TRANSIENT PERFORMANCE OF STEAM REFORMERS IN THE CONTEXT OF AUTOMOTIVE FUEL CELL SYSTEM INTEGRATION

Transcription

1 TRANSIENT PERFORMANCE OF STEAM REFORMERS IN THE CONTEXT OF AUTOMOTIVE FUEL CELL SYSTEM INTEGRATION By DANIEL AUGUSTO BETTS A DISSERTATION PRESENTED TO THE GRADUATE SCHOOL OF THE UNIVERSITY OF FLORIDA IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF DOCTOR OF PHILOSOPHY UNIVERSITY OF FLORIDA 005

2 Copyright 005 by Daniel Augusto Betts

3 During my PhD studies a lot of happiness and tragedy occurred in my life. For this reason I would like to dedicate this work, which represents the culmination of all these things, to the following individuals: Erica E. Carr-Betts, my wife (thank you for always supporting me. I love you very much); Carmen A. Carrington Betts, my mother (thank you for being strong and fighting the good fight. You have immense strength and courage); Claude D. Betts, my father (thank you for your reason and courage); Lydia C. Carrington, my grandmother (I am so sorry that we could not save you. Thank you for teaching me that I can); and to Matilda Eva Carr-Betts, my daughter (thank you for bringing me so much joy).

4 ACKNOWLEDGMENTS I would like to acknowledge the following individuals who directly or indirectly contributed to this dissertation. First, my wife Erica Eva Carr-Betts was instrumental in the elaboration of this study. She funded this endeavor, and maintained my sanity during difficult times. Her support and love had no bounds. My advisors (Dr. William Lear and Dr. Vernon Roan) were very patient with me. Their instruction and encouragement fueled the hope that an end was near, even when it was not. I am eternally grateful to them. Those who worked at the Ford Fuel Cell Laboratory are also responsible for this work coming to fruition. Special thanks are given to Dr. Timothy Simons and Ryotaro Honjo, who were always available to give me a hand. Dr. Paul Erickson, an alumni of UF and an old friend, was gracious enough to allow me to conduct experiments in his reformers at UC-Davis. His collaboration made this work possible. The US Department of Education and the UF Mechanical and Aerospace Engineering Department supported my research with a fellowship. Without this fellowship I would not have been able to obtain this education. I am grateful. I promise to repay this favor by dedicating part of my efforts to providing educational and professional opportunities to others. iv

5 TABLE OF CONTENTS page ACKNOWLEDGMENTS... iv LIST OF TABLES... vii LIST OF FIGURES... viii ABSTRACT... xi CHAPTER 1 INTRODUCTION...1 BACKGROUND...7 Fuel Processing...7 Steam Reforming...8 Partial Oxidation Reforming...9 Autothermal Reforming...10 Transportation Fuel Cells...11 The Transportation Load Environment...1 CO and Hydrogen Concentration Effects on PEMFC performance...16 Methanol as Fuel LITERATURE REVIEW...3 Kinetics of the Catalytic Methanol Steam Reformer Reactor...4 Reformer Performance ANALYSIS...9 Steam Reformer Model...9 Heat Transfer...31 Continuity...35 Conservation of Specie...38 Chemical Model...39 Finite Difference Solution...41 FEMLAB Solution...44 Steam Reformer Model Validation...45 General Description...46 v

6 UC-Davis Reformer Data...49 Reformer Design...57 Design Generalization RESULTS...65 Transient Efficiency...65 Case 1 - Infinite Reformer...65 Case - The Medium Sized Reformer...67 Case 3 - The Short Reformer...68 Carbon Monoxide Concentration...70 Hydrogen Concentration CONCLUSIONS...78 Steam Reformer Modeling...78 Steady-state Reformer Operation...78 Unsteady Reformer Operation SUGGESTED FUTURE WORK...80 APPENDIX A B COPY OF FEMLAB SOLUTION REPORT...81 PRELIMINARY LOAD FORECASTING STUDY...99 C PHYSICAL DESCRIPTION OF THE UC-DAVIS REFORMER General Description Pumping Subassembly Catalyst Bed Housing Subassemblies Condensing Unit Subassembly LIST OF REFERENCES BIOGRAPHICAL SKETCH vi

7 LIST OF TABLES Table page 3-1. Test parameters for Nakagaki tests Estimated maximum times for vaporizer volume flow change Values constants used to develop virtual UC-Davis reformer Steady-state design of 60 kw reformer with an 11.9 mm radius and a catalyst bed density of 1983 kg/m Case 1: Summary of the transient response of the infinite reformer to changes in fuel flow rate Case : Summary of the transient response of the medium size reformer to changes in fuel flow rate Case 3: Summary of the transient efficiency response of the short reformer to changes in fuel flow rate...69 C-1. UC-Davis reformer subassemblies vii

8 LIST OF FIGURES Figure page 1-1. Types of fuel cells and fuel choice used for fuel cell vehicle prototypes Energy density of different fuels on a lower heating value (LHV) basis compared to hydrogen Typical polarization curves of various types of fuel cells Driving load of a 30 ft phosphoric acid fuel cell bus operated at the University of Florida. The bus weighs is approximately 30,000 lbs Indirect methanol automotive fuel cell system power delivery schematic TBB- fuel cell bus methanol steam reformer and steam reformer burner temperatures during transient bus operation University of Florida bus indirect methanol fuel cell system Effect of hydrogen concentration and utilization on PEMFC performance Effect of CO poisoning on PEMFC performance Use of air bleed to recover PEMFC performance General schematic of modeled steam reformer reactor Control volume diagram Convection-conduction ratio in the UC-Davis reformer fuel entry regions Calculated values of dt/dt for the UC-Davis reformer at the entrance region Measured temperature distributions of the UC-Davis steam reforming test rig operating at varying space velocities and varying conversion efficiencies Schematic representation chemical reaction model in the control volume Equilibrium products of the steam reforming reaction (CH 3 OH + H O ah O + bco + cco + dh ) with respect to temperature...41 viii

