Morphology. Chemical Engineering Research and Design, 1999. 77(1): p. 21-38.

39. Liu, B., et al., Fluidization of non-spherical particles: Sphericity, Zingg factor and other fluidization parameters. Particuology, 2008. 6(2): p. 125-129.

40. Nandiyanto, A.B.D. and K. Okuyama, Progress in developing spray-drying methods for the production of controlled morphology particles: From the nanometer to submicrometer size ranges. Advanced Powder Technology, 2011. 22(1): p. 1-19.

41. Lukasiewicz, S.J., Spray-Drying Ceramic Powders. Journal of the American Ceramic Society, 1989. 72(4): p. 617-624.

42. Jr. Walker, W.J., J.S. Reed, and S.K. Verma, Influence of Slurry Parameters on the Characteristics of Spray-Dried Granules. Journal of the American Ceramic Society, 1999. 82(7): p. 1711-1719.

43. Cao, X.Q., et al., Spray-drying of ceramics for plasma-spray coating. Journal of the European Ceramic Society, 2000. 20(14–15): p. 2433-2439.

44. Tsetsekou, A., et al., Optimization of the rheological properties of alumina slurries for ceramic processing applications Part II: Spray-drying. Journal of the European Ceramic Society, 2001. 21(4): p. 493-506.

45. Bertrand, G., et al., Influence of slurry characteristics on the morphology of spray-dried alumina powders. Journal of the European Ceramic Society, 2003. 23(2): p. 263-271.

46. Bertrand, G., et al., Spray-dried ceramic powders: A quantitative correlation between slurry characteristics and shapes of the granules. Chemical Engineering Science, 2005. 60(1): p. 95-102.

47. Garnier, J., A thermodynamic study on CaO-doped nonstoichiometric cerium dioxide. Journal of Physics and Chemistry of Solids, 1976. 37(4): p. 368-378.

48. Mahdjoub, H., et al., The effect of the slurry formulation upon the morphology of spray-dried yttria stabilised zirconia particles. Journal of the European Ceramic Society, 2003. 23(10): p. 1637-1648.

49. Roy, P., G. Bertrand, and C. Coddet, Spray drying and sintering of zirconia based hollow powders. Powder Technology, 2005. 157(1–3): p. 20-26.

50. Eslamian, M. and N. Ashgriz, Effect of precursor, ambient pressure, and temperature on the morphology, crystallinity, and decomposition of powders prepared by spray pyrolysis and drying. Powder Technology, 2006. 167(3): p. 149-159.

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51. Naglieri, V., et al., Optimized Slurries for Spray Drying: Different Approaches to Obtain Homogeneous and Deformable Alumina-Zirconia Granules. Materials, 2013. 6(11): p. 5382.

52. Iskandar, F., L. Gradon, and K. Okuyama, Control of the morphology of nanostructured particles prepared by the spray drying of a nanoparticle sol. Journal of Colloid and Interface Science, 2003. 265(2): p. 296-303.

53. Bale, C.W., et al., FactSage thermochemical software and databases — recent developments. Calphad, 2009. 33(2): p. 295-311.

54. Eriksson, G. and K. Hack, ChemSage—A computer program for the calculation of complex chemical equilibria. Metallurgical Transactions B, 1990. 21(6): p. 1013-1023.

55. Kanerva, U., et al., Evaluation of crushing strength of spray-dried MgAl2O4 granule beds. Ceramics International, 2015. 41(7): p. 8494-8500.

56. Srinivasan, R., et al., Identification of tetragonal and cubic structures of zirconia using synchrotron x-radiation source. Journal of Materials Research, 1991. 6(6): p. 1287-1292.

57. Gilewicz-Wolter, J., et al., Diffusion of chromium, manganese, and iron in MnCr2O4 spinel. Journal of Phase Equilibria and Diffusion, 2005. 26(5): p. 561-564.

58. Baek, J.-I., et al., Performance Comparison of Spray-Dried Oxygen Carriers: The Effect of NiO and Pseudoboehmite Content in Raw Materials. Energy & Fuels, 2010. 24(10): p. 5757- 5764.

59. Liu, H., et al., A novel FCC catalyst synthesized via in situ overgrowth of NaY zeolite on kaolin microspheres for maximizing propylene yield. Catalysis Today, 2007. 125(3–4): p. 163-168.

60. Daniell, W., et al., Enhanced surface acidity in mixed alumina–silicas: a low-temperature FTIR study. Applied Catalysis A: General, 2000. 196(2): p. 247-260.

61. Caiut, J.M.A., et al., Elaboration of boehmite nano-powders by spray-pyrolysis. Powder Technology, 2009. 190(1–2): p. 95-98.

62. Albano, M.a.P. and L.B. Garrido, Rheological properties of Si3N4, pseudoboehmite and bayerite coated Si3N4 suspensions with ammonium polyacrylate dispersant. Colloids and Surfaces A: Physicochemical and Engineering Aspects, 2002. 203(1–3): p. 117-128.

63. Davis, R.H. and K.H. Birdsell, Hindered settling of semidilute monodisperse and polydisperse suspensions. AIChE Journal, 1988. 34(1): p. 123-129.

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CHAPTER IV

ROLE OF METAL IONS IN HERCYNITE SPINEL

4.1. Abstract

Cobalt-doped hercynite materials are examined for two-step solar thermochemical water splitting. Computational simulations have been performed using DFT+U for iron and cobalt aluminate. These calculations predict that electron density transfer on the creation of oxygen vacancies is almost exclusively to the atoms that are first-nearest-neighbor to the oxygen vacancy.

Cases that have a single cobalt next to the oxygen vacancies tend to be very favorable both in terms of a lower oxygen vacancy formation energy and also a lower host structure energy, and cobalt appears to preferentially change oxidation state on formation of oxygen vacancies. However, sites with more than one Co next to the O-vacancy tend to be less favorable, due to cobalt being less favorable on the octahedral sites. Current XPS results seem to contradict the computational results, as the Co-doped hercynite sample does not show Co oxidation state changes while the un-doped hercynite shows Fe oxidation state changes. This is attributable to the excess cobalt in the Co- doped sample; more Co make oxygen vacancies less likely to form. Hercynite samples with less cobalt should show more oxidation state change for the Co-doped sample. Cobalt-doped hercynite has been showed to have faster reaction rates but less hydrogen production capacity compared to the un-doped hercynite. This might be due to kinetic/catalytic impacts of Co on the reaction mechanism, but can also be at least partially affected by the thermodynamics of the reaction enthalpy change as well. This work suggests that the hercynite cycle can be improved by lowering the amount of Co in the material or by introducing new dopants that can assume the octahedral site more easily.

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4.2. Introduction

Solar thermochemical water splitting (STWS) has the potential to efficiency produce renewable and carbon-free hydrogen. This chapter performs an examination of elemental doping of a promising STWS material in order to obtain a more complete understanding of how dopants may behave in a broader class of materials.

4.2.1. Doping in Solar Thermochemical Water Splitting Materials

Doping elements into host oxides that are primarily composed of another element has been used extensively for tuning material properties in general, and for tuning the performance of STWS materials in particular. Many different elements have been attempted as dopants for ceria, which is considered the best current oxygen vacancy-driven STWS material. Some di- and trivalent elements have been doped into ceria to replace the tetravalent cerium in order to reduce the enthalpy required to reduce the material, thus making it able to be reduced at lower temperatures

[1-3]. However, many of these dopants lowered the reduction enthalpy too much, making the reduction of the material go forward but leaving the reduced oxide unable to re-oxidize with water

[1]. Dopants that have a different valence level than the host metal will generate permanent oxygen vacancies in the material in order to maintain charge neutrality, but these oxygen vacancies do not contribute toward STWS.

However, other dopants that differ in valence from the host have shown very promising results with ceria and perovskites. Doping strontium into the lanthanum-site in perovskites has been shown to increase H2 production capacity at the cost of reducing the reaction rate [4] . This has been explained by computation results that show the creation of un-occupied states in the band gap of the perovskite material on doping with Sr, lowering the energy required for oxygen vacancy formation [5]. Additionally, uranium-doped ceria has shown very promising STWS performance

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[6, 7]. This increase in water splitting behavior is attributed to computationally predicted charge transfer from the multivalent U to Ce, making reduction of both U and Ce more energetically favorable [8]; this is consistent with experimental results of U and Ce oxidation states [7]. This shows the benefit of dopants that can take multiple oxidation states.

Other dopants that have been attempted for ceria are tetravalent, meaning they do not create the same permanent oxygen vacancies as dopants with different valences. Instead, these elements form metal ions with the same charge but different ionic radii than the host cerium. If the dopant has a smaller ionic radius than the host ion, it introduces a tensile lattice strain into the material.

This can counteract the compressive lattice strain introduced by oxygen vacancy formation; the lattice can “relax” as an oxygen atom is removed. Different ionic radii can thus affect ceria differently, allowing for more fine-tuning of the net energy required for reduction. This has been demonstrated for Zr and Hf dopants into ceria [9, 10]; the ionic radii of Zr4+, Hf4+, and Ce4+ are

0.84, 0.83, and 0.97 Å, respectively [10, 11]. However, it should be noted that these dopants do have an effect on the rate of reaction of ceria, as they slow the oxidation rate [9, 10]. Thus, dopants need to consider both thermodynamic capacity as well as kinetic effects.

Other elements can lower the production capacity of a material but provide other benefits, increasing the overall performance of a material. Recent investigations of various perovskite-type materials have shown that perovskites made up of strontium, lanthanum, and manganese (SLM) can produce more H2 than ceria at lower temperatures [12]. Similar perovskites that have aluminum added (SLMA) also show higher H2 production than ceria at lower temperatures, but suffer from less sintering than the SLM perovskites [13]. This indicates that the aluminum ions can stabilize the perovskite structure and allow for oxygen vacancy formation without phase change and with less sintering. This is conceptually similar to doping Al into the B-site of iron

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oxide spinel (Fe3O4) to form hercynite (FeAl2O4); this also lowers the H2 production capacity but leads to a material which sinters significantly less than pure iron oxide.

4.2.2. Hercynite Cycle History

The hercynite cycle was originally suggested by Scheffe et al. They discovered that cobalt ferrite deposited on alumina would produce H2 during steam oxidation after being reduced at

1,200°C, which is 200-300°C lower than cobalt ferrite on a zirconia support [14]. This reaction was hypothesized to operate through a stoichiometric mechanism, where the CoFe2O4 reacts with the Al2O3 support during reduction to produce a solid solution of FeAl2O4 (hercynite) and

CoAl2O4: CoFe2O4 + 3Al2O3  CoAl2O4 + 2FeAl2O4 + ½ O2. This solid solution was expected to then revert back to two separate phases (CoFe2O4 spinel and Al2O3 corundum) on oxidation by steam [14]. More recently, Muhich et al. have proposed that the hercynite cycle actually appears to operate through an oxygen vacancy mechanism [15]. Computational simulations were used to suggest that the reduction enthalpy of the stoichiometric reaction was not energetic enough to produce H2 on re-oxidation. This was consistent with demonstrated H2 production experiments as well as in-situ high temperature XRD, which showed lattice constant changes on reduction and oxidation but no observable phase changes [15]. Furthermore, Muhich et al. predicted that un- doped hercynite (FeAl2O4) would have higher H2 production capacity than Co-doped hercynite

(CoxFe1-xAl2O4). This was experimentally confirmed, though it was shown that the Co-doped hercynite had faster reaction rates for H2 production [15].

While the computational predictions of Muhich et al. were correct in their predictions of

H2 capacity, it was not well understood what specific role the cobalt played in doping the material.

The effect on oxidation kinetics was hypothesized to be due to some kind of surface catalytic effect

[15], but more work is needed to understand this fully. The prediction of H2 production capacity

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was based on energetic favorability of oxygen sites being surrounded by 4 nearest neighbors that were either Fe or Co instead of Al [15]. This is an important finding, and provides a way to estimate

H2 production capacity based on similar materials by calculating O-vacancy formation energies and estimating the number of favorable active sites. However, a deeper understanding of why Fe,

Co, and Al behave in this way could lead to further improvements in material discovery either through doping or selection of a new host material.

This work uses density functional theory to further explore differences between hercynite

(FeAl2O4), cobalt aluminate (CoAl2O4), and Co-doped hercynite (CoxFe1-xAl2O4). Specifically, charge density transfer on formation of oxygen vacancies is examined as a way to determine what role each metal ion plays in the material. This is then compared to analytical results in which samples of these materials are characterized with X-ray photoelectron spectroscopy (XPS) while being heated to high temperatures in vacuum. This in-situ analysis allows for the determination of oxidation state changes in real samples as they reduce, allowing for electron density charge transfer to be experimental observed directly in order to verify computational predictions.

4.3. Method

Experimental results are compared to various analytical methods of the sample powders and to theoretical simulations. The computational simulations will be described first, followed by the experimental and analytical methods used.

4.3.1. Computational Simulations

The Vienna Ab-initio Simulation Package (VASP) was used to perform plane wave periodic boundary conditions density functional theory (DFT) simulations [16, 17]. These calculations used the Perdew-Burke-Ernzerhof (PBE) generalized gradient approximation (GGA) exchange and correlation functional [18]. This was coupled with projector augmented wave

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(PAW) pseudopotentials [19] for the explicit description of the 4s and 3d Co orbitals, 4s and 3d

Fe orbitals, 2s and 2p O orbitals, and 2s and 2p Al orbitals. The degrees of freedom for electron spin were treated explicitly. A limited search for the ground-state magnetic configuration was performed, which included system initialization in ferromagnetic, anti-ferromagnetic, and multiple paramagnetic configurations for un-doped hercynite structures. Spin sampling in the un-doped structures indicated that the trends for charge transfer and oxygen vacancy formation energy were qualitatively similar regardless of spin state; therefore, for Co-doped hercynite structures, only the ferromagnetic structures were calculated in order to reduce computational requirements.