9 4-8. Separation of the reformer volume into annular elements with constant gas properties in space for the finite difference solution method used Experimental steam reformer developed at UC Davis Example of CO spike in the reformate of the UC-Davis reformer after an increase in premix fuel flow (transient condition). Also shown is a decrease of the steady-state CO concentration in the reformate with increased premix fuel flow rate Reformate CO composition relationship to changes in temperature in the reformer catalyst bed CO and H concentrations in the reformate gas in the UC-Davis reformer Estimation of catalyst bed thermal conductivity using UC-Davis reformer data Virtual UC-Davis reformer steady-state CO composition with varying premix fuel flow rates Hydrocarbon signal output obtained during UC-Davis reformer run Virtual UC-Davis Reformer CO spike from 10 to 15ml/min premix flow rate change Virtual UC-Davis reformer temperature distribution (color plot) and methanol concentration (contour plot) for premix flow rate equal to 5 ml/min Virtual UC-Davis reformer temperature distribution (color plot) and methanol concentration (contour plot) for premix flow rate equal to 30 ml/min Reformer Efficiency with varying catalyst effectiveness at varying fuel flow rates Case 1: Infinite reformer operation when undergoing step variations in fuel flow rate Case 1: Reformer efficiency changes due to change of fuel flow rate for the infinite reformer Case : Transient efficiency of the medium size reformer under transient load changes Case : Reformer efficiency changes due to change of fuel flow rate for the medium size reformer Case 3: Transient efficiency of the short reformer under transient load changes...70 ix

10 5-6. Case 3: Reformer efficiency changes due to change of fuel flow rate for the short reformer Infinite reformer transient CO concentrations for various for changes in power output Infinite reformer transient off-steady-state CO concentrations for various changes in reformer power output Medium reformer transient CO concentrations for various changes in power output Medium reformer off-steady-state CO concentrations for various changes in reformer power output Short reformer transient CO concentrations for various changes in power output Short reformer off-steady-state CO concentration for various changes in reformer power output Transient CO comparison of various size reformers operating between 40 and 50 kw hydrogen power output Comparison of transient changes in H concentration induced by step changes in fuel flow corresponding to hydrogen yields from 10 to 60 kw (lower heating value based) for various size reformers...77 B-1. Transient load profile for TBB- under the University of Florida Fuel Cell Lab driving cycle data obtained on B-. Power Change Index (PCI) for TBB- under UFFL driving cycle data obtained on B-3. Multi-layer Perceptron B-4. The MLP performance as a 10 second PCI predictor C-1. General process diagram of the UC-Davis reformer C-. Premix reservoir and gear pump C-3. Vaporizer design C-4. Superheater design C-5. The UC-Davis reactor design...11 x

11 Abstract of Dissertation Presented to the Graduate School of the University of Florida in Partial Fulfillment of the Requirements for the Degree of Doctor of Philosophy TRANSIENT PERFORMANCE OF STEAM REFORMERS IN THE CONTEXT OF AUTOMOTIVE FUEL CELL SYSTEM INTEGRATION By Daniel Augusto Betts December 005 Chair: William Lear Cochair: Vernon Roan Major Department: Mechanical and Aerospace Engineering Proton exchange membrane fuel cells (PEMFCs) and, to a lesser degree, phosphoric acid fuel cells (PAFCs) have been widely studied as possible replacements for transportation internal combustion engines (ICE). These fuel cells consume hydrogen as fuel, which is electrochemically oxidized through an acid electrolyte. Because of the low energy density and scarcity of elemental hydrogen, alternatives such as methanol and natural gas have been investigated as primary fuels for PEMFC and PAFC fuel cell systems. Of these fuels, methanol is the most easily reformed into a hydrogen-rich gas. The most efficient way of doing this is through catalytic steam reforming. Therefore a clear understanding of the performance of steam reformers may lead to better integration of these devices into fuel cell engines. In this dissertation the results of studies regarding the transient and steady-state performance of methanol steam reformers for automotive fuel cell system integration are provided. To power an automobile, a fuel cell system needs to be capable of changing xi

12 power output rapidly, to adapt to changing driving loads. Although most fuel cell systems currently use batteries to reduce the required response time associated with the fuel cells and other balance of plant components (including the reformer), the ultimate goal may be to reduce the reliance of the system on batteries (which add cost, weight, and complexity to the system). This means that the fuel cell stack and the reformer must each be capable of changing power outputs quickly and efficiently. Fuel cell stack efficiency is intimately related to reformate gas composition (especially CO concentration). Since the reformer is upstream of the fuel cell stack, it has great influence on the overall power output and efficiency response of the fuel cell system. The results of this study are based on data obtained from an experimental reformer and numerical models. Non-dimensionalization of the governing equations derived for reformer model development resulted in identification of potential reformer similarity variables. Based on these reformer sizing and scaling theories were developed. Of particular importance was the further demonstration of the insufficiency of space velocity and aspect ratio as sole reformer similarity variables. Step changes in fuel flow into the experimental reformer produced transient CO concentration spikes. These spikes have also been identified in reformer literature. Through the use of the transient reformer model, potential physical mechanisms that cause these CO concentration spikes were identified. Studies in conversion efficiency and transient and steady-state reformate hydrogen concentration were also carried out. xii

13 CHAPTER 1 INTRODUCTION Interest in fuel cell technology for transportation has led to an ever-increasing number of prototype fuel cell vehicle demonstrations in recent years. For example, 10 new prototype fuel cell vehicles were built in 004. This is more than the cumulative number of fuel cell prototypes developed between 1959 and 00 [1]. Enthusiasm for fuel cells centers on their high energy-conversion efficiencies, the cleanliness of its primary fuel (hydrogen), and the cleanliness of its operational by-products (water, electricity and heat). Yet commercialization of fuel cell vehicles faces major hurdles including design of low-cost and robust systems, proper water-management design for transient operation at various power levels and conditions, improvement of stack life, improvement of hydrogen storage technology, and improvement of hydrogen-production technology. Generally, hydrogen supplied to the fuel cell stack comes from one of two sources: on-board hydrogen storage, or a fuel processor. Currently, most fuel cell vehicle manufacturers are developing vehicles that use hydrogen as their primary fuel (Figure 1-1). On-board hydrogen storage is intended to simplify fuel cell engines and reduce costs. However, even with pure hydrogen, fuel cell systems are far from achieving operational goals, especially when the power output of the fuel cell system is cycled through time. In addition, hydrogen storage technology is far from achieving the energy density goals needed for widespread vehicle commercialization. Figure 1- compares the energy densities of common hydrogen-rich fuels. Even with the most advanced hydrogen 1