Simulations were calculated using a 3x3x3 k-point mesh, centered on the Gamma-point. A 500 eV plane wave energy cutoff was used, as determined by a convergence analysis. All total energies converged to below 1x10-5 eV. Hubbard on-site correction terms (U) were used to treat correlation of 3d electrons [20]. A value of 3 eV was used for U for both Co and Fe as a compromise between oxidation energy and bandgap based on literature [21, 22]. A 112 atom supercell was used, corresponding to a 2x2x2 expansion of the primitive cell. In a “normal” spinel structure, A2+ atoms occupy tetrahedral sites, and B3+ atoms occupy octahedral sites. In an “inverse” spinel structure, half of the B3+ ions occupy tetrahedral sites while the other half of the B3+ ions and all of the A2+ ions occupy octahedral sites.

Oxygen vacancy structures were generated by the removal of one neutral O atom from the

112 atom structure. In the case of a normal spinel, the oxygen vacancy is surrounded by three Al atoms in octahedral sites and one A atom on a tetrahedral site, while a fully inverse spinel has an oxygen vacancy surrounded by four A atoms and zero Al atoms. Lattice parameters and atomic positions were optimized in both the defect and host structure. Non-equivalent O sites were sampled and the lowest defect site was used in subsequent calculations. The O-vacancy formation

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defect energy (EV) was calculated as per Equation (4.60), in which 퐸tot is the total energy of the defect

host FERE structure, 퐸tot is the total energy of the host structure, and 휇O is the standard state O chemical potential. Bader charge analysis was conducted using software from the Henkelman group [23,

24].

푑푒푓푒푐푡 ℎ표푠푡 퐹퐸푅퐸 (4.60) 퐸푉 = 퐸푡표푡 − 퐸푡표푡 + 휇푂

4.3.2. Experimental Characterization

Hercynite particles were made via spray drying. Hercynite slurries are prepared by combining purchased commercial suspensions of aluminum oxide (NYACOL Nano Technology,

AL20DW, Colloidal Alumina Sol) and iron oxide (Aldrich, 720704, Iron(III) oxide, dispersion).

The appropriate amount of each suspension was weighed out and combined in a beaker, which was then stirred for 20 minutes on a stir plate at 400 rpm. 1M NaOH (2.5M for hercynite slurries) or 1M HCl were added dropwise to the slurry to affect the pH of the suspension, which was measured with a digital pH meter. After each drop, the suspension was allowed to settle and stir for 2 minutes before the pH was measured again. This process continued until the desired pH of 7 was achieved for each sample. Water was added as needed to make slurry able to spray. Spray- dried metal oxide particles are made using a Büchi B-290 Mini Spray Dryer. The slurry is fed into the nozzle of the spray dryer via a peristaltic pump where it is nebulized with compressed air to form droplets. The liquid droplets are entrained in a flow of hot air (~200C) in the drying chamber.

Particles form as the droplets dry, and are collected and separated from ultra-fines using a cyclone.

After spray drying, particles are calcined in a box furnace under ambient atmosphere at 1,200ºC for 8 hours. This was done to form the thermodynamically favorable mixed metal oxide phases, and to sinter the particles prior to analysis.

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XPS was performed in a separate UHV system with a base pressure of 1×10−9 mbar equipped with SPECS XR50 dual anode X-ray source (Mg Kα was utilized) and Scienta R3000 hemispherical electrostatic energy analyzer. The system is also equipped with an E-beam heater

HEAT3-PS by PREVAC, that was utilized to heat the samples to the desired temperatures. The temperature of the samples was monitored using Sirius pyrometer by Process Sensors and a calibrated K-type thermocouple. The samples were pressed into pellets (10mm diameter) and loaded on a tantalum plate. The XPS data analysis was performed using THERMO AVANTAGE software by Thermo Fisher Scientific. XPS spectra of Fe 2p, Co 2p, Al 2p, C 1s and O 1s were collected. Typical acquisition conditions were a pass energy of 50 eV and scan rate of 0.1 eV per

200 ms.

The H2 production capacity was measured in an in-house made SFR with a MKS Cirrus2 mass spectrometer with a capillary sampling port. The active materials were allowed to reduce for

60 min at 1,500°C and a total pressure of 760 Torr with a 300 sccm flowrate of He to swept the generated O2 out of the reactor. Oxidation occurred for 15 min at 1,350°C and a 760 Torr total pressure with a 300 sccm 50% H2O/50% He gas flow. Steam was fed using a RSIRC RainMaker steam feeder. Blank runs with no active material present in the reactor were conducted and the H2 production from heterolytic water splitting was subtracted from the sample runs.

4.4. Results

Results and insights from computational simulations will be described first, followed by experimental characterization of the sample materials. This will then be followed by a discussion of how the computational and experimental results combine to provide a more complete picture of how the hercynite spinels react.

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4.4.1. Computational Results

The oxygen vacancy formation energy has been calculated for pure iron and cobalt oxides as well as cobalt aluminate, iron aluminate, and cobalt-doped iron aluminate. All oxides considered are in the spinel structure, making direct comparison possible. The pure oxides (Co3O4 and Fe3O4) are much more easily reduced, with the aluminate O-vacancy value nearly 3X the value of the pure oxides for a single vacancy as shown in Figure 4.1.

Figure 4.1: Calculated oxygen vacancy formation energy for pure oxides and aluminates for both cobalt and iron Computational simulations were also used to determine the effects of charge transfer on formation of an oxygen vacancy. Calculations were performed for un-doped hercynite (FeAl2O4),

CoAl2O4, and hercynite doped with a single Co atom adjacent to the oxygen vacancy. These results are shown in Figure 4.2, showing the final positions of atoms as well as iso-surfaces to show charge density change on creation of the oxygen vacancy. Charge density is transferred away from the O- vacancy, and is transferred to the nearest neighbor A-site atom. This is true for both the doped and un-doped cases.

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Figure 4.2: Structure of un-doped a) FeAl2O4, b) CoAl2O4, and FeAl2O4 doped with a single Co. All images display oxygen atoms as red, aluminum atoms as light blue, iron atoms as gold, and cobalt atoms as dark blue. In each image, charge density change iso-surfaces are shown as semi- transparent to illustrate charge transfer: the teal iso-surface denotes electron density loss, while the yellow iso-surface shows electron density gain. Bader charge analysis is used to quantify the effect of the results of charge transfer for un- doped and Co-doped hercynite, and this is shown in Figure 4.3. The same general trends are shown for each case, in which the charge is primarily transferred to the first nearest neighbor A-site atom.

The Bader charge analysis assigns some charge transfer to nearby B-site atoms, but this is not reflected in Figure 4.2. Therefore, this is attributed to the charge density transfer being assigned to new volume segments after the O-vacancy is created.

Figure 4.3: Net charge transfer for the 30 closet nearest neighbors (NN) of the oxygen vacancy. Each atom is labeled by the location relative to the oxygen vacancy, as well as the type (A or B) of the metal ion site. Bader charge analysis was used to assign charge to specific atoms. In addition to normal spinel structures, inverse spinel structures were also simulated. This was done by changing the local inversion by placing different numbers of Al atoms (on the

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octahedral site in a normal spinel) onto tetrahedral sites near the oxygen vacancy. The oxygen vacancy formation energy was calculated for all different possible local environments around the oxygen vacancy. In addition, the relative structure energy was calculated for each of the three broad material categories (FeAl2O4, CoAl2O4, and (Co,Fe)Al2O4) by subtracting the minimum total energy for the perfect structure from each material category from the rest of the structures in that category. To account for different number of dopant atoms in this calculation, the change in atomic chemical potential for substitution of an iron atom with a cobalt atom is added to the relative structure energy. The results of these calculations are shown in Figure 4.4 for all structures considered here. In general, these results show lower O-vacancy formation energies when there are fewer Al atoms next to the O-vacancy. However, these structures also tend to have higher relative structure energies, which indicates they may be less likely to form and thus be available for oxygen vacancy formation.

Figure 4.4: Calculated oxygen vacation formation energy and relative structure energy values.

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Example structures for these inverse spinel calculations are shown in Figure 4.5. The charge from a formed oxygen vacancy is primarily transferred to Co and Fe first-nearest- neighbors, which show the same general trends as above.

Figure 4.5: Structure of a) FeAl2O4, b) CoAl2O4, and FeAl2O4 doped with a single Co. Each of the above images are for a 50% inverse structure, in which 2 Al atoms are on tetrahedral sites that are 1st nearest neighbor to the O-vacancy. All images display oxygen atoms as red, aluminum atoms as light blue, iron atoms as gold, and cobalt atoms as dark blue. In each image, charge density change iso-surfaces are shown as semi-transparent to illustrate charge transfer: the teal iso-surface denotes electron density loss, while the yellow iso-surface shows electron density gain. The inverse spinel structure was also considered, and generally the same trends are followed for observed charge density transfer. Figure 4.6 shows the charge difference before and after the creation of an oxygen vacancy for the 30 nearest neighbors. While the elements in each position may differ, the charge is almost exclusively transferred to the four nearest neighbors of the vacancy. This is true for cobalt aluminate, iron aluminate, and Co-doped iron aluminate.

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Figure 4.6: Net charge transfer for inverse spinel structures for the 30 closet nearest neighbors (NN) of the oxygen vacancy. Each atom sorted by the location relative to the oxygen vacancy. Bader charge analysis was used to assign charge to specific atoms. The structures used for the net charge transfer (shown in Figure 4.6) is shown in Figure 4.7 with iso-surfaces to denote charge density gains and losses. As the Bader charge analysis quantified, the charge transfer is focused on the four nearest neighbors to the oxygen vacancy. This is true even though dipoles seem to induced in those and some other nearby atoms.

Figure 4.7: Structures of a) FeAl2O4, b) CoAl2O4, and Co-doped FeAl2O4. Each of the above images are for a fully inverse structure, in which 0 Al atoms are 1st nearest neighbor to the O- vacancy. CoAl2O4 is an exception, where 1 Al is nearest neighbor. All images display oxygen atoms as red, aluminum atoms as light blue, iron atoms as gold, and cobalt atoms as dark blue. In each image, charge density change iso-surfaces are shown as semi-transparent to illustrate charge transfer: the teal iso-surface denotes electron density loss, while the yellow iso-surface shows electron density gain. The visualization and quantification of charge density transfer is useful to observe the effects of oxygen vacancy formation, but a more detailed examination of the role of each atom

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type can show a more completely picture. Projected density of states (DOS) were calculated and plotted in Figure 4.8 for both host and oxygen vacancy structures. The host un-doped hercynite structure (Figure 4.8a) shows a bandgap of ~2 eV, and the highest occupied molecular orbital

(HOMO) and lowest un-occupied molecular orbital (LUMO) states immediately on either side of this gap are dominated by states from iron orbitals. After the creation of an oxygen vacancy in the same structure (Figure 4.8b), mid-gap states are created immediately below the valence band. The aluminum and oxygen orbitals are relatively un-changed between these two simulations, which supports the fact that charge is mostly transferred to Fe on oxygen vacancy creation.

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Figure 4.8: Projected density of states for a) FeAl2O4, b) FeAl2O4 after O-vacancy creation, c) Co-doped FeAl2O4, d) Co-doped FeAl2O4 after O-vacancy creation, and e) Co-doped FeAl2O4 after O-vacancy creation with the doped Co as a 2nd-nearest neighbor instead of 1st. All simulations are for 2 Al first nearest neighbors. Energy values on the horizontal axis are relative to the Fermi energy for each system. Positive and negative vertical axis values are for spin-up and spin-down states, respectively. Density of states are scaled using different values for ease of viewing, and are not meant to be directly additive to obtain the total DOS.

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The DOS for the host Co-doped hercynite (Figure 4.8c) is very similar to the un-doped structure, with one important difference; the presence of additional states from cobalt at the bottom of the valence band, lowering the band gap slightly. However, the DOS for the Co-doped structure after the creation of the oxygen vacancy (Figure 4.8d) has two important differences than the un- doped hercynite oxygen vacancy structure (Figure 4.8b). Significant mid-gap states are created, this time driven by cobalt instead of iron. Importantly, these states extend further into the band gap, leading to a significantly smaller band gap in this reduced material. The other important difference is the valence band states created immediately at and below the Fermi energy level.

These states are driven by cobalt, and the iron states that appeared in the un-doped structure are shifted to lower energies (relative to the Fermi energy). This suggests that the cobalt conduction band states in the host structure provide a slightly easier place for electrons to occupy when the oxygen vacancy is created. As the oxygen vacancy structure relaxes, these cobalt states stay filled and are transferred to the valence band. Further, the cobalt appears to stabilize the iron by shifting the iron states in the valence band to lower relative energies.

Finally, a simulation is shown for Co-doped hercynite but for when the cobalt atom is not a first nearest neighbor to the oxygen vacancy. This is shown in Figure 4.8e, and shows a striking similarity to the un-doped oxygen vacancy structure (Figure 4.8b) more than the Co-doped structure (Figure 4.8d). The same mid-gap LUMO iron states just below the conduction band are shown, and the cobalt states do not appear to be in the HOMO. This suggests that the cobalt dopant must be directly adjacent (first nearest neighbor) to created oxygen vacancies to have a significant impact.

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4.4.2. Particle Characterization

X-ray photoelectron spectroscopy (XPS) was performed on the spray dried hercynite samples. The samples were heated to various temperatures of interest and measured in-situ, and the results are shown in Figure 4.9. The XPS measured results are fit to various component curves which can be attributed to various metal oxidation states. The Fe2p spectra for un-doped hercynite

(Figure 4.9a) show that on heating to 1,400°C, the hercynite undergoes reduction. This is especially visible in the formation of the peak attributed to Fe0 on the leading right-hand edge of the figure. Results for Co-doped hercynite (Figure 4.9b and c) show a change in oxidation state on heating from 25 to 1,300°C, but after this the results diverge. The Fe2p component peaks continue to shift oxidation states (Figure 4.9b) while the Co oxidation state appears to stay constant with further increases to temperature (Figure 4.9c).

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Figure 4.9: XPS measurements with fitted component curves for a) Fe2p in un-doped hercynite, b) Fe2p for Co-doped hercynite, and c) Co2p for Co-doped hercynite. Hydrogen and oxygen production were measured in a stagnation flow reactor (SFR) and the average production over multiple cycles are shown in Figure 4.10 for both un-doped and Co- doped hercynite. The Co-doped hercynite has significantly less production in both H2 and O2 than the un-doped hercynite.

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Figure 4.10: Hydrogen and oxygen production values per cycle measured in stagnation flow reactor. The average production per unit mass of solid sample is shown for the second to the twelfth cycle, as typically the first cycle is much different. Error bars show the 95% confidence interval. 4.5. Discussion

4.5.1. Differences Between Oxides and Aluminates

The focus of this work is on aluminate spinels, but a direct comparison to pure oxides does help inform this choice. Many pure binary oxides have been studied for STWS, but no ideal material has been identified. Some suffer from operational issues, such as volatile or molten phases, while others simply do not undergo the reduction/oxidation cycle at reasonable conditions

[1]. The results in Figure 4.1 can help to explain this trend, at least for iron and cobalt oxides. The oxygen vacancy formation energy values for pure iron and cobalt oxides are much lower than for the corresponding iron and cobalt aluminates. This is consistent with other results that show that the Fe or Co atom is much harder to change oxidation state when in an aluminate than when a pure

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oxide [25, 26]. This is also consistent with materials (such as manganese or cobalt oxide) being able to reduce at lower temperature, but not being able to oxidize with water [27]. This has been attributed to the enthalpy of reaction, which is directly related to the oxygen vacancy formation energy; the enthalpy of the reduction reaction must be high enough to also drive the exothermic oxidation reaction to complete the cycle [15].

4.5.2. Preferential Charge Transfer On Aluminate Oxygen Vacancy Formation

The simulations reported in Section 4.4.1 show that the oxygen vacancy formation causes a transfer of charge density to specific nearby metal ions. This is certainly related to the oxidation state, and this is reflected in the Bader charge analysis performed. An intuitive result is the fact that no significant charge is transferred to the aluminum atoms. Aluminum atoms are known to form very stable bonds with oxygen, and do not tend to form oxides based on multiple oxidation states. By contrast, there are many different iron and cobalt oxides for these metals in the 2+ and

3+ oxidations states. This indicates that the aluminum is much less likely to change oxidation state, especially in a point-defect fashion without a corresponding phase change.

Other results are not so intuitive. The computational predictions show that charge transfer tends to occur on both the iron and cobalt atoms fairly consistently: significant iso-surfaces appear next to both iron and cobalt in Figure 4.5 and Figure 4.7, and Figure 4.8 shows cobalt orbital states play a significant role in oxygen vacancy formation. However, the XPS results (Figure 4.9) show that the Fe oxidation state changes significantly more than the Co oxidation state. This apparent contradiction could be explained by the fact that the computational simulations focused on an infinite structure, analogous to the bulk of the material, whereas XPS is a surface sensitive technique. However, XPS tends to be sensitive to penetration depths on the order of nanometers, whereas the computational cell is only ~11.6 Å deep; this difference in distance of orders of

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magnitude means that most of the surface sampled by XPS should be fairly accurately reflected in the “bulk” computations. Thus, the surface sensitivity of XPS is likely not the main driver of this difference.

Many different local environments were sampled for the oxygen vacancy formation energy. However, all of these resulting structures are not equivalent; some are more stable and therefore likely to form than others. In order to illustrate the total effect of both starting structure relative energies and the energy required to form oxygen vacancies, the values from Figure 4.4 are added together, sorted, and displayed in Figure 4.11. It should be noted that this summation is not a direct thermodynamic energy cost, but rather an illustration of general trends of likelihood of formation of a given vacancy (with respect to prevalence of sites) and ease of formation of a vacancy at those sites. As can be seen, considering a low oxygen vacancy formation energy alone may be misleading, as some of the structures to the right of the figure have very low O-vacancy formation energy values, but the relative structure energy makes these sites unlikely to form in the first place. Conversely, some of the moderate-to-high values of O-vacancy formation energy have the lowest total bar height, as the relative structure energy is lower for some of these.

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Figure 4.11: Calculated oxygen vacation formation energy and relative structure energy values, sorted by sum of total energy for each case. Considering both the oxygen vacancy formation energy and the relative structure energy provides a more complete picture of the likelihood of different oxygen vacancies to form and how this behavior can explain the observed XPS results. The lower energy oxygen vacancies (left side of Figure 4.11) are likely to form first and most often in a given material. While these lower energy vacancies do almost all occur with cobalt present, none of them have more than a single cobalt and so charge transfer is also likely to go to these nearby Fe atoms. For example, the second-lowest energy vacancy is for un-doped hercynite. The computational results in Section 4.4.1 show that oxygen vacancies with a single Co dopant transfers charge (and thus fills states) of primarily cobalt rather than iron. However, Figure 4.11 also shows that vacancies with multiple cobalt first nearest neighbors tend to be much higher energy. Thus, for a sample with a large amount of cobalt relative to iron is more likely to show changes in the Fe2p peaks rather than the Co2p peaks.

Other important trends can be identified in Figure 4.11, such as the fact that the oxygen vacancies for zero or one Al nearest neighbor are all on the right side of the plot (higher energy).

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This is despite the fact that electron density transfer only appears to go to transition metals (Co and Fe) rather than Al, which might suggest that as more of those transition metals are nearby it would be more stable to create an oxygen vacancy due to more available locations for electron transfer. However, this suggests that instead, the Al stabilizes the material during oxygen vacancy formation. Pure transition metal oxides might be more likely to change phase, while the aluminum in these materials can hold the structure as oxygen vacancies are created and destroyed.

4.5.3. Octahedral and Tetragonal Site Energy Penalties

An important aspect of these results is the reason why small amounts of cobalt are favorable for oxygen vacancies, but larger amounts are not. This can be explained by the fact that as more cobalt are nearest neighbors of potential oxygen vacancies, they must eventually start occupying octahedral sites instead of tetrahedral. This is much less favorable for cobalt than for iron due to the degeneracy of d-orbitals and the number of valence electrons for cobalt and iron. Crystal field theory describes how d-orbitals will form into two different degenerate energy levels when exposed to a non-spherical coordination [28]. An octahedral geometry will form three degenerate energy levels (collectively referred to as t2g) that are at a lower energy than two other degenerate energy levels (eg), whereas a tetrahedral geometry will have two lower energy levels (e) and three degenerate higher energy levels (t2). When the material is in a low spin state, the difference in energy between the higher and lower energy states (Δoct and Δtet for octahedral and tetrahedral, respectively) is large, so electrons fill and pair in the low energy states first before any higher energy states are filled. Conversely, in a high spin state, the energy difference is small, meaning that one electron will fill each orbital first before pairing (consistent with Hund’s rule).

These degenerate energy states have an important bearing on cobalt and iron in particular.

Fe has 6 valence d-electrons while Co has 7. For an octahedral configuration, there can be a

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maximum of 6 paired electrons in the triply degenerate t2g states, meaning that Fe can fill only t2g states in the low spin state, whereas Co must have one electron in the higher energy eg states. This means that Co will suffer a higher energy penalty to be in a octahedral site than Fe. Conversely, the high or low spin state does not matter very much for Co or Fe in a tetrahedral configuration, because a maximum of 2 un-paired or 4 paired electrons can be in the lower energy e states; this means that electrons will be in the higher energy t2 states with either element. Additionally, Δtet is approximately half the value of Δoct (all else being equal), meaning the energy penalty for electrons in the higher energy states is smaller for the tetragonal configuration.

The effect of cobalt in octahedral sites can be observed directly when comparing different simulations of CoAl2O4. Figure 4.4 shows that the relative structure energy is 2.83 eV for the

3Co,1Al case (which has 2 Co on octahedral sites and 1 Co on a tetrahedral site), meaning that the host structure has a full cell energy that is 2.83 eV higher (less favorable) than the 1Co,3Al case

(which has only 1 Co on a tetrahedral site) for the same material. The crystal field theory explanation above would suggest that this is because the cobalt in the octahedral sites would have d-electrons in higher energy states, and this is seen directly in Figure 4.12, which shows projected

DOS plots for these two cases. As can be seen, Figure 4.12b has additional d-orbital states which form the HOMO at the top of the valence band. The reason that the additional d-orbital states are described as being added on top of the valence band (as opposed to the conduction band being shifted downwards) is that the Fermi level of the two cases is also higher for the octahedral-Co

(3Co,1Al) case: 6.38 eV for the 1Co,3Al case and 7.32 eV for the 3Co,1Al case. This suggests that the 3Co,1Al case is shifting filled states to right above the previous conduction band. This reduces the band gap, which is shown in Figure 4.12.

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Figure 4.12: Projected density of states for CoAl2O4 with a) 1 and b) 2 Co as a nearest neighbor to the oxygen that will be removed. Energy values on the horizontal axis are relative to the Fermi energy for each system. Positive and negative vertical axis values are for spin-up and spin-down states, respectively. Density of states are scaled using different values for ease of viewing, and are not meant to be directly additive to obtain the total DOS.

Low inversion for CoAl2O4 was noted by Muhich et al. and used as an explanation for why

CoAl2O4 would produce such small amounts of hydrogen relative to hercynite [15]. The current calculated values for oxygen vacancy formation and relative structure energy are consistent with

Muhich et al., and this works shows the fundamental reasons behind this behavior. Deml et al. found that A-site doping in perovskites tended to lower band gaps and oxygen vacancy formation energies [5], and this work also sees a smaller band gap material giving smaller oxygen vacancy formation energies. However, this work suggests that the prevalence and favorability of different types of oxygen vacancy sites also plays a significant role. Al-Shankiti et al. found that U-doping into ceria gave much higher H2 production [6]; Al-Salik et al. used XPS to show that increased oxidation states for U led to more lower oxidation states for Ce [7], which Sacaranto and Idriss confirmed computationally with DFT+U [8]. This mechanism by which the dopant is destabilized in order to stabilize the active ion does not appear to be consistent with the current behavior in aluminate spinel materials. Instead, tetragonal Co is easier to reduce than Fe.

For future STWS materials development with iron aluminate spinels, this work suggests that small amounts of Co can be highly beneficial. It also suggests that doping other elements into

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the A-site of AAl2O4 compounds could be very beneficial. Lower atomic number elements (e.g.

Mn) can also take on 2+ and 3+ oxidation states but not have the same energy penalties to taking octahedral sites. Elements with different radii might also have beneficial effects, similar to ceria as described in Section 4.2. Many elements can take these oxidation states, and thus have the possibility to dope onto the A-site. This works shows the importance of taking the structure into account in addition to the oxygen vacancy formation energy.

4.5.4. Trade-Off Between Kinetics and Capacity

The cobalt-doped hercynite material has much lower capacity than un-doped hercynite, but reacts much more quickly. This has been reported previously [15], and is also observed in the current samples (Figure 4.13). This indicates that the cobalt does seem to have some kinetic effect on the material, in addition to changing the H2 production capacity. This can be partially explained by the results in Figure 4.11, in which the highest energy oxygen vacancies form near multiple Co atoms. This means that at equilibrium, relatively fewer of these vacancies will form on reduction, limiting capacity. However, during oxidation, these vacancies will have a higher exothermic reaction energy, which ensures that they are relatively more likely to re-oxidize. This can also contribute to a more rapid reaction rate through the Bell-Evans-Polanyi principle, in which changes to the reaction enthalpy can linearly affect the activation energy [29]. This suggests that this is a fundamental barrier in STWS materials development: lower reduction enthalpy values will lead to less capacity but faster reactions. Additional work is needed to further explore this kinetic effect, but the results here indicate that cobalt in this material does appear to be removing production capacity. Thus, this dopant should be added in the smallest possible amount to achieve the kinetic benefit, so as to remove as little of the capacity as possible. Additionally, other kinetic effects of

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dopants should ideally offset this kinetic loss, e.g. though catalyzing surface reactions or oxygen vacancy bulk transfer.

Figure 4.13: Hydrogen production rates for un-doped and Co-doped hercynite for an example oxidation reaction. For both materials, the fifth cycle oxidation is shown. 4.6. Conclusion

Iron aluminate spinel materials are examined for two-step solar thermochemical water splitting. Computational simulations have been performed using DFT+U for iron and cobalt aluminate spinel materials. These calculations predict that electron density transfer on the creation of oxygen vacancies is almost exclusively to the atoms that are first-nearest-neighbor to the oxygen vacancy. Calculations with different Co/Fe second nearest neighbors do not show different results.

This suggests that dopants must be prevalent enough to be next to any oxygen vacancies that are likely to form to have an effect. This also suggests that dopants that are near to the main transition metal (Fe) but not to the oxygen vacancy will likely not have an effect.