14 energy-storage mechanisms, the energy density of hydrogen is far less than that of other fuels. Currently, hydrogen-fueled fuel cell vehicles have achieved a maximum driving range of approximately 300 km [1]. Major research efforts are now aiming to improve hydrogen storage capacity. These efforts include using carbon nanotubes as a hydrogen storage medium, and developing safe hydrogen tanks capable of operating at pressures of over 700 bars. Figure 1-1. Types of fuel cells and fuel choice used for fuel cell vehicle prototypes. This figure was published with the permission of its original authors [1]. Introduction of reformation technology into fuel cell vehicles has been proposed as a possible inexpensive way to solve the energy density problems associated with hydrogen []. Industry analysts have pointed out that alternative liquid fuels could make use of parts of the current global fuel delivery and storage infrastructure, potentially reducing the cost of the fuels. In addition, alternative liquid- fuels could still continue to be used in internal combustion engine vehicles without major modifications. Of the fuels typically considered, methanol seems the most attractive because it is easily reformed at relatively low temperatures and can be produced from a wide range of feedstock. In terms of energy density, methanol has approximately half the energy density of gasoline, and methanol is approximately 1.45 times more energy dense than hydrogen stored in advanced hydride beds (Figure 1-). It is important to point out that the process of steam

15 3 and autothermal reformation includes the dissociation of water to produce hydrogen, therefore yielding greater hydrogen per carbon concentration than the simple methanol dissociation reaction. Energy Density (MJ/L) Cetane (diesel) Octane (gasoline) Carbon Hydrogen heptane hexane pentane butane ethane propane ethanol methane methanol ammonia liq. Hydrogen Hydride Water Figure 1-. Energy density of different fuels on a lower heating value (LHV) basis compared to hydrogen [3]. Using methanol as primary fuel for PEMFC vehicles is technically challenging. Methanol-derived reformate contains hydrogen diluted mostly in carbon dioxide, water, carbon monoxide, and methanol. Among the problems is that the dilution of hydrogen produces a drop in cell efficiency. This drop in cell performance can be measured as a drop in cell voltage (from V 1 to V ) that corresponds to a drop in cell hydrogen partial pressure from (P 1 to P ) as shown in Equation 1-1.

16 4 P V = V1 C ln (1-1) P 1 The value of the constant C has been measured to range from 0.03 to 0.06 for PEMFCs. This drop in voltage leads to a drop in fuel cell efficiency and operational power density range. A positive aspect of reformation is that water vapor is among the diluents in the anode fuel stream. This reduces the need for anode gas humidification, which in hydrogen-fueled fuel cell systems can be very complicated. The reformate gas also contains carbon monoxide, which has a poisoning effect on low-temperature fuel cells 1. The PEMFCs have been poisoned with CO concentrations from 30 ppm. The equilibrium concentration of CO for the steam reforming reaction of methanol at 5 o C and 300 o C (the typical range of temperature for catalytic steam reforming) are 0.8 and.3 mole percent, which is beyond the limits of most current PEMFC technology tolerance. Various reformate gas clean-up schemes aim to reduce the concentration of CO in the anode fuel stream. Notable among these are preferential oxidizers and specialized membranes (generally palladium) capable of filtering CO and other gases from the reformate stream. Coupled with these technical issues is the need for controlling the fuel cell and the reformer when changes in load are required, in such a way that system efficiency drops are minimized and response speed is maximized. Figure 1-3 shows the recorded efficiencies of one of the 30ft Georgetown fuel cell buses (TBB-) during a typical drive 1 CO poisoning in fuel cells is a reversible phenomenon. Small concentrations of oxygen can be used to reverse the CO poisoning effects. Other techniques include increasing cell current to promote the electrochemical water-gas shift reaction.

17 5 cycle. 1 These buses use methanol as their fuel and use a steam reformer to produce hydrogen that is consumed by a phosphoric acid fuel cell (PAFC). The transient changes in system efficiency shown are caused by variations in anode hydrogen stoichiometric flow, changes in reformate hydrogen gas concentration, and changes in reformate carbon monoxide concentration. Temperature variations in the fuel cell stack may also contribute to these efficiency changes. The goal of this study is to provide a broader understanding of the role of the reformer in the dynamic performance of automotive load-following fuel cell systems, and provide a basis for reformer design and modeling. Analyses in this study are derived from a mixture of physical modeling results and experimental observations from which general trends in reformer steady-state and transient conversion, hydrogen concentration, and carbon monoxide concentration were derived. 1 The TBB- bus is a 30 ft phosphoric acid fuel cell bus housed at the University of Florida FORD fuel cell laboratory. This bus uses on-board methanol steam reformation for hydrogen production.

18 Max Power = 50kW Rate of Chemical Energy in the Premix flow was calculated using the lower heating value of MeOH in liquid form Run Time (min) FC System Efficiency Ratio (FC Power/Max Power) Stack Efficiency Figure 1-3. Measured transient efficiency of the 30 ft Georgetown Fuel Cell Bus (TBB-)

19 CHAPTER BACKGROUND In this chapter, several basic concepts essential to the general understanding of the analysis and results of this dissertation are provided. The chapter describes key fuel processing techniques and compares these techniques to each other. In addition, a short description of low temperature automotive fuel cells is provided. Also, general systems integration issues of indirect methanol fuel cell systems 1 are presented. In this section, data obtained from the Georgetown fuel cell test bed buses is used to show transient interactions between the fuel processor and the fuel cell in an actual system. Finally, a general discussion regarding methanol as a potential automotive fuel cell fuel is given. Fuel Processing Because of the difficulties of direct electrochemical oxidation of hydrocarbons, hydrocarbon fuels are normally processed (or reformed) to give a hydrogen-rich fuel mixture. Reforming plays an especially important role in low-temperature fuel cells, namely proton exchange membrane fuel cells (PEMFCs) and phosphoric acid fuel cells (PAFCs), since they are unable to oxidize any other fuel but hydrogen (H ). There are three basic reforming methods: steam reforming, partial oxidation reforming, and 1 Indirect methanol fuel cell systems refers to the use of methanol as the system fuel. The methanol in this systems use a reformer to produce hydrogen, which ultimately is the fuel cell fuel. The term indirect is used to differentiate these systems from direct methanol fuel cell systems, which are capable of oxidizing methanol in the fuel cell without the use of a reformer. Direct methanol fuel cells (DMFCs) commonly use similar electrolyte material, namely Nafion, as PEMFCs. In this text the term PEMFC refers to hydrogen proton exchange membrane fuel cells and excludes DMFCs. 7

20 8 autothermal reforming. This study concentrates on steam reforming since, of the aforementioned methods, it is the most efficient form of hydrogen production from hydrocarbon fuels. Steam Reforming Steam reforming is the reacting of a hydrocarbon with steam to produce a hydrogen-rich gas (reformate). Using methanol as an example feedstock, steam reforming can theoretically produce a maximum three moles of diatomic hydrogen gas for every mole of carbon involved in the reaction (3:1). Steam reforming processes can occur with or without a catalyst. 1 However, most low-temperature fuel cell systems that use steam reformation make use of catalysts in the reformer. Whereas non-catalytic steam reforming requires temperatures in excess of 1100 o C for simple hydrocarbons, catalytic steam reforming can occur at temperatures ranging from 00 o C to 700 o C for simple hydrocarbons. Nonetheless, non-catalytic steam reforming is still under investigation as a possible method of reducing the thermal mass and volume of reformers. Equation -1 shows the general, ideal steam reforming reaction for alcohols. C H l m O n m + ah O + a H + ( n + a l) CO + ( l n a)co (-1) l, m, and n represent the number of moles of carbon, hydrogen and oxygen, respectively, in the hydrocarbon being reformed. The variable a represents the number of moles of water per mole of hydrocarbon reacted. Greater concentrations of water (a) in the reactants reduce the concentration of CO, and increase the quantity of CO in the products. Moreover, the amount of hydrogen is 1 Non-catalytic steam reformation requires very high temperatures.