Computational results show that cases that have a single cobalt next to the oxygen vacancies tend to be very favorable both in terms of a lower oxygen vacancy formation energy and

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also a lower host structure energy, meaning those sites will likely to be more common in a real material. Projected density of states show that this is due to cobalt forms additional electron filled states in the valence band, suggested that the dopant changes oxidation state by taking the charge density from the oxygen vacancy. However, sites with more than one Co next to the O-vacancy tend to be less favorable. This is shown to be due to cobalt being less favorable on the octahedral sites due to d-electron orbital splitting as suggested by crystal field theory.

Current XPS results seem to contradict the computational results, as the Co-doped hercynite sample does not show Co oxidation state changes while the un-doped hercynite shows

Fe oxidation state changes. This is attributable to the excess cobalt in the Co-doped sample; more

Co make oxygen vacancies less likely to form. Hercynite samples with less cobalt should show more oxidation state change for the Co-doped sample.

Cobalt-doped hercynite has been showed to have faster reaction rates but less hydrogen production capacity compared to the un-doped hercynite. This might be due to kinetic/catalytic impacts of Co on the reaction mechanism, but can also be at least partially affected by the thermodynamics of the reaction enthalpy change as well. Reactions with similar mechanisms have shown a lower activation energy when the enthalpy of the reaction is increased. This suggests that the trade-off between capacity and kinetics for doped materials is likely a fundamental effect, unless dopants can introduce other kinetic effects to offset the effect.

This work suggests how new dopants can avoid problems for future development of materials. Co-doped hercynite can be improved by lowering the amount of cobalt and ensuring it is well-mixed in the material. Other dopants that can also assume the 2+/3+ oxidation states can potentially have less trouble in assuming the octahedral site and be very promising in improving the hercynite cycle.

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4.7. References

1. Muhich, C.L., et al., A review and perspective of efficient hydrogen generation via solar thermal water splitting. Wiley Interdisciplinary Reviews: Energy and Environment, 2016. 5(3): p. 261-287.

2. Ramos-Fernandez, E.V., N.R. Shiju, and G. Rothenberg, Understanding the solar-driven reduction of CO2 on doped ceria. RSC Advances, 2014. 4(32): p. 16456-16463.

4+ 4+ 3. Jiang, Q., et al., Thermochemical CO2 splitting reaction with CexM1−xO2−δ (M = Ti , Sn , Hf4+, Zr4+, La3+, Y3+ and Sm3+) solid solutions. Solar Energy, 2014. 99: p. 55-66.

4. Yang, C.-K., et al., Thermodynamic and kinetic assessments of strontium-doped lanthanum manganite perovskites for two-step thermochemical water splitting. Journal of Materials Chemistry A, 2014. 2(33): p. 13612-13623.

5. Deml, A.M., et al., Oxide enthalpy of formation and band gap energy as accurate descriptors of oxygen vacancy formation energetics. Energy & Environmental Science, 2014.

6. Al-Shankiti, I., et al., Solar Thermal Hydrogen Production from Water over Modified CeO2 Materials. Topics in Catalysis, 2013. 56(12): p. 1129-1138.

7. Al-Salik, Y., I. Al-Shankiti, and H. Idriss, Core level spectroscopy of oxidized and reduced CexU1−xO2 materials. Journal of Electron Spectroscopy and Related Phenomena, 2014. 194: p. 66-73.

8. Scaranto, J. and H. Idriss, The Effect of Uranium Cations on the Redox Properties of CeO2 Within the Context of Hydrogen Production from Water. Topics in Catalysis, 2014: p. 1-6.

9. Le Gal, A., S. Abanades, and G. Flamant, CO2 and H2O Splitting for Thermochemical Production of Solar Fuels Using Nonstoichiometric Ceria and Ceria/Zirconia Solid Solutions. Energy & Fuels, 2011. 25(10): p. 4836-4845.

10. Scheffe, J.R., et al., Synthesis, Characterization, and Thermochemical Redox Performance of 4+ 4+ 3+ Hf , Zr , and Sc Doped Ceria for Splitting CO2. The Journal of Physical Chemistry C, 2013. 117(46): p. 24104-24114.

11. Le Gal, A. and S. Abanades, Catalytic investigation of ceria-zirconia solid solutions for solar hydrogen production. International Journal of Hydrogen Energy, 2011. 36(8): p. 4739-4748.

12. Scheffe, J.R., D. Weibel, and A. Steinfeld, Lanthanum–Strontium–Manganese Perovskites as Redox Materials for Solar Thermochemical Splitting of H2O and CO2. Energy & Fuels, 2013. 27(8): p. 4250-4257.

13. McDaniel, A.H., et al., Sr- and Mn-doped LaAlO3-δ for solar thermochemical H2 and CO production. Energy & Environmental Science, 2013. 6: p. 2424-2428.

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14. Scheffe, J.R., J.H. Li, and A.W. Weimer, A spinel ferrite/hercynite water-splitting redox cycle. International Journal of Hydrogen Energy, 2010. 35(8): p. 3333-3340.

15. Muhich, C.L., et al., Predicting the solar thermochemical water splitting ability and reaction mechanism of metal oxides: a case study of the hercynite family of water splitting cycles. Energy & Environmental Science, 2015. 8(12): p. 3687-3699.

16. Kresse, G. and J. Furthmüller, Efficient iterative schemes for ab initio total-energy calculations using a plane-wave basis set. Physical Review B, 1996. 54(16): p. 11169-11186.

17. Kresse, G. and J. Furthmüller, Efficiency of ab-initio total energy calculations for metals and semiconductors using a plane-wave basis set. Computational Materials Science, 1996. 6(1): p. 15-50.

18. Perdew, J.P., K. Burke, and M. Ernzerhof, Generalized Gradient Approximation Made Simple. Physical Review Letters, 1996. 77(18): p. 3865-3868.

19. Kresse, G. and D. Joubert, From ultrasoft pseudopotentials to the projector augmented-wave method. Physical Review B, 1999. 59(3): p. 1758-1775.

20. Hubbard, J., Electron Correlations in Narrow Energy Bands. Proceedings of the Royal Society of London. Series A. Mathematical and Physical Sciences, 1963. 276(1365): p. 238- 257.

21. Chen, J., X. Wu, and A. Selloni, Electronic structure and bonding properties of cobalt oxide in the spinel structure. Physical Review B, 2011. 83(24): p. 245204.

22. Stevanović, V., et al., Correcting density functional theory for accurate predictions of compound enthalpies of formation: Fitted elemental-phase reference energies. Physical Review B, 2012. 85(11): p. 115104.

23. Henkelman, G., A. Arnaldsson, and H. Jónsson, A fast and robust algorithm for Bader decomposition of charge density. Computational Materials Science, 2006. 36(3): p. 354-360.

24. Sanville, E., et al., Improved grid-based algorithm for Bader charge allocation. Journal of Computational Chemistry, 2007. 28(5): p. 899-908.

25. Hansteen, O.H., H. Fjellvåg, and B.C. Hauback, Reduction, Crystal Structure and Magnetic Properties of Co3-xAlxO4-δ (0.0 ≤ x ≤ 2.0, 0.0 ≤ δ ≤ 1.0). Comparison with the Co/γ-Al2O3 Fischer-Tropsch Catalyst. Acta Chemica Scandinavica, 1998. 52: p. 1285-1292.

26. Ewbank, J.L., et al., Effect of preparation methods on the performance of Co/Al2O3 catalysts for dry reforming of methane. Green Chemistry, 2014. 16(2): p. 885-896.

27. Kodama, T. and N. Gokon, Thermochernical cycles for high-temperature solar hydrogen production. Chemical Reviews, 2007. 107(10): p. 4048-4077.

28. Housecroft, C.E. and A.G. Sharpe, Inorganic Chemistry. 2008: Pearson Prentice Hall.

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29. Dill, K. and S. Bromberg, Molecular Driving Forces: Statistical Thermodynamics in Biology, Chemistry, Physics, and Nanoscience. 2 ed. 2011: Garland Science, Taylor & Francis Group.

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CHAPTER V

CONCLUSIONS AND RECOMMENDATIONS FOR FUTURE WORK

5.1. Conclusions

Overall, the potential to produce H2 efficiently using only water and sunlight remains a major opportunity and challenge. This work contributes to an improved understanding of active materials and system efficiency. Overall conclusions will be summarized here, followed by suggestions for further improvements and critical efforts in this area.

5.1.1. Solar Thermochemical Water Splitting Materials and Reactors

Based on the current understanding of the many possible STWS cycles, two-step cycles provide the most promise in achieving the solar to hydrogen efficiencies which are required for economical hydrogen production. While stoichiometric cycles have high hydrogen production capacities, the propensity to form gas or liquid phases upon reduction limits the practical implementation of these materials. Doped-hercynite materials show potential due to their lower reduction temperatures and thus ability to operate isothermally, and higher H2 production capacity, but are currently limited by slow reaction rates. Among the O-vacancy-type cycles studied, ceria has been identified as the current standard STWS material for comparison, but still requires very high reduction temperatures despite years of research and doping with many various elements.

Perovskites reduce at lower temperatures and produce significant quantities of H2 at relatively high rates, and are largely unexplored for their use in STWS. As the field moves forward, research will likely concentrate on doped-hercynite and perovskite materials as they have the most benefits and fewest disadvantages of the STWS cycles.

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Proposed STWS reactors have ranged from simple and reliable cavity reactors with stationary reactive materials to complex reactors involving moving mechanical parts using the

STWS materials formed into both particles and monoliths. Based on lessons learned from various reactor configurations, a reactor based on flowing particles which requires a minimum number of moving parts is preferable.

5.1.2. System Efficiency Modeling

A ferrite/zirconia composite demonstrated higher thermodynamic efficiency values than ceria, primarily due to the increased hydrogen production capacity of the material. However, system efficiency based on thermodynamics alone may not be able to be realized. Kinetic limitations from oxidation reaction kinetics for ceria are minimal except at very high ΔT values.

The kinetic limitations for ferrite/zirconia are larger, causing large decreases in efficiency at even moderate ΔT values. Conditions that appear to be optimal for a particular material when considering only thermodynamic end-points can yield a very different set of operating conditions that may be different when kinetics are considered.

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Generally, the optimal system efficiency increases with both gas and solid heat recuperation. The optimal ΔT to operate the process increases with increasing solid heat recuperation but decreases with increasing gas heat recuperation. The impact of gas heat recuperation on the optimal STH efficiency is somewhat muted based on the current assumptions of a minimum inert gas flow rate and a low oxygen partial pressure. However, a more realistic system (with either or both of these assumptions relaxed) will show a much larger impact of both types of heat recuperation. Conversely, the current impact of solid heat recuperation on the optimal

STH efficiency is somewhat overstated, especially since this optimal efficiency occurs at unrealistic conditions. A higher inert gas flowrate will shift the balance of importance more towards gas heat recuperation rather than solids heat recuperation.

A high separation temperature for the recycled inert gas loop gives a higher system efficiency, especially gas heat recuperation is low and the inert gas flowrate is high. The increase in separation work that accompanies this is offset by greatly reduced sensible energy requirements to reach the very high reduction temperature. Furthermore, the high separation temperature means that the un-recuperated sensible heat from the inert gas stream is able to be used elsewhere in the system. However, a high temperature separation is only beneficial for a recycled inert gas system, not open loop continuous production.

Increasing the water/hydrogen separation temperature was examined as a way to decrease the importance of high temperature gas heat recuperation. An increase in this separation temperature has the largest impact on efficiency for smaller ΔT values and moderate values of gas heat recuperation, though the increase to the maximum ηSTH and shift of ΔTopt to smaller values are both fairly small. This relatively small increase with the current configuration is unlikely to be worth the increased system complexity and parasitic losses.

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Lower vacuum pump efficiencies suggest that currently available vacuum pumping results in low STH efficiencies. Further, the efficiency of a cascade pressure reduction has been examined, which uses multiple vacuum pumps operating at different pressures to achieve high levels of vacuum. It was found that cascade pressure reduction does increase the effective pump efficiency relative to a single vacuum pump, but this increase (~5X) is not large enough to cause a significant change in the system efficiency. However, a novel recycled inert gas sweep with a high temperature separation step maintains relatively high STH efficiencies over a range of gas purification efficiencies and flow rates. Thermodynamic separation and pumping efficiency values of around 10% are likely to result in high system efficiency values for both vacuum and inert gas technologies.

5.1.3. Engineered Particles for High Temperature Solar Thermochemical Cycles

A trade-off between spray dried particle robustness during redox cycling and particle activity has been shown for manganese oxide particles. Material formulations with added alumina and zirconia retained individual particle structure after being heated to 1,200°C, but did not perform as well during redox cycling. On the other hand, pure manganese oxide showed good redox activity but was unable to reliably form spray dried particles without the help of a colloidal suspension, and thus calcining the material produced mostly sintered agglomerates. Manganese oxide to iron oxide in a molar ratio of 1:2 was able to form spherical spray dried particles which agglomerated and lost spherical shape during calcining, but also exhibited the best activity of any material formulation tested. This thermochemical performance was obtained over multiple cycles despite this apparent loss of surface area.

Zirconia acts as a stable solid support material at the temperatures tested, and did not form a solid solution with manganese oxide. This suggests the inclusion of zirconia as an inert support

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for higher activity materials. Alumina and iron oxide both formed mixed metal oxides with a spinel structure after calcining. The addition of alumina and iron oxide did not improve the redox activity of the materials. The difference in performance of the two iron-containing samples is attributed to the more rapid attainment of thermodynamic equilibrium by the lower-iron containing material.

This has been attributed to the difference in ionic diffusion rates of iron and manganese.