21 9 increased with greater steam. The maximum concentration of water that can be reacted in the general steam reforming reaction is amax = l n (-) In practice, the catalytic steam reforming reaction is dependent on three-phase chemical kinetics with mechanisms that are complex and currently not well understood. Modern steam reforming catalysts can be designed to suppress the formation of CO and promote the formation of CO. The most commonly used methanol steam reformer catalysts in published literature are composed of copper-oxide and zinc-oxide (CuO/ZnO). As a general rule, the steam reformation of methanol across a CuO/ZnO catalyst is executed with 1 to 1.5 moles of water per mole of methanol. Increasing the H O concentration to 1.5 has been found to reduce the CO content in the products while increasing the hydrogen yield when the reformation process is maintained in the range of 50 o C-300 o C. Although higher water concentrations in the fuel increases hydrogen yield and decreases reformate CO concentrations, these benefits may be counterbalanced by the energy penalty associated with water vaporization and superheating. Partial Oxidation Reforming Partial oxidation reforming is a process in which the fuel undergoes combustion in a fuel-rich environment. The elevated temperatures generated from this combustion breaks down the hydrocarbon fuel into a hydrogen rich gas. Assuming, ideally, that the reaction products are only hydrogen, carbon monoxide and carbon dioxide, the general partial oxidation reaction for an alcohol is given in Equation -3. C H O m + ao e H + ( n + a l e) CO + ( l n a + e) CO + eh O l m n (-3)

22 10 To obtain the maximum possible amount of hydrogen in a partial oxidation reaction, the amount of water (e) formed must be negligible (e 0). Taking methanol as an example, the maximum, ideal, diatomic hydrogen to carbon ratio for the partial oxidation reaction is :1. This is /3 of the maximum possible obtainable hydrogen to carbon ratio from steam reforming of methanol. It is important to note that for most applications the use of pure oxygen for partial oxidation is impractical. Thus nitrogen will be present in the products of the reaction, which further dilutes the reformate hydrogen concentration. Some advantages of partial oxidation over steam reforming are that water is not required for the process, the equipment is more compact, it allows rapid start-up, the same reactor can reform various types of fuel, and the high-temperature reactor can tolerate many impurities (including sulfur). On the other hand, care must be taken to prevent carbon deposition. Also, since the process is exothermic, relatively high local temperatures are produced, causing possible materials problems. For integration with lower temperature fuel cells, a gas clean-up step is necessary to reduce the inherently large CO content. Autothermal Reforming Autothermal reforming combines aspects of steam reforming and partial oxidation reforming. A catalyst bed, typically nickel, and steam promote the reforming reaction. The necessary heat is generated through combustion of a portion of the reformer fuel. The general, idealized autothermal reforming reaction is given by Equation -4. C H l m O n m + ao + bh O + b H + ( n + a + b l) CO + ( l n a b) CO (-4)

23 11 In autothermal reforming, the maximum amount of water (b) that can be reacted is given by Equation -5. b = l n a (-5) max Compared to the ideal steam reforming reaction, the maximum amount of water that can be reformed is reduced by the presence of oxygen, thus reducing the overall hydrogen yield. On the other hand, the addition of water to the reaction allows for higher hydrogen yields and lower carbon monoxide concentrations in the products. Autothermal reformers have had a large amount of interest because of their capacity to reconcile between the benefits and detriments of using either catalytic steam reforming or partial oxidation reforming. Transportation Fuel Cells For transportation applications and small-scale residential power production, PEMFCs and PAFCs have been most widely considered and explored. These fuel cells operate similarly, oxidizing hydrogen through an electrolyte and a current collector as shown: Anode: H 4e - +H + Cathode: ½ O + 4e - +H + H O However, PEMFCs operate at a lower temperature than PAFCs (5 o C-80 o C as opposed to 150 o C to 00 o C). In addition, PEMFCs, due to their high efficiency over a wide range of current densities, typically have higher energy density than PAFCs. Mainly for this reason, PEMFCs are widely considered the preferred automotive fuel cell. This distinction does not mean that PEMFCs are trouble-free. Major engineering challenges are yet to be resolved in order to commercialize automotive PEMFC stacks. Foremost,

24 1 these fuel cells must dramatically improve their operational life; they must achieve consistency in performance under mass production, which has not yet been demonstrated; they must be designed to have effective humidity control and positive water balance during all operating regimes; and their cost must be dramatically reduced. In terms of life and humidity control, PAFCs have been demonstrated to far surpass PEMFCs. However, PAFCs are much more expensive to produce than PEMFCs and their cost is harder to reduce to target levels, even through mass production. Figure -1 shows typical polarization curves of various types of fuel cells including molten carbonate fuel cells (MFCs), solid oxide fuel cells (SOFCs), and alkaline fuel cell (AFCs). The reader might find it curious that AFCs are not considered in this study even though they exhibit much higher voltages than all other fuel cells. The reason for this is that AFCs do not tolerate air (due to the presence of CO ). This limits their use to space applications and as electrolyzers. The Transportation Load Environment The power load associated with automobiles and buses is high and extremely transient. Vehicles are driven with a rapid succession of speed adjustments, especially in city traffic. Any fuel cell engine that is going to be incorporated into a vehicle must be able to change power delivery over a wide range of conditions fast and efficiently. Figure - shows driving loads obtained from a 30 ft fuel cell transit bus operated at the University of Florida (UF) on a low speed circuit around the city of Gainesville, Florida. The transient response of fuel cell systems is limited by their capacity to deliver fuel (hydrogen) and oxidant (oxygen) to the fuel cell stack. In a reformer based fuel cell system, as shown in Figure -3, the fuel processing subsystem must be capable of changing hydrogen production levels at the same rate as current draw changes occur in