A manganese iron oxide spinel with a 1:2 Mn:Fe molar ratio outperformed pure manganese oxide by a factor of 1.5, but does not retain spherical particle shape and undergoes sintering at high temperature. However, this high performance was shown over multiple cycles that were done after agglomeration and sintering. Thus, future research should investigate the combination of this active material with robust, inert particle supports to find a balance between activity and physical stability while still maintaining higher activity than pure manganese oxide.

Spray dried hercynite particles have been successfully produced using a simple mixture and pH modification method. This type of sol preparation has been demonstrated in the past, but typically the suspension had much higher sedimentation. This shows that partial flocculation, especially using a common charge stabilized pseudo-boehmite sol, can be an effective method to produce solid spherical particles. The incorporation of iron oxide has not been used with this material previously, and indicates that the flocculation behavior of the boehmite sol seems to drive performance. This suggests that this system could be used to produce different mixed oxides based on an aluminate base.

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simple synthesis technique of pH adjustment is able to be implemented successfully on a mixed oxide system illustrates a path forward for future synthesis at scale.

5.1.4. Role of Metal Ions in Hercynite Thermochemical Cycles

Iron aluminate spinel materials have been examined for two-step solar thermochemical water splitting. Computational simulations have been performed using DFT+U for iron and cobalt aluminate spinels. These simulations predict that charge transfer from oxygen vacancies is almost exclusively transferred to the first-nearest-neighbors to the oxygen vacancy, while different Co/Fe second-nearest-neighbors do not show different results. Thus, dopants should be prevalent enough to be next to any oxygen vacancies that are likely to form to have an effect.

Computational results also show that a single cobalt next to the oxygen vacancy tends to be favorable for a lower oxygen vacancy formation energy and also a lower host structure energy, meaning those sites will likely to be more common in a real material. Projected density of states show that this is due to cobalt forms additional electron filled states in the valence band, suggested that the dopant changes oxidation state by taking the charge density from the oxygen vacancy.

However, sites with more than one Co next to the O-vacancy tend to be less favorable. This was shown to be due to cobalt being less favorable on the octahedral sites from d-electron orbital splitting. Current XPS results seem to contradict the above computational results, as the Co-doped hercynite sample does not show significant Co oxidation state changes while the un-doped hercynite shows Fe oxidation state changes. This is attributable to the excess cobalt in the Co- doped sample; more Co make oxygen vacancies less likely to form. Hercynite samples with less cobalt should show more oxidation state change for the Co-doped sample.

Un-doped hercynite has been showed to have slower reaction rates but more H2 production capacity than Co-doped hercynite. This may be at least partially affected by the thermodynamics

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of the reaction enthalpy change. Reactions with similar mechanisms have shown a lower activation energy when the enthalpy of the reaction is increased. This suggests that the trade-off between capacity and kinetics for doped materials is likely a fundamental effect, unless dopants can introduce other kinetic effects to offset the effect. However, kinetic/catalytic impacts of Co on the reaction mechanism almost certainly contribute to the kinetics as well.

5.2. Recommendations for Future Work

Much progress has been made in solar thermochemical water splitting, both in materials advances and reactor concept design. Different reactor concepts need to be experimentally validated. Solar simulators can be very useful in this respect, as they are able to provide consistent and reproducible test conditions while allowing a reactor to be tested using semi-realistic thermal radiation. This would allow for direct and fair comparisons between some of the trade-offs identified in different reactor concepts. For example, monolithic reactors with no moving parts need to operate in a batch system, and experimental testing could help validate how quickly the reaction cycle can actually occur. Similarly, reactors with particles or moving parts need to be experimentally shown to be able to operate consistently. Direct and in-direct heating could have direct comparisons, which would be very useful for future designs. However, while more difficult, true solar testing is also important. Solar simulators typically do not have very fast response times, and one of the great challenges of high temperature solar thermal operation is the rapid changes in flux. While large swings in flux would be minimized in a well-controlled system, reactor concepts need to be able to handle these rapid changes without catastrophic failure.

5.2.1. System-Level Considerations

When considering operating conditions of various units within a solar thermochemical system, changes in any of the many parameters can shift ideal operating conditions drastically,

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meaning that a flexible reactor design is important. A reactor that allows for different operating temperatures, pressures, and flow rates between the reduction and oxidation reactions would allow for the system to always operate at the most efficient conditions, instead of being tied to a particular method or mode of operation. Additionally, realistic systems will require high gas heat recuperation. Materials research should be directed towards developing high temperature refractory steam/steam heat exchangers which can operate above 1,273 K, providing high levels of gas heat recuperation.

Achieving a low oxygen partial pressure in the reduction reaction enables highly efficiency operation. Future research should focus on the development of high temperature ionic transport membranes for efficiently separating O2 from a recycled inert gas sweep stream without having been substantially cooled. Alternately, additional efforts to achieve a thermodynamic vacuum pump efficiency of 10% would drastically increase the possibility of that approach to thermochemical reduction.

5.2.2. Improvements in Active Particles

Spray drying is a commercially ubiquitous method for producing many different kinds of particles, and has been used here to provide pathways to large scale production of STWS particles.

However, this work has only explored a tiny part of the potential for this technology. More secondary oxides should be explored for all different types of materials. Some materials have been shown to be nominally inert (such as zirconia when added to manganese oxide) but different formulations should be explored more in depth to see if the structural benefits of the inert phase can offset the reduced reactivity. This testing needs to be done with moving particles to validate the physical robustness as well, such as a fluidized bed.

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Additionally, more works needs to be done in determining best practices for making mixed metal oxides via spray drying. The high temperatures required for phase formation and sintering for mechanical strength can lead to inter-particle agglomeration as well, especially when the particles are stationary. The use of a rotating kiln and the determination of necessary temperatures and time to reach equilibrium are important for current and new formulations that may be developed.

5.2.3. Improved Doping in Aluminate Spinel Materials

This works has highlighted the importance of tying theoretical and experimental work for a more complete understanding of materials and reactions. The combination of computational simulations with various types of analytical techniques make for a much more complete picture than either one of these aspects individually. Experimentation can provide evidence for theoretical predictions, and theoretical simulations can help to explain analytical observations. Matching experimental and theoretical results can be challenging, but when this is accomplished it provides a much more convincing argument for the way that systems, materials, and atoms behave.

This work has also highlighted the usefulness of in-situ analytical techniques. These techniques should be expanded and improved so that other types of measurements can be done at the reaction conditions of interest. X-ray transparent materials allow for heating to high temperature while still performing X-ray-dependent techniques (e.g., XRD and XPS). Other techniques such as TGA allow for direct observation of mass changes while at high temperatures.

It is not immediately obvious how in-situ methods could be directly applied to currently used analytical techniques. High temperature capillaries could allow for local gas sampling as close as possible to the material of interest. Electron gun heating could be used in conjunction with

SEM/TEM techniques for microscopy and elemental analysis. While difficult, these techniques

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can give direct observation of important aspects to reactions that occur at challenging conditions, and should be used wherever possible.

This work also suggests possible areas for future improvements to spinel-type materials.

Co-doped hercynite can be improved by lowering the amount of cobalt and ensuring it is well- mixed in the material. This should allow for Co to provide easier and more rapid reduction and oxidation while not sacrificing nearly as much capacity. Other dopants that can also assume the

2+/3+ oxidation states can potentially have less trouble in assuming the octahedral site and be very promising in improving the hercynite cycle. Preferably these elements would not have the same un-favorability in residing on the octahedral site in spinels, so that increased levels of spinel inversion would be possible. Furthermore, dopants with different atomic radii could be used in order to superimpose tensile/compressive strain fields onto the material to further fine-tine the oxygen vacancy formation energy.

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BIBLIOGRAPHY

Hydrogen Analysis Resource Center: Lower and Higher Heating Values of Hydrogen and Fuels. Available from: http://hydrogen.pnl.gov/hydrogen-data/lower-and-higher-heating-values- hydrogen-and-other-fuels.

The Impact of Increased Use of Hydrogen on Petroleum Consumption and Carbon Dioxide Emissions. 2008, U.S. Energy Information Administration (EIA): Report No. SR/OIAF- CNEAF/2008-04.

SunShot Concentrating Solar Power (CSP) R&D Funding Opportunity Announcement, Department of Energy, DE-FOA-0000595 CFDA Number 81.087. 2011: Golden, CO.

Abanades, S., CO2 and H2O reduction by solar thermochemical looping using SnO2/SnO redox reactions: Thermogravimetric analysis. International Journal of Hydrogen Energy, 2012. 37(10): p. 8223-8231.

Abanades, S., P. Charvin, and G. Flamant, Design and simulation of a solar chemical reactor for the thermal reduction of metal oxides: Case study of zinc oxide dissociation. Chemical Engineering Science, 2007. 62(22): p. 6323-6333.

Abanades, S., et al., Screening of water-splitting thermochemical cycles potentially attractive for hydrogen production by concentrated solar energy. Energy, 2006. 31(14): p. 2805-2822.

Abanades, S., et al., Novel two-step SnO2/SnO water-splitting cycle for solar thermochemical production of hydrogen. International Journal of Hydrogen Energy, 2008. 33(21): p. 6021-6030.

Abanades, S. and G. Flamant, Thermochemical hydrogen production from a two-step solar- driven water-splitting cycle based on cerium oxides. Solar Energy, 2006. 80(12): p. 1611- 1623.

Abanades, S., et al., Investigation of reactive cerium-based oxides for H2 production by thermochemical two-step water-splitting. Journal of Materials Science, 2010. 45(15): p. 4163-4173.

Abu Hamed, T., J.H. Davidson, and M. Stolzenburg, Hydrolysis of Evaporated Zn in a Hot Wall Flow Reactor. Journal of Solar Energy Engineering, 2008. 130(4): p. 041010-041010.

Adanez, J., et al., Progress in Chemical-Looping Combustion and Reforming technologies. Progress in Energy and Combustion Science, 2012. 38(2): p. 215-282.

Agrafiotis, C., M. Roeb, and C. Sattler, A review on solar thermal syngas production via redox pair-based water/carbon dioxide splitting thermochemical cycles. Renewable and Sustainable Energy Reviews, 2015. 42: p. 254-285.

198

Albano, M.a.P. and L.B. Garrido, Rheological properties of Si3N4, pseudoboehmite and bayerite coated Si3N4 suspensions with ammonium polyacrylate dispersant. Colloids and Surfaces A: Physicochemical and Engineering Aspects, 2002. 203(1–3): p. 117-128.

Alioshin, Y., et al., High Temperature Electrolysis of CO2 for Fuel Production. Journal of The Electrochemical Society, 2016. 163(2): p. F79-F87.

Allen, K.M., et al., Study of a Magnetically Stabilized Porous Structure for Thermochemical Water Splitting via TGA, High-Temperature-XRD, and SEM Analyses. Industrial & Engineering Chemistry Research, 2013. 52(10): p. 3683-3692.

Allendorf, M.D., et al., Two-Step Water Splitting Using Mixed-Metal Ferrites: Thermodynamic Analysis and Characterization of Synthesized Materials. Energy & Fuels, 2008. 22(6): p. 4115-4124.

Alonso, E., et al., Kinetics of Mn2O3–Mn3O4 and Mn3O4–MnO Redox Reactions Performed under Concentrated Thermal Radiative Flux. Energy & Fuels, 2013. 27(8): p. 4884-4890.

Al-Salik, Y., I. Al-Shankiti, and H. Idriss, Core level spectroscopy of oxidized and reduced CexU1−xO2 materials. Journal of Electron Spectroscopy and Related Phenomena, 2014. 194: p. 66-73.

Al-Shankiti, I., et al., Solar Thermal Hydrogen Production from Water over Modified CeO2 Materials. Topics in Catalysis, 2013. 56(12): p. 1129-1138.

Arifin, D., Study of redox reactions to split water and carbon dioxide, in Chemical and Biological Engineering. 2013, University of Colorado: Boulder.

Arjmand, M., et al., Screening of Combined Mn-Fe-Si Oxygen Carriers for Chemical Looping with Oxygen Uncoupling (CLOU). Energy & Fuels, 2015. 29(3): p. 1868-1880.

Azimi, G., et al., Chemical-looping with oxygen uncoupling using combined Mn-Fe oxides, testing in batch fluidized bed. Energy Procedia, 2011. 4: p. 370-377.

Azimi, G., et al., Comprehensive study of Mn–Fe–Al oxygen-carriers for chemical-looping with oxygen uncoupling (CLOU). International Journal of Greenhouse Gas Control, 2015. 34: p. 12-24.

Babiniec, S.M., et al., Investigation of LaxSr1−xCoyM1−yO3−δ (M = Mn, Fe) perovskite materials as thermochemical energy storage media. Solar Energy, 2015. 118: p. 451-459.

Babiniec, S.M., et al., Doped calcium manganites for advanced high-temperature thermochemical energy storage. International Journal of Energy Research, 2016. 40(2): p. 280-284.

Bader, R., et al., Thermodynamic Analysis of Isothermal Redox Cycling of Ceria for Solar Fuel Production. Energy & Fuels, 2013. 27(9): p. 5533-5544.

199

Baek, J.-I., et al., Performance Comparison of Spray-Dried Oxygen Carriers: The Effect of NiO and Pseudoboehmite Content in Raw Materials. Energy & Fuels, 2010. 24(10): p. 5757- 5764.

Bala Chandran, R., R.M. De Smith, and J.H. Davidson, Model of an integrated solar thermochemical reactor/reticulated ceramic foam heat exchanger for gas-phase heat recovery. International Journal of Heat and Mass Transfer, 2015. 81: p. 404-414.