25 13 the fuel cell stack. In turn, fuel cell voltage at a certain current draw depends greatly on H and CO partial pressures in the reformate gas. The process of steam reforming is complicated by the endothermicity associated with hydrogen production. The composition of the reformate gas is largely dependent on the temperature at which the reforming reaction takes place. Yet, in steam reformers the heat required to maintain the reforming reaction is provided by an external source and thus is limited by heat transfer considerations. This results in varying reforming reaction temperatures during steady-state and transient operation. As an example, the University of Florida fuel cell bus (TBB-) reformer temperatures during transient conditions are shown in Figure In TBB- the reformer catalyst bed is heated via an external burner that consumes stack anode flue gas (Figure -5). The anode flue gas contains varying quantities of hydrogen, depending on stack current draw and anode stoichiometry. Anode stoichiometry is dependent upon fuel flow rate and reformate quality. Current-draw depends on the load environment. For TBB- the stack power draw and the reformer hydrogen production are coupled. University of Florida researchers have found that this interrelation between the stack current draw and reforming heating has led to instances of reformer overheating and under-heating, both of which have led to drops in overall system performance []. 1 Reformer top, center, and bottom refers to the position of thermocouples in different regions of the catalyst bed. The catalyst bed entrance is located close to the reformer bottom thermocouple. The center thermocouple is located close to the midpoint of the reformer and the top thermocouple is located close to the exit of the reformer. The reformer burner thermocouple is located close to the reformer burner flame.

26 14 Figure -1. Typical polarization curves of various types of fuel cells. This figure was published with the permission of its original authors [4] test Power (kw) Transient Load Profile for TBB- under the UFFL driving cycle Run Time (s) Figure -. Driving load of a 30 ft phosphoric acid fuel cell bus operated at the University of Florida. The bus weighs is approximately 30,000 lbs.

27 15 Parasitic and Aux. Loads Methanol:Water Premix Tank Steam Reformer Phosphoric Acid Fuel Cell Stack Power Conditioning High Voltage Distribution Traction Drive Batteries Wheels Figure -3. Indirect methanol automotive fuel cell system power delivery schematic Idle Ramp Up Period Steady State High Power Steady State, Idle 550 Ramp Down Period 500 Temperature ( o C) Reformer Burner Reformer Bottom Reformer Top Reformer Center 00 00: :1.0 14:4.0 1:36.0 8: : :1.0 50:4.0 57:36.0 Time (minutes) Reformer Top Reformer Center Reformer Bottom Reformer Burner Figure -4. TBB- fuel cell bus methanol steam reformer and steam reformer burner temperatures during transient bus operation

28 16 Combustion Products Anode Flue Gas Air Reformate Bypass Fuel Cell Stack Anode Air Reformer Burner Steam Reformer Excess Air and Water Cathode Heat Exchange Plate Vaporizer Neat Methanol Water Methanol Premix Air Heat Exchanger Start-up Burner Combustion Products Figure -5. University of Florida bus indirect methanol fuel cell system CO and Hydrogen Concentration Effects on PEMFC performance Small increases in reformate CO concentration lead to drastic reductions in fuel cell performance in low-temperature fuel cells that use platinum catalyst to promote the cell anodic reaction. In these cells, CO competes directly with hydrogen for active sites in the anode due to its high affinity for platinum at low-temperatures. Springer et al. propose that the reactions depicted in Equation -6, Equation -7, Equation -8 and Equation -9 represent well the competition for active sites at a PEMFC anode [5]. kfc / bfc CO + M ( M CO) (-6) kfh / bfh H + M ( M CO) (-7) keh + ( M H ) H + e + M (-8) H kec + O + ( M CO) M + CO + H + e (-9)

29 17 Equations -6 and -7 represent the competition between H and CO for active sites, where kfc and kfh represent the forward reaction-rate constants. The electrooxidation of hydrogen and CO are represented in Equations -8 and -9, with their respective rate constants (keh and kec). PEMFC models and experimental results have demonstrated that drops in anode hydrogen concentration produce very mild decreases in cell performance (Figure -6). However, high CO concentrations have been shown to drastically drop cell performance, especially when diluted in reformate bearing low hydrogen concentrations (Figure -7). In order to further reduce the CO concentrations produced from steam reformers, many fuel processing systems use preferential oxidizers and anode air bleeds (the introduction of small quantities of air into the fuel cell anode chamber). Preferential oxidizers require the addition of air into a catalytic bed that promotes the oxidation of CO. Typically a small percentage of the hydrogen in the reformate is also oxidized. Currently, control of air delivery systems for preferential oxidizers is complicated due to the high sensitivity of the chemical rate equations to gas and catalyst temperature and CO concentration. In addition, the amount of air associated with preferential oxidation operation is very low and difficult to meter. In the overall reforming process high upstream variations in CO concentrations may lead to increased concentrations of CO leaving preferential oxidizers, thus increasing the probability of anode catalyst poisoning. This is especially problematic since CO spikes occur during periods of increasing power demand. This phenomenon will be further explained in subsequent chapters. Various solutions have been proposed for solving the problem of CO poisoning of the fuel cell anode catalyst. Of these the most commonly reported has been the

30 18 introduction of small quantities of air into the cell anode during periods of decreased cell performance due to CO poisoning. This produces catalytic oxidation of the CO in the presence of anode platinum. Nonetheless, the air bleed also has a tendency to oxidize a portion of the hydrogen introduced into the fuel cell. This results in a further reduction in cell efficiency. Figure -8 shows how the effect of air bleed introduction in the anode reduces the poisoning effects of CO. Also, it is difficult to distinguish between cell performance deterioration due to CO poisoning and other operating conditions (humidity variations, temperature variations, pressure variations, cell degradation, etc). Based on this, a reformer that produces minimal increases in CO level during increased fuel flow events is more desirable than one that produces high hydrogen concentrations (within typical reformate hydrogen and carbon monoxide concentrations). Figure -6. Effect of hydrogen concentration and utilization on PEMFC performance. This figure was published with the permission of its original authors [5].

31 19 Figure -7. Effect of CO poisoning on PEMFC performance. This figure was published with the permission of its original authors [5]. Figure -8. Use of air bleed to recover PEMFC performance. This figure was published with the permission of its original authors [5]. Methanol as Fuel Since most of the results of this research are based on the study of methanol reformation, a brief discussion of methanol as a potential fuel cell fuel is relevant.