Bale, C.W., et al., FactSage thermochemical software and databases — recent developments. Calphad, 2009. 33(2): p. 295-311.

Banerjee, A., R. Bala Chandran, and J.H. Davidson, Experimental investigation of a reticulated porous alumina heat exchanger for high temperature gas heat recovery. Applied Thermal Engineering, 2015. 75: p. 889-895.

Barreto, L., A. Makihira, and K. Riahi, The hydrogen economy in the 21st century: a sustainable development scenario. International Journal of Hydrogen Energy, 2003. 28(3): p. 267- 284.

Berman, A. and M. Epstein, The kinetics of hydrogen production in the oxidation of liquid zinc with water vapor. International Journal of Hydrogen Energy, 2000. 25(10): p. 957-967.

Bertrand, G., et al., Influence of slurry characteristics on the morphology of spray-dried alumina powders. Journal of the European Ceramic Society, 2003. 23(2): p. 263-271.

Bertrand, G., et al., Spray-dried ceramic powders: A quantitative correlation between slurry characteristics and shapes of the granules. Chemical Engineering Science, 2005. 60(1): p. 95-102.

Bhosale, R., et al., H2 generation from two-step thermochemical water-splitting reaction using sol-gel derived SnxFeyOz. Journal of Renewable and Sustainable Energy, 2011. 3(6): p. -.

Bhosale, R.R., et al., Sol–gel derived CeO2–Fe2O3 nanoparticles: Synthesis, characterization and solar thermochemical application. Ceramics International, 2016. 42(6): p. 6728- 6737.

Brendelberger, S., et al., Counter flow sweep gas demand for the ceria redox cycle. Solar Energy, 2015. 122: p. 1011-1022.

Brendelberger, S. and C. Sattler, Concept analysis of an indirect particle-based redox process for solar-driven H2O/CO2 splitting. Solar Energy, 2015. 113: p. 158-170.

Bulfin, B., et al., Thermodynamics of CeO2 Thermochemical Fuel Production. Energy & Fuels, 2015. 29(2): p. 1001-1009.

Bulfin, B., et al., Analytical Model of CeO2 Oxidation and Reduction. The Journal of Physical Chemistry C, 2013. 117(46): p. 24129-24137.

200

Caiut, J.M.A., et al., Elaboration of boehmite nano-powders by spray-pyrolysis. Powder Technology, 2009. 190(1–2): p. 95-98.

Cao, X.Q., et al., Spray-drying of ceramics for plasma-spray coating. Journal of the European Ceramic Society, 2000. 20(14–15): p. 2433-2439.

Carrillo, A.J., et al., Thermochemical energy storage at high temperature via redox cycles of Mn and Co oxides: Pure oxides versus mixed ones. Solar Energy Materials and Solar Cells, 2014. 123: p. 47-57.

Carrillo, A.J., et al., Improving the Thermochemical Energy Storage Performance of the Mn2O3/Mn3O4 Redox Couple by the Incorporation of Iron. ChemSusChem, 2015. 8(11): p. 1947-1954.

Chambon, M., S. Abanades, and G. Flamant, Kinetic investigation of hydrogen generation from hydrolysis of SnO and Zn solar nanopowders. International Journal of Hydrogen Energy, 2009. 34(13): p. 5326-5336.

Chambon, M., S. Abanades, and G. Flamant, Solar thermal reduction of ZnO and SnO2: Characterization of the recombination reaction with O2. Chemical Engineering Science, 2010. 65(11): p. 3671-3680.

Chambon, M., S. Abanades, and G. Flamant, Thermal dissociation of compressed ZnO and SnO2 powders in a moving-front solar thermochemical reactor. AIChE Journal, 2011. 57(8): p. 2264-2273.

Charvin, P., et al., Two-step water splitting thermochemical cycle based on iron oxide redox pair for solar hydrogen production. Energy, 2007. 32(7): p. 1124-1133.

Charvin, P., et al., Experimental study of SnO2/SnO/Sn thermochemical systems for solar production of hydrogen. AIChE Journal, 2008. 54(10): p. 2759-2767.

Chen, J., X. Wu, and A. Selloni, Electronic structure and bonding properties of cobalt oxide in the spinel structure. Physical Review B, 2011. 83(24): p. 245204.

Chueh, W.C. and S.M. Haile, A thermochemical study of ceria: exploiting an old material for new modes of energy conversion and CO2 mitigation. Philosophical transactions. Series A, Mathematical, physical, and engineering sciences, 2010. 368(1923): p. 3269-94.

Coker, E.N., et al., Oxygen transport and isotopic exchange in iron oxide/YSZ thermochemically- active materials via splitting of C(18O)2 at high temperature studied by thermogravimetric analysis and secondary ion mass spectrometry. Journal of Materials Chemistry, 2012. 22(14): p. 6726-6732.

Daniell, W., et al., Enhanced surface acidity in mixed alumina–silicas: a low-temperature FTIR study. Applied Catalysis A: General, 2000. 196(2): p. 247-260.

Dasari, H.P., et al., Hydrogen production from water-splitting reaction based on RE-doped

201

ceria–zirconia solid-solutions. International Journal of Hydrogen Energy, 2013. 38(14): p. 6097-6103.

Davis, R.H. and K.H. Birdsell, Hindered settling of semidilute monodisperse and polydisperse suspensions. AIChE Journal, 1988. 34(1): p. 123-129.

Deml, A.M., et al., Oxide enthalpy of formation and band gap energy as accurate descriptors of oxygen vacancy formation energetics. Energy & Environmental Science, 2014.

Demont, A., S. Abanades, and E. Beche, Investigation of Perovskite Structures as Oxygen- Exchange Redox Materials for Hydrogen Production from Thermochemical Two-Step Water-Splitting Cycles. The Journal of Physical Chemistry C, 2014. 118(24): p. 12682- 12692.

Dill, K. and S. Bromberg, Molecular Driving Forces: Statistical Thermodynamics in Biology, Chemistry, Physics, and Nanoscience. 2 ed. 2011: Garland Science, Taylor & Francis Group.

Diver, R.B., et al., Solar Thermochemical Water-Splitting Ferrite-Cycle Heat Engines. Journal of Solar Energy Engineering, 2008. 130(4): p. 041001-041001-8.

Domalski, E.S. and E.D. Hearing, Condensed Phase Heat Capacity Data, in NIST Chemistry WebBook, Standard Reference Database Number 69, P.J. Linstrom and W.G. Mallard, Editors, National Institute of Standards and Technology: Gaithersburg, MD.

Dyer, P.N., et al., Ion transport membrane technology for oxygen separation and syngas production. Solid State Ionics, 2000. 134(1–2): p. 21-33.

Ehrensberger, K., et al., Comparative experimental investigations of the water-splitting reaction with iron oxide Fe1−yO and iron manganese oxides (Fe1−xMnx)1−yO. Solid State Ionics, 1995. 78(1–2): p. 151-160.

Ehrhart, B., et al., Thermochemical Cycle of a Mixed Metal Oxide for Augmentation of Thermal Energy Storage in Solid Particles. Energy Procedia, 2014. 49: p. 762-771.

Ehrhart, B.D. and D.D. Gill, Evaluation of Annual Efficiencies of High Temperature Central Receiver Concentrated Solar Power Plants With Thermal Energy Storage. 2013, Sandia National Laboratories: Report No. SAND2013-5493 Albuquerque, NM.

Ehrhart, B.D., et al., System efficiency for two-step metal oxide solar thermochemical hydrogen production – Part 1: Thermodynamic model and impact of oxidation kinetics. International Journal of Hydrogen Energy, 2016. 41(44): p. 19881-19893.

Ehrhart, B.D., et al., System efficiency for two-step metal oxide solar thermochemical hydrogen production – Part 2: Impact of gas heat recuperation and separation temperatures. International Journal of Hydrogen Energy, 2016. 41(44): p. 19894-19903.

Ehrhart, B.D., et al., System efficiency for two-step metal oxide solar thermochemical hydrogen

202

production – Part 3: Various methods for achieving low oxygen partial pressures in the reduction reaction. International Journal of Hydrogen Energy, 2016. 41(44): p. 19904- 19914.

Eriksson, G. and K. Hack, ChemSage—A computer program for the calculation of complex chemical equilibria. Metallurgical Transactions B, 1990. 21(6): p. 1013-1023.

Ermanoski, I., Cascading pressure thermal reduction for efficient solar fuel production. International Journal of Hydrogen Energy, 2014. 39(25): p. 13114-13117.

Ermanoski, I., J.E. Miller, and M.D. Allendorf, Efficiency maximization in solar-thermochemical fuel production: challenging the concept of isothermal water splitting. Physical Chemistry Chemical Physics, 2014. 16(18): p. 8418-8427.

Ermanoski, I., N.P. Siegel, and E.B. Stechel, A New Reactor Concept for Efficient Solar- Thermochemical Fuel Production. Journal of Solar Energy Engineering, 2013. 135(3): p. 031002-031002-10.

Eslamian, M. and N. Ashgriz, Effect of precursor, ambient pressure, and temperature on the morphology, crystallinity, and decomposition of powders prepared by spray pyrolysis and drying. Powder Technology, 2006. 167(3): p. 149-159.

Evdou, A., V. Zaspalis, and L. Nalbandian, La1−xSrxFeO3−δ perovskites as redox materials for application in a membrane reactor for simultaneous production of pure hydrogen and synthesis gas. Fuel, 2010. 89(6): p. 1265-1273.

Ewbank, J.L., et al., Effect of preparation methods on the performance of Co/Al2O3 catalysts for dry reforming of methane. Green Chemistry, 2014. 16(2): p. 885-896.

Falcone, P.K., Handbook for solar central receiver design. 1986, Sandia National Laboratories: Report No. SAND86-8009 Livermore, CA.

Fan, L.-S., Chemical Looping systems for Fossil Energy Conversions. 2010, Hoboken, NJ: American Institute of Chemical Engineers (AIChE) and John Wiley & Sons, Inc.

Felinks, J., et al., Heat recovery concept for thermochemical processes using a solid heat transfer medium. Applied Thermal Engineering, 2014. 73(1): p. 1004-1011.

Felinks, J., et al., Particle–particle heat transfer coefficient in a binary packed bed of alumina and zirconia-ceria particles. Applied Thermal Engineering, 2016. 101: p. 101-111.

Francis, T.M., P.R. Lichty, and A.W. Weimer, Manganese oxide dissociation kinetics for the Mn2O3 thermochemical water-splitting cycle. Part 1: Experimental. Chemical Engineering Science, 2010. 65(12): p. 3709-3717.

Furler, P., et al., Thermochemical CO2 splitting via redox cycling of ceria reticulated foam structures with dual-scale porosities. Physical Chemistry Chemical Physics, 2014. 16(22): p. 10503-10511.

203

Garnier, J., A thermodynamic study on CaO-doped nonstoichiometric cerium dioxide. Journal of Physics and Chemistry of Solids, 1976. 37(4): p. 368-378.

Gilewicz-Wolter, J., et al., Diffusion of chromium, manganese, and iron in MnCr2O4 spinel. Journal of Phase Equilibria and Diffusion, 2005. 26(5): p. 561-564.

Gokon, N., et al., New Solar Water-Splitting Reactor With Ferrite Particles in an Internally Circulating Fluidized Bed. Journal of Solar Energy Engineering, 2009. 131(1): p. 011007-011007.

Hablanian, M.H., High-Vacuum Technology: A Practical Guide, Second Edition. 1997: Taylor & Francis.

Hallberg, P., et al., Chemical Looping Combustion and Chemical Looping with Oxygen Uncoupling Experiments in a Batch Reactor Using Spray-Dried CaMn1–xMxO3−δ (M = Ti, Fe, Mg) Particles as Oxygen Carriers. Energy & Fuels, 2013. 27(3): p. 1473-1481.

Hansteen, O.H., H. Fjellvåg, and B.C. Hauback, Reduction, Crystal Structure and Magnetic Properties of Co3-xAlxO4-δ (0.0 ≤ x ≤ 2.0, 0.0 ≤ δ ≤ 1.0). Comparison with the Co/γ-Al2O3 Fischer-Tropsch Catalyst. Acta Chemica Scandinavica, 1998. 52: p. 1285-1292.

Hao, Y., C.-K. Yang, and S.M. Haile, High-temperature isothermal chemical cycling for solar- driven fuel production. Physical Chemistry Chemical Physics, 2013. 15(40): p. 17084- 17092.

Heat Transfer International. High Temperature Ceramic Heat Exchangers. [cited 2015 July 14]; Available from: http://www.heatxfer.com/products/htche.html.

Henkelman, G., A. Arnaldsson, and H. Jónsson, A fast and robust algorithm for Bader decomposition of charge density. Computational Materials Science, 2006. 36(3): p. 354- 360.

Hildebrandt, A.F. and K.A. Rose, Receiver design considerations for solar central receiver hydrogen production. Solar Energy, 1985. 35(2): p. 199-206.

Holladay, J.D., et al., An overview of hydrogen production technologies. Catalysis Today, 2009. 139(4): p. 244-260.

Housecroft, C.E. and A.G. Sharpe, Inorganic Chemistry. 2008: Pearson Prentice Hall.

Hubbard, J., Electron Correlations in Narrow Energy Bands. Proceedings of the Royal Society of London. Series A. Mathematical and Physical Sciences, 1963. 276(1365): p. 238-257.

Imhof, A., The cyclone reactor — an atmospheric open solar reactor. Solar Energy Materials, 1991. 24(1–4): p. 733-741.

Iskandar, F., L. Gradon, and K. Okuyama, Control of the morphology of nanostructured particles prepared by the spray drying of a nanoparticle sol. Journal of Colloid and Interface

204

Science, 2003. 265(2): p. 296-303.