32 0 Of the alternative fuels which are being considered for vehicles, many view hydrogen as the ultimate long-term alternative fuel. However, while hydrogen has some very desirable attributes, there are many extremely difficult issues to resolve, which might make the time scale for a hydrogen economy far longer than the time scale of readily available and affordable petroleum. Fuels which can be stored as liquids offer major advantages for transportation applications. Of the alternative liquid fuels, in many ways the most promising is methanol. This is because methanol reforming technology is well developed as compared to reforming technology for other common automotive fuels. Also methanol can be produced from natural gas, biomass, electricity, coal, and any other hydrocarbon. These result in a fuel that can potentially be relatively inexpensive in most regions of the globe. In addition, methanol can be used in internal combustion engines, gas turbine engines and in fuel cell engines. Most of the methanol produced in the U.S. is derived from natural gas. Twelve major methanol production plants exist in the U.S., which in 001 produced over 1.5 billion gallons of methanol (94. trillion BTU or equivalent to million gallons of gasoline). The overall U.S. methanol consumption in 001 was, in energy terms, equivalent to ~1,507 million gallons of gasoline, half of which was imported [6]. The price of delivered methanol has averaged ~$0.50 per gallon (~$8/million BTU) since If the price of natural gas did not figure into the cost of methanol, the methanol cost would be ~$1.50/million BTU. Currently, most of the U.S. produced hydrogen is derived from natural gas steam reforming. The U.S. consumes around 9 million tons of hydrogen (95.8 trillion BTU or

33 1 equivalent to ~7,407 million gallons of gasoline) per year, of which approximately 17% is sold to chemical plants and refineries [7]. Current hydrogen costs at large scale chemical plants, where hydrogen is consumed on site, are generally around $5.31/million BTU. If transported, the price of hydrogen is much higher. The typical price of delivered liquid hydrogen oscillates between $1.00 and $1.40 per pound ($0/million BTU to $7/million BTU). Therefore hydrogen prices are comparable to methanol prices only if hydrogen is not pressurized, liquefied, or transported. The cost of producing hydrogen using steam reforming of natural gas minus the cost of the natural gas, for a large plant, would be ~$4.15/million BTU. This price does not include transportation, storage, compression or liquefaction of the gas. While for methanol, storage and transportation costs are not essential components of the overall fuel cost, this is not the case with hydrogen. It is expected that, due to its lower energy density, methanol transportation costs per unit energy basis will be twice the cost for transporting gasoline, in energy terms (~$6.9 per million BTU) [7]. Hydrogen transportation cost, in contrast, may vary drastically depending on the distance transported and the method of transportation. Taking current gasoline transportation costs as a baseline, it could be assumed, as an initial estimate, that because of its lower energy density, liquid-hydrogen transportation would cost at least 5 times more than gasoline transportation if similar transportation forms were used (mainly pipeline and truck transport). Thus, hydrogen transportation costs could be expected to be around $18.94 per million BTU [7].

34 Hydrogen distribution costs could also be estimated by assuming that the eventual hydrogen distribution infrastructure will be similar to that of natural gas. This assumption is legitimate since both fuels are gaseous. The average distribution cost for natural gas, from wellhead to residential consumer 1 in the U.S., from 1991 to 003, was $4.80/million BTU. Thus, by applying an energy density penalty it could be estimated that the average hydrogen distribution price in the U.S. will be ~$18.51/million BTU (similar to the previous estimation). Note however that these estimates do not include additional cost penalties associated with hydrogen transport such as boil-off, gas leakage, and special materials needs. Based on this basic analysis, it is fair to assume that methanol could be an important fuel for transportation and fuel cells since it has the potential to be much cheaper and more easily distributed than hydrogen. A study by the University of Florida found that based on future fuel cell vehicle fleet estimates for 00, and according to various economic scenarios, methanol derived hydrogen from coal would be the least expensive automotive fuel for fuel cell vehicles in the US if environmental concerns are not taken into account [7]. 1 Residential natural gas prices were used to estimate hydrogen distribution prices because it most closely resembles the distribution requirements of the automotive sector.

35 CHAPTER 3 LITERATURE REVIEW This literature survey has the purpose of framing the work and results of this dissertation as part of the latest research and scientific debates regarding the design and integration of methanol steam reformers into fuel cell systems. Generally, research conducted on reformers can be catalogued into two different approaches, the chemical and the physical. Chemical research has concentrated in trying to understand the chemical mechanisms and rate of reaction relations of the steam reforming reaction. A large volume of papers and presentations have been published in the last couple of years regarding this subject. Common among these is a general concern among researchers that almost every research case (even when identical catalysts are used) has lead to different rate of reaction equations, although in the same mathematical form. The physical approach to reformer research has dealt with understanding the complete reformer assembly and its behavior. The volume of work in this area is limited. In addition, most of the work has concentrated on the steady-state case. Dynamic models and studies on fuel cell systems in published literature typically rely on user-defined linear exponential decay functions for each fuel cell system component, including the reformer. These investigations concentrate on the question of fuel cell systems control. Typical of this approach is the study conducted by El-Sharkh et al. 004 [8]. 3

36 4 Kinetics of the Catalytic Methanol Steam Reformer Reactor Although the actual kinetics of the methanol steam reforming reaction were outside the scope of this study, a general review of the work in this field is herein presented. Of note is the diversity of rate equations that have been presented in literature. This demonstrates the complexity of the reaction and the variability that can be obtained based on catalyst selection, catalyst preparation, and reaction conditions. Choi et al., 005, proposed that the methanol steam reformer products can be obtained by with the rate equations for 3 main reactions in a Cu/ZnO/Al O 3 catalyst manufactured by Sud-Chemie. These were [9], CH OH CO + H 3 CH OH + H O CO 3 CO + H O CO + H + 3H The study conducted by Choi greatly simplified the widely cited 1 reaction rate model previously given by Peppley et al. [10]. Peppley and Choi point out in their studies the large degree of disagreement in literature regarding reaction mechanisms and the debate on identifying the physical location of the steam reforming catalyst active site. The results of their experiments were a series of rate equation for each of the proposed methanol steam reformer reactions. Based on these reactions a methanol steam reformer, a water gas shifter and a PROX reactor model were developed. The results of their models show an increase in conversion with greater temperature and catalyst weight in the reactor, however this also increased CO. The study conducted by Choi did not model heat transfer in the reactor bed or the transient behavior of the reformer. Lee et al., 004, [11] recognized the importance of steam reforming kinetics in sizing methanol steam reformers. They also recognized the large diversity of rate

37 5 equations found in literature. Lee, in his paper includes a list of 9 different catalytic methanol steam reforming rate of reaction equations found in recent literature. Many of these reaction rate equations are contradicting even when the same catalyst is being used. Lee observes that higher water concentration in the steam reforming reaction promotes effluent CO concentrations closer to equilibrium. The discrepancy between experimental CO concentrations and equilibrium concentrations were between ~1mol% to <0.1mol% for a water feed concentration between 15mol% to 30mol%. Finally Lee, in his study, also proposed a new rate of reaction of methanol for the catalyst used. The correlation proposed by Nakagaki et al. [1] was used in the models produced in this study. Nakagaki obtained his correlations through a series of experiments that closely resembled the test conditions in which data was obtained for this study. The study conducted by Nakagaki used a temperature controlled packed-bed reformer operated at various fuel flow rates. In obtaining his correlation, Nakagaki conducted constant catalyst temperature studies and constant wall temperature studies (the latter allowed for temperature gradients to exist in the reformer catalyst bed). Finally his correlation was validated through actual reformer operation and modeling. This type of experimentation varies from others in literature in that the effects of fuel flow (transport phenomena) and varying reactant concentrations were considered. Also considered are the effects of temperature gradients in the catalyst. The following is a general description of the study conducted by Nakagaki, et al. Nakagaki et al. obtained their correlation through experiments conducted in a packed-bed experimental reformer. The catalyst used was Cu-Zn in alumina support commercially available from Nissan (G66-B). The catalysts were 1/8in by 1/8 in

38 6 cylindrical tablets. Various concentrations of methanol diluted in simulated reformate gas assuming conversions equivalent to 0, 10, 0, 50, and 75 percent were flowed through the packed bed reformer. The reformer temperature was maintained constant and controlled through the use of sheath heaters. This first test was used to determine the methanol rate of reaction in a constant temperature reaction. A second test controlled the temperature of the reforming reaction by maintaining a constant wall temperature. In the second test the temperature distribution and the conversion were recorded. The exiting reformate gas composition was recorded using a TCD gas chromatograph. Based on second test results, the Nakagaki correlation was refined. Table 3-1 shows a summary of the tests conducted by Nakagaki. Nakagaki also conducted modeling studies through a combined energy, momentum (Darcian equation), and conservation of species equations to further validate his correlation. However, since the application of the study was for employment in gas turbine engines, the study of CO concentration in the reformate outlet stream was neglected. Table 3-1. Test parameters for Nakagaki tests Parameter Test 1 Test Pressure (atm) 1, 5, 10, 15 1, 10 Catalyst Temperature (K) 473, 498, 513, Wall Temperature (K) , 548 Mass Flux (kg/(sm )) , 0.36 Simulated conversion percentage 0, 10, 0, 50, 75 0, 80 Reformer Performance Various studies have been published on reformer performance and design but the majority of these have concentrated on reformer steady-state performance and catalyst

39 7 dynamic behavior (rates of reaction correlations). The work of Betts et al. [13] provided a mechanistic model for observed fluctuations in reformer CO output. The identification of CO spikes as being caused primarily by heat transfer effects in the reformer was discussed in that study. A study by Horng [14] showed for an autothermal reformer the existence of CO spikes once fuel was introduced into the reformer. However, the study concentrated in the start-up process of the reformer and did not place much emphasis on the CO phenomenon. The importance of heat transfer and temperature on the performance of methanol steam reformers was studied by Hohlein et al. [15]. Hohlein concludes that reformer heating is a very strong determinant of reformate CO concentration. Perhaps the most comprehensive experimental study on reformer performance was conducted by Davieau [16]. In this study, reformer performance was measured in terms of methanol conversion for two reformers of differing diameters. The initial interest of Davieau was to study the degree in which space velocity and aspect ratio determines reformer similarity. The reformers were operated at various space velocities. The results demonstrate that space velocity and aspect ratio were not a good determinant of reformer performance. That is, two reformers operating with the same space velocity, same catalyst loading, same wall temperature and same aspect ratio had drastically different performances. Yet, Davieau points out that space velocity is the most cited reformer similarity relation in literature. Additionally, Davieau notes that temperature distribution within the reformer is a critical performance parameter. Furthermore, it is noted that diameter influences heat transfer and therefore has an effect on performance that is

40 8 beyond catalyst loading. The results obtained by Davieau were based on experimental observations. His study did not focus on the transient behavior of the reformer.

41 CHAPTER 4 ANALYSIS In accordance with the objectives of this research, a general methanol steam reforming model was developed. The details of the derivations and major assumptions applied to the development of this model are included in this chapter. In addition, this chapter includes a general comparison of the steam reformer model results and experiments conducted at UC-Davis using an experimental reformer test rig developed at by Dr. Paul Erickson (Appendix C). Part of the analysis includes the development of a generalized design and comparison criteria for steam reformers through similitude parameters. Finally, in this chapter transient operation of the modeled steam reformer is presented. Steam Reformer Model Packed bed catalytic steam reformers are generally composed of a reaction chamber filled with catalyst pellets (Figure 4-1). The methanol water mix along with the product gases flows around and through the catalyst pellets. The chemical reactions occur at the surface and within the catalyst pellets. In the development of this model, it was recognized that the main driver of reformate chemistry is temperature. A control volume approach was used to obtain differential conservation of species and conservation of energy relationships. In choosing the scale of the finite differential control volume, it was assumed that the catalyst pellets and the gases exist in the control volume as a homogenous mixture. That implies that the control volume used, although small compared to the overall reformer, has length scales larger than that of the catalyst. 9

42 30 It was also assumed that the gas and the catalyst exist locally at the same temperature. This comes as a result of the assumption that the catalyst and the gases are regarded as part of a continuum. This also implies that the surface area of the catalyst inside the control volume is very large, resulting in the gas being in thermal equilibrium with the catalyst. In addition, the flow was assumed to occur in the axial direction (i.e. the radial component of the velocity was assumed negligible). Figure 4- shows the control volume with the major heat inputs and outputs at the boundaries. Steam Reformer Walls Methanol + Water Reaction Chamber Reformate Catalyst Pellets Figure 4-1. General schematic of modeled steam reformer reactor Figure 4-. Control volume diagram

43 31 Heat Transfer Based on the chosen control volume and the aforementioned assumptions, the heat transfer equation can be derived (Equation 4-1). T keff T T T deg T ρ eff Ceff = + keff + keff + uρgas( 1 ε ) C p, gas (4-1) t r r r x dv x The variables in Equation 4-1 are defined as follows: ρ eff is the effective mass density of the catalyst-gas mixture, C eff is the effective specific heat of the catalyst-gas mixture, T is the temperature inside the control volume, k eff is the catalyst-gas mixture effective conductivity, r is the radial coordinate, x is the axial coordinate, d is the net rate of heat absorbed by chemical reactions, V is the reformer volume, ρgas is the density E g of the gas, e is the void factor of the reformer, and C, is the average specific heat of p gas the gas. The thermal capacitance of the mass inside the control volume is assumed to be mostly because of the catalyst. Therefore ρ eff C eff is approximately equal to the catalyst density (ρ c ) multiplied by the catalyst specific heat (C c ) and the void factor (ε). The use of effective properties for modeling of catalytic reactors is a fairly common practice. Incoprera et al., in their widely used heat transfer text book, Fundamentals of Heat and Mass Transfer, describes heat transfer correlations for packed bed systems using effective gas/packed-bed properties [17]. In published literature, various methods for determining packed-bed effective thermal conductivity have been explored. Typical of this work are the studies conducted by Munagavalasa et al. [18], and Brucker et al. [19].