Jarrett, C., et al., Critical limitations on the efficiency of two-step thermochemical cycles. Solar Energy, 2016. 123: p. 57-73.

4+ 4+ 4+ Jiang, Q., et al., Thermochemical CO2 splitting reaction with CexM1−xO2−δ (M = Ti , Sn , Hf , Zr4+, La3+, Y3+ and Sm3+) solid solutions. Solar Energy, 2014. 99: p. 55-66.

Jr. Walker, W.J., J.S. Reed, and S.K. Verma, Influence of Slurry Parameters on the Characteristics of Spray-Dried Granules. Journal of the American Ceramic Society, 1999. 82(7): p. 1711-1719.

Kaneko, H., et al., Cerium ion redox system in CeO2–xFe2O3 solid solution at high temperatures (1,273–1,673 K) in the two-step water-splitting reaction for solar H2 generation. Journal of Materials Science, 2008. 43(9): p. 3153-3161.

Kaneko, H., et al., Decomposition of Zn-ferrite for O2 generation by concentrated solar radiation. Solar Energy, 2004. 76(1–3): p. 317-322.

Kaneko, H., et al., Reactive ceramics of CeO2–MOx (M=Mn, Fe, Ni, Cu) for H2 generation by two-step water splitting using concentrated solar thermal energy. Energy, 2007. 32(5): p. 656-663.

Kanerva, U., et al., Evaluation of crushing strength of spray-dried MgAl2O4 granule beds. Ceramics International, 2015. 41(7): p. 8494-8500.

Kanezashi, M. and M. Asaeda, Hydrogen permeation characteristics and stability of Ni-doped silica membranes in steam at high temperature. Journal of Membrane Science, 2006. 271(1–2): p. 86-93.

Kang, M., et al., Enhanced thermochemical CO2 splitting over Mg- and Ca-doped ceria/zirconia solid solutions. RSC Advances, 2014. 4(11): p. 5583-5590.

Kang, M., et al., CO2 splitting via two step thermochemical reactions over doped ceria/zirconia solid solutions. RSC Advances, 2013. 3(41): p. 18878-18885.

Kasahara, S., et al., Flowsheet study of the thermochemical water-splitting iodine-sulfur process for effective hydrogen production. International Journal of Hydrogen Energy, 2007. 32(4): p. 489-496.

Keene, D.J., J.H. Davidson, and W. Lipiński, A Model of Transient Heat and Mass Transfer in a Heterogeneous Medium of Ceria Undergoing Nonstoichiometric Reduction. Journal of Heat Transfer, 2013. 135(5): p. 052701-052701-9.

Kim, J., et al., Fuel production from CO2 using solar-thermal energy: system level analysis. Energy & Environmental Science, 2012. 5(9): p. 8417-8429.

Kim, J., et al., Comparative analysis of environmental impact of S2P (Sunshine to Petrol) system

205

for transportation fuel production. Applied Energy, 2013. 111(0): p. 1089-1098.

Kodama, T. and N. Gokon, Thermochernical cycles for high-temperature solar hydrogen production. Chemical Reviews, 2007. 107(10): p. 4048-4077.

Kodama, T., et al., Thermochemical hydrogen production by a redox system of ZrO2-supported Co(II)-ferrite. Solar Energy, 2005. 78(5): p. 623-631.

Koepf, E., et al., A novel beam-down, gravity-fed, solar thermochemical receiver/reactor for direct solid particle decomposition: Design, modeling, and experimentation. International Journal of Hydrogen Energy, 2012. 37(22): p. 16871-16887.

Koepf, E.E., et al., Experimental Investigation of Vortex Flow in a Two-Chamber Solar Thermochemical Reactor. Journal of Fluids Engineering, 2013. 135(11): p. 111103- 111103.

Kreider, P.B., et al., Manganese oxide based thermochemical hydrogen production cycle. International Journal of Hydrogen Energy, 2011. 36(12): p. 7028-7037.

Krenzke, P.T. and J.H. Davidson, Thermodynamic Analysis of Syngas Production via the Solar Thermochemical Cerium Oxide Redox Cycle with Methane-Driven Reduction. Energy & Fuels, 2014. 28(6): p. 4088-4095.

Krenzke, P.T. and J.H. Davidson, On the Efficiency of Solar H2 and CO Production via the Thermochemical Cerium Oxide Redox Cycle: The Option of Inert-Swept Reduction. Energy & Fuels, 2015. 29(2): p. 1045-1054.

Kresse, G. and J. Furthmüller, Efficient iterative schemes for ab initio total-energy calculations using a plane-wave basis set. Physical Review B, 1996. 54(16): p. 11169-11186.

Kresse, G. and J. Furthmüller, Efficiency of ab-initio total energy calculations for metals and semiconductors using a plane-wave basis set. Computational Materials Science, 1996. 6(1): p. 15-50.

Kresse, G. and D. Joubert, From ultrasoft pseudopotentials to the projector augmented-wave method. Physical Review B, 1999. 59(3): p. 1758-1775.

Lange, M., et al., T–S diagram efficiency analysis of two-step thermochemical cycles for solar water splitting under various process conditions. Energy, 2014. 67: p. 298-308.

Lange, M., et al., Efficiency assessment of a two-step thermochemical water-splitting process based on a dynamic process model. International Journal of Hydrogen Energy, 2015. 40(36): p. 12108-12119.

Lange, M., et al., Entropy Analysis of Solar Two-Step Thermochemical Cycles for Water and Carbon Dioxide Splitting. Entropy, 2016. 18(1): p. 24.

Lapp, J., J.H. Davidson, and W. Lipiński, Efficiency of two-step solar thermochemical non-

206

stoichiometric redox cycles with heat recovery. Energy, 2012. 37(1): p. 591-600.

Le Gal, A. and S. Abanades, Catalytic investigation of ceria-zirconia solid solutions for solar hydrogen production. International Journal of Hydrogen Energy, 2011. 36(8): p. 4739- 4748.

Le Gal, A. and S. Abanades, Dopant Incorporation in Ceria for Enhanced Water-Splitting Activity during Solar Thermochemical Hydrogen Generation. The Journal of Physical Chemistry C, 2012. 116(25): p. 13516-13523.

Le Gal, A., et al., Reactivity of Doped Ceria-Based Mixed Oxides for Solar Thermochemical Hydrogen Generation via Two-Step Water-Splitting Cycles. Energy & Fuels, 2013. 27(10): p. 6068-6078.

Le Gal, A., S. Abanades, and G. Flamant, CO2 and H2O Splitting for Thermochemical Production of Solar Fuels Using Nonstoichiometric Ceria and Ceria/Zirconia Solid Solutions. Energy & Fuels, 2011. 25(10): p. 4836-4845.

Levêque, G. and S. Abanades, Kinetic analysis of high-temperature solid–gas reactions by an inverse method applied to ZnO and SnO2 solar thermal dissociation. Chemical Engineering Journal, 2013. 217(0): p. 139-149.

Levêque, G. and S. Abanades, Investigation of thermal and carbothermal reduction of volatile oxides (ZnO, SnO2, GeO2, and MgO) via solar-driven vacuum thermogravimetry for thermochemical production of solar fuels. Thermochimica Acta, 2015. 605: p. 86-94.

Levêque, G., et al., Characterization of Two-Step Tin-Based Redox System for Thermochemical Fuel Production from Solar-Driven CO2 and H2O Splitting Cycle. Industrial & Engineering Chemistry Research, 2014. 53(14): p. 5668-5677.

Lewis, N.S. and D.G. Nocera, Powering the planet: Chemical challenges in solar energy utilization. Proceedings of the National Academy of Sciences, 2006. 103(43): p. 15729- 15735.

Lichty, P., et al., Atomic layer deposited thin film metal oxides for fuel production in a solar cavity reactor. International Journal of Hydrogen Energy, 2012. 37(22): p. 16888-16894.

Lichty, P., et al., Rapid High Temperature Solar Thermal Biomass Gasification in a Prototype Cavity Reactor. Journal of Solar Energy Engineering, 2010. 132(1): p. 011012-011012.

Lin, M. and S. Haussener, Solar fuel processing efficiency for ceria redox cycling using alternative oxygen partial pressure reduction methods. Energy, 2015. 88: p. 667-679.

Linderholm, C., T. Mattisson, and A. Lyngfelt, Long-term integrity testing of spray-dried particles in a 10-kW chemical-looping combustor using natural gas as fuel. Fuel, 2009. 88(11): p. 2083-2096.

Liu, B., et al., Fluidization of non-spherical particles: Sphericity, Zingg factor and other

207

fluidization parameters. Particuology, 2008. 6(2): p. 125-129.

Liu, H., et al., A novel FCC catalyst synthesized via in situ overgrowth of NaY zeolite on kaolin microspheres for maximizing propylene yield. Catalysis Today, 2007. 125(3–4): p. 163- 168.

Loutzenhiser, P.G., A. Meier, and A. Steinfeld, Review of the two-step H2O/CO2-splitting solar thermochemical cycle based on Zn/ZnO redox reactions. Materials, 2010. 3(11): p. 4922- 4938.

Lukasiewicz, S.J., Spray-Drying Ceramic Powders. Journal of the American Ceramic Society, 1989. 72(4): p. 617-624.

Lundberg, M., Model calculations on some feasible two-step water splitting processes. International Journal of Hydrogen Energy, 1993. 18(5): p. 369-376.

Mahdjoub, H., et al., The effect of the slurry formulation upon the morphology of spray-dried yttria stabilised zirconia particles. Journal of the European Ceramic Society, 2003. 23(10): p. 1637-1648.

Mallapragada, D.S. and R. Agrawal, Limiting and achievable efficiencies for solar thermal hydrogen production. International Journal of Hydrogen Energy, 2014. 39(1): p. 62-75.

Marbán, G. and T. Valdés-Solís, Towards the hydrogen economy? International Journal of Hydrogen Energy, 2007. 32(12): p. 1625-1637.

Martinek, J., R. Viger, and A.W. Weimer, Transient simulation of a tubular packed bed solar receiver for hydrogen generation via metal oxide thermochemical cycles. Solar Energy, 2014. 105(0): p. 613-631.

Martinek, J. and A.W. Weimer, Design considerations for a multiple tube solar reactor. Solar Energy, 2013. 90: p. 68-83.

Martinek, J. and A.W. Weimer, Evaluation of finite volume solutions for radiative heat transfer in a closed cavity solar receiver for high temperature solar thermal processes. International Journal of Heat and Mass Transfer, 2013. 58(1–2): p. 585-596.

Marugán, J., et al., Study of the first step of the Mn2O3/MnO thermochemical cycle for solar hydrogen production. International Journal of Hydrogen Energy, 2012. 37(8): p. 7017- 7025.

Marugán, J., et al., Study of the hydrogen production step of the Mn2O3/MnO thermochemical cycle. International Journal of Hydrogen Energy, 2014. 39(10): p. 5274-5282.

McDaniel, A.H., et al., Nonstoichiometric Perovskite Oxides for Solar Thermochemical H2 and CO Production. Energy Procedia, 2014. 49(0): p. 2009-2018.

McDaniel, A.H., et al., Sr- and Mn-doped LaAlO3-δ for solar thermochemical H2 and CO

208

production. Energy & Environmental Science, 2013. 6: p. 2424-2428.

Meixner, D.L., et al., Electrochemical Oxygen Separation Using Solid Electrolyte Ion Transport Membranes. Journal of the Electrochemical Society, 2002. 149(9): p. D132-D136.

Melchior, T., et al., A cavity-receiver containing a tubular absorber for high-temperature thermochemical processing using concentrated solar energy. International Journal of Thermal Sciences, 2008. 47(11): p. 1496-1503.

Melchior, T., N. Piatkowski, and A. Steinfeld, H2 production by steam-quenching of Zn vapor in a hot-wall aerosol flow reactor. Chemical Engineering Science, 2009. 64(5): p. 1095- 1101.

Meng, Q.-L., et al., Reactivity of CeO2-based ceramics for solar hydrogen production via a two- step water-splitting cycle with concentrated solar energy. International Journal of Hydrogen Energy, 2011. 36(21): p. 13435-13441.

Meng, Q.-L., et al., Solar hydrogen production using Ce1−xLixO2−δ solid solutions via a thermochemical, two-step water-splitting cycle. Journal of Solid State Chemistry, 2012. 194(0): p. 343-351.

Meredig, B. and C. Wolverton, First-principles thermodynamic framework for the evaluation of thermochemical H2O or CO2 splitting materials. Physical Review B, 2009. 80(24): p. 245119.

Miller, J.E., et al., Sunshine to Petrol: Solar thermochemistry for liquid fuels. Abstracts of Papers of the American Chemical Society, 2011. 241.

Miller, J.E., A.H. McDaniel, and M.D. Allendorf, Considerations in the Design of Materials for Solar-Driven Fuel Production Using Metal-Oxide Thermochemical Cycles. Advanced Energy Materials, 2014. 4(2): p. 1300469.

Muhich, C.L., et al., A review and perspective of efficient hydrogen generation via solar thermal water splitting. Wiley Interdisciplinary Reviews: Energy and Environment, 2016. 5(3): p. 261-287.

Muhich, C.L., et al., Predicting the solar thermochemical water splitting ability and reaction mechanism of metal oxides: a case study of the hercynite family of water splitting cycles. Energy & Environmental Science, 2015. 8(12): p. 3687-3699.

Muhich, C.L., et al., Efficient Generation of H2 by Splitting Water with an Isothermal Redox Cycle. Science, 2013. 341(6145): p. 540-542.

Müller, R., P. Haeberling, and R.D. Palumbo, Further advances toward the development of a direct heating solar thermal chemical reactor for the thermal dissociation of ZnO(s). Solar Energy, 2006. 80(5): p. 500-511.

Naglieri, V., et al., Optimized Slurries for Spray Drying: Different Approaches to Obtain

209

Homogeneous and Deformable Alumina-Zirconia Granules. Materials, 2013. 6(11): p. 5382.

Nakamura, N., et al., Hydrogen production via thermochemical water-splitting by lithium redox reaction. Journal of Alloys and Compounds, 2013. 580, Supplement 1(0): p. S410-S413.

Nakamura, T., Hydrogen production from water utilizing solar heat at high temperatures. Solar Energy, 1977. 19(5): p. 467-475.

Nalbandian, L., A. Evdou, and V. Zaspalis, La1−xSrxMO3 (M = Mn, Fe) perovskites as materials for thermochemical hydrogen production in conventional and membrane reactors. International Journal of Hydrogen Energy, 2009. 34(17): p. 7162-7172.

Nandiyanto, A.B.D. and K. Okuyama, Progress in developing spray-drying methods for the production of controlled morphology particles: From the nanometer to submicrometer size ranges. Advanced Powder Technology, 2011. 22(1): p. 1-19.

Naterer, G., et al., Recent Canadian advances in nuclear-based hydrogen production and the thermochemical Cu-Cl cycle. International Journal of Hydrogen Energy, 2009. 34(7): p. 2901-2917.

Naterer, G.F., et al., Thermochemical hydrogen production with a copper-chlorine cycle. I: oxygen release from copper oxychloride decomposition. International Journal of Hydrogen Energy, 2008. 33(20): p. 5439-5450.

Ogden, J.M., Prospects for Building a Hydrogen Energy Infrastructure. Annual Review of Energy and the Environment, 1999. 24(1): p. 227-279.

Oles, A.S. and G.S. Jackson, Modeling of a concentrated-solar, falling-particle receiver for ceria reduction. Solar Energy, 2015. 122: p. 126-147.

Panlener, R.J., R.N. Blumenthal, and J.E. Garnier, A Thermodynamic Study of Nonstoichiometric Cerium Dioxide. Journal of Physics and Chemistry of Solids, 1975. 36: p. 1213-1222.

Pardo, P., et al., A review on high temperature thermochemical heat energy storage. Renewable and Sustainable Energy Reviews, 2014. 32: p. 591-610.

Perdew, J.P., K. Burke, and M. Ernzerhof, Generalized Gradient Approximation Made Simple. Physical Review Letters, 1996. 77(18): p. 3865-3868.

Perkins, C., P. Lichty, and A.W. Weimer, Determination of aerosol kinetics of thermal ZnO dissociation by thermogravimetry. Chemical Engineering Science, 2007. 62(21): p. 5952- 5962.

Perkins, C., P.R. Lichty, and A.W. Weimer, Thermal ZnO dissociation in a rapid aerosol reactor as part of a solar hydrogen production cycle. International Journal of Hydrogen Energy, 2008. 33(2): p. 499-510.

210

Perkins, C. and A.W. Weimer, Likely near-term solar-thermal water splitting technologies. International Journal of Hydrogen Energy, 2004. 29(15): p. 1587-1599.

Perkins, C. and A.W. Weimer, Solar-thermal production of renewable hydrogen. AIChE Journal, 2009. 55(2): p. 286-293.

Ramos-Fernandez, E.V., N.R. Shiju, and G. Rothenberg, Understanding the solar-driven reduction of CO2 on doped ceria. RSC Advances, 2014. 4(32): p. 16456-16463.

Rao, C.N.R., et al., Solar photochemical and thermochemical splitting of water. Philosophical Transactions of the Royal Society of London A: Mathematical, Physical and Engineering Sciences, 2016. 374(2061).

Roeb, M., et al., Materials-Related Aspects of Thermochemical Water and Carbon Dioxide Splitting: A Review. Materials, 2012. 5(11): p. 2015-2054.

Roeb, M., et al., Operational strategy of a two-step thermochemical process for solar hydrogen production. International Journal of Hydrogen Energy, 2009. 34(10): p. 4537-4545.

Roy, P., G. Bertrand, and C. Coddet, Spray drying and sintering of zirconia based hollow powders. Powder Technology, 2005. 157(1–3): p. 20-26.

Sander, R., Henry's Law Constants, in NIST Chemistry WebBook, Standard Reference Database Number 69, P.J. Linstrom and W.G. Mallard, Editors, National Institute of Standards and Technology: Gaithersburg, MD.

Sanville, E., et al., Improved grid-based algorithm for Bader charge allocation. Journal of Computational Chemistry, 2007. 28(5): p. 899-908.

Scaranto, J. and H. Idriss, The Effect of Uranium Cations on the Redox Properties of CeO2 Within the Context of Hydrogen Production from Water. Topics in Catalysis, 2014: p. 1- 6.

Scheffe, J.R., et al., Synthesis, Characterization, and Thermochemical Redox Performance of 4+ 4+ 3+ Hf , Zr , and Sc Doped Ceria for Splitting CO2. The Journal of Physical Chemistry C, 2013. 117(46): p. 24104-24114.

Scheffe, J.R., et al., Kinetics and mechanism of solar-thermochemical H2 production by oxidation of a cobalt ferrite-zirconia composite. Energy & Environmental Science, 2013. 6(3): p. 963-973.

Scheffe, J.R. and A. Steinfeld, Oxygen exchange materials for solar thermochemical splitting of H2O and CO2: a review. Materials Today, (0).

Scheffe, J.R. and A. Steinfeld, Thermodynamic Analysis of Cerium-Based Oxides for Solar Thermochemical Fuel Production. Energy & Fuels, 2012. 26(3): p. 1928-1936.

Scheffe, J.R., D. Weibel, and A. Steinfeld, Lanthanum–Strontium–Manganese Perovskites as

211

Redox Materials for Solar Thermochemical Splitting of H2O and CO2. Energy & Fuels, 2013. 27(8): p. 4250-4257.

Schunk, L.O. and A. Steinfeld, Kinetics of the thermal dissociation of ZnO exposed to concentrated solar irradiation using a solar-driven thermogravimeter in the 1800–2100 K range. AIChE Journal, 2009. 55(6): p. 1497-1504.

Siegel, N.P., et al., Factors Affecting the Efficiency of Solar Driven Metal Oxide Thermochemical Cycles. Industrial & Engineering Chemistry Research, 2013. 52(9): p. 3276-3286.

Simpson, A.P. and A.E. Lutz, Exergy analysis of hydrogen production via steam methane reforming. International Journal of Hydrogen Energy, 2007. 32(18): p. 4811-4820.

Singh, A.K., et al., Thermal Reduction of Iron Oxide under Reduced Pressure and Implications on Thermal Conversion Efficiency for Solar Thermochemical Fuel Production. Industrial & Engineering Chemistry Research, 2015. 54(26): p. 6793-6803.

Smestad, G.P. and A. Steinfeld, Review: Photochemical and Thermochemical Production of Solar Fuels from H2O and CO2 Using Metal Oxide Catalysts. Industrial & Engineering Chemistry Research, 2012. 51(37): p. 11828-11840.

Smith, J.M., H.C. Van Ness, and M.M. Abbott, Introduction to chemical engineering thermodynamics. 7th ed. 2005: McGraw-Hill.

Srinivasan, R., et al., Identification of tetragonal and cubic structures of zirconia using synchrotron x-radiation source. Journal of Materials Research, 1991. 6(6): p. 1287-1292.

Steinfeld, A., Solar hydrogen production via a two-step water-splitting thermochemical cycle based on Zn/ZnO redox reactions. International Journal of Hydrogen Energy, 2002. 27(6): p. 611-619.

Steinfeld, A., S. Sanders, and R. Palumbo, Design Aspects of Solar Thermochemical Engineering—A Case Study: Two-Step Water-Splitting Cycle Using the Fe3O4/FeO Redox System. Solar Energy, 1999. 65(1): p. 43-53.

Stekli, J., L. Irwin, and R. Pitchumani, Technical Challenges and Opportunities for Concentrating Solar Power With Thermal Energy Storage. Journal of Thermal Science and Engineering Applications, 2013. 5(2): p. 021011-021011.

Stevanović, V., et al., Correcting density functional theory for accurate predictions of compound enthalpies of formation: Fitted elemental-phase reference energies. Physical Review B, 2012. 85(11): p. 115104.

Takacs, M., J.R. Scheffe, and A. Steinfeld, Oxygen nonstoichiometry and thermodynamic characterization of Zr doped ceria in the 1573-1773 K temperature range. Physical Chemistry Chemical Physics, 2015. 17(12): p. 7813-7822.

212

Tamaura, Y., et al., Production of solar hydrogen by a novel, 2-step, water-splitting thermochemical cycle. Energy, 1995. 20(4): p. 325-330.

Tregambi, C., et al., Heat transfer in directly irradiated fluidized beds. Solar Energy, 2016. 129: p. 85-100.

Tsetsekou, A., et al., Optimization of the rheological properties of alumina slurries for ceramic processing applications Part II: Spray-drying. Journal of the European Ceramic Society, 2001. 21(4): p. 493-506.

Turner, J., et al., Renewable hydrogen production. International Journal of Energy Research, 2008. 32(5): p. 379-407.

U.S. Energy Information Administration (EIA). Few transportation fuels surpass the energy densities of gasoline and diesel. Today in Energy 2013; Available from: http://www.eia.gov/todayinenergy/detail.cfm?id=9991.

U.S. Energy Information Administration (EIA). Hydrogen for refineries is increasingly provided by industrial suppliers. Today in Energy 2016; Available from: http://www.eia.gov/todayinenergy/detail.cfm?id=24612.

U.S. Energy Information Administration (EIA), Monthly Energy Review. April 2016: Report No.

U.S. Geological Survey (USGS), Mineral Commodity Summaries 2016. 2016.

Venstrom, L.J., et al., Applicability of an Equilibrium Model To Predict the Conversion of CO2 to CO via the Reduction and Oxidation of a Fixed Bed of Cerium Dioxide. Energy & Fuels, 2015. 29(12): p. 8168-8177.

Venstrom, L.J., et al., Efficient Splitting of CO2 in an Isothermal Redox Cycle Based on Ceria. Energy & Fuels, 2014. 28(4): p. 2732-2742.

Vishnevetsky, I. and M. Epstein, Tin as a Possible Candidate for Solar Thermochemical Redox Process for Hydrogen Production. Journal of Solar Energy Engineering, 2009. 131(2): p. 021007-021007.

Walton, D.E. and C.J. Mumford, Spray Dried Products—Characterization of Particle Morphology. Chemical Engineering Research and Design, 1999. 77(1): p. 21-38.

Wang, Z., et al., Comparison of thermochemical, electrolytic, photoelectrolytic and photochemical solar-to-hydrogen production technologies. International Journal of Hydrogen Energy, 2012. 37(21): p. 16287-16301.

Weidenkaff, A., et al., Thermogravimetric analysis of the ZnO/Zn water splitting cycle. Thermochimica Acta, 2000. 359(1): p. 69-75.

Winter, C.-J., Hydrogen energy — Abundant, efficient, clean: A debate over the energy-system- of-change. International Journal of Hydrogen Energy, 2009. 34(14, Supplement 1): p. S1-

213

S52.

Wong, B., et al., Oxide based thermochemical heat storage, in SolarPACES. 2010: Perpignan, France.

Wong, B. and Project Staff, Thermochemical Heat Storage for Concentrated Solar Power: Thermochemical System Reactor Design for Thermal Energy Storage. 2011, General Atomics: Report No. GA-C27137.

Wu, J. and X.f. Long, Research progress of solar thermochemical energy storage. International Journal of Energy Research, 2015. 39(7): p. 869-888.

Xu, B., Y. Bhawe, and M.E. Davis, Low-temperature, manganese oxide-based, thermochemical water splitting cycle. Proceedings of the National Academy of Sciences of the United States of America, 2012. 109(24): p. 9260-9264.

Xu, R. and T.F. Wiesner, Conceptual design of a two-step solar hydrogen thermochemical cycle with thermal storage in a reaction intermediate. International Journal of Hydrogen Energy, 2014. 39(24): p. 12457-12471.

Yang, C.-K., et al., Thermodynamic and kinetic assessments of strontium-doped lanthanum manganite perovskites for two-step thermochemical water splitting. Journal of Materials Chemistry A, 2014. 2(33): p. 13612-13623.

Yang, W., et al., Quantum Chemical Calculations on the Reaction of Zinc and Water in Gas Phase. Combustion Science and Technology, 2013. 186(1): p. 24-33.

Yildiz, B. and M.S. Kazimi, Efficiency of hydrogen production systems using alternative nuclear energy technologies. International Journal of Hydrogen Energy, 2006. 31(1): p. 77-92.

Yoshida, T. and K. Kojima, Toyota MIRAI Fuel Cell Vehicle and Progress Toward a Future Hydrogen Society. The Electrochemical Society Interface, 2015. 24(2): p. 45-49.

Yuan, C., et al., A new solar fuels reactor concept based on a liquid metal heat transfer fluid: Reactor design and efficiency estimation. Solar Energy, 2015. 122: p. 547-561.

Yun, S. and S. Ted Oyama, Correlations in palladium membranes for hydrogen separation: A review. Journal of Membrane Science, 2011. 375(1–2): p. 28-45.

Zou, Z., et al., Direct splitting of water under visible light irradiation with an oxide semiconductor photocatalyst. Nature, 2001. 414(6864): p. 625-627.

214