44 3 The convection term in Equation 4-1, uρ C gas p, gas T x, for modeling purposes was assumed to be small as compared to the radial conduction terms. Typically, temperature gradients in the radial direction are typically higher than those in the axial direction through most of the steam reformer. A possible exception to this could occur at the entrance of the reformer if the entrance temperature is significantly different than the reformer catalyst bed temperature. As an example of the disparity between the convection and conduction terms, an order of magnitude analysis of these, based on data obtained from the UC-Davis reformer (see Appendix C), is presented in Figure 4-3. The analysis is executed for a run in which various changes in methanol flow were executed. The results were obtained from temperature data recorded at the entrance regions of the reformer. A finite difference approximation of the derivative and second derivative of temperature in the axial and radial directions, respectively, were calculated using K-type thermocouple readings. The radial thermocouples were placed at the wall, at half the radius and at the midpoint, where the radius of the tube was 11.9 mm. The axial thermocouples were 67.6 mm apart. It is recognized that the significantly larger axial distances used to approximate the finite difference axial temperature gradient leads to greater averaging. This in turn may lead to an underestimation of the axial gradient in the entrance region. In spite of its limitations, this analysis provides a general relative magnitude of the axial and radial temperature gradients. During most of the run, the convection term did not exceed 0.4% of the value of the conduction term. This, in order of magnitude terms, supports the decision to disregard the convection term in the energy equation. The UC-Davis reformer is discussed in greater detail at the latter sections of this chapter and in Appendix C.

45 33 Figure 4-3. Convection-conduction ratio in the UC-Davis reformer fuel entry regions The heat dissipated in the catalyst bed is due to the reforming reaction. Consequently, d E g [ ] is a function of the rate in which methanol is reacted in the steam reforming reaction, d CH 3OH dt. This in turn is a function of the temperature (T), catalyst volume (dv c ), and the molar concentration of methanol in the reformate x 3. The rate of methanol consumed per unit mass of catalyst (r M ) and can be gas ( ) CH OH obtained using Equation 4- [ CH OH ] ( r dr) [ CH OH ] 1 d 3 1 d 3 r M = = = m dt πρ drdx dt c c f ( T, P, x CH3OH ) (4-) The calculation of the rate of methanol reacted in a CuO/ZnO catalyst for methanol-steam reforming reactions was extensively studied by Nakagaki et al. at Toshiba Power Systems [1]. One of the results of this study has provided a good correlation for r M, shown here:

46 34 r k M o = k o m l T P 513 e E RT 6 = mol /( g n CH3OH s atm) m = 10 if T > 513K otherwise it's equal to 0 5 E = 1 10 J / mol l = 0.13 n = 1.3 x cat (4-3) The Nakagaki correlation is an empirical variation of the Arhenius equation and was obtained experimentally. Further information regarding the Nakagaki study is found in the literature review chapter of this text. Given these relationships, d E & g can be defined explicitly (Equation 4-4). d [ CH OH ] d E g = E 3 r = cπdxdr(r + dr) dt ερ E r (4-4) where the expression for the heat of reaction per mole of methanol consumed is expressed as E r Therefore the energy equation for the control volume can be rewritten as Equation r M T t α T = r r T + α r T + α x Err C c M (4-5) The quantity E r r M is a function of temperature, thus Equation 4-5 is non-linear. This non-linearity is the source of the complex and sometimes unexpected behavior associated with packed bed steam-reformers. The boundary conditions for the heat equation (Equation 4-5) are

47 35 T r T x r= 0 T ( r = R) = T T ( x = 0) = T x= L = 0 = 0 T ( t = 0) = T i w g, in (4-6) Given these boundary conditions the energy equation can be solved by using a finite difference scheme or using finite element analysis. Also, Equation 4-5 can be made non-dimensional to yield: T 1 T T R T * * * * M r = + + * * * * Fo r r r L x αtw T r x αt T =, r =, x =, Fo= * * * * Tw r L R R r E (4-7) For purposes of this study, the heat transfer model was solved in dimensional terms in order to relate results to real physical parameters which can be readily verified through experimentation. Continuity The fluid flow within the steam reformer tube is transient, compressible and viscous. If it is assumed that the mean flow travels in the axial direction (u x >> u r ), then the continuity equation can be written as Equation 4-8. ρ g t = ( ρ u ) g x x (4-8) where u x is the axial component of the fluid velocity vector, and ρ g is the density of the fluid. If the fluid is assumed to be an ideal gas, then the continuity equation can be written in terms of fluid temperature. In addition, it is convenient to use mass flow rate,

48 36 m&, instead of the velocity because cross-sectional averages of the velocity will be used. This yields m Acs P T = x RT t (4-9) In essence Equation 4-9 states that variations in the axial mass flow are a function of the transient change in temperature at any given location inside the reformer. However, reformer data collected from the UC-Davis reformer suggests that T t for methanol steam reformers is small even at the entrance, centerline region (Zone 1 and Zone ). This is because the thermal capacitance of reformer catalysts is relatively high, and because the primary mode of heat transfer is through conduction, which is relatively slow. Figure 4-4 shows recorded values of T t from the UC-Davis reformer. For illustration purposes, even if the maximum value exhibited was held at ~ K/s (the maximum recorded from the UC-Davis reformer during reformer warm-up) and assuming that this level of heating were to be carried out through out the entire reformer, the change in reformate mass flow rate that would be detectable at the reformer exit would be 0.018% for a 1 meter long reformer, based on Equation L AcsP T m& out m& in = dx (4-10) 0 RT t Therefore, the mass flow rate within the reformer will have very small transient variations. Despite this, Equation 4-10 was used in the elaboration of the reformer model. Temperature distribution plots for the UC-Davis reformer at varying flow rates are shown in Figure 4-5. Note: for modeling purposes the pressure drop in the reformer was neglected.

49 37.5 UC-Davis Reformer dt/dt values Calculated dt/dt [K/s] Run Time (min) Zone1 dt/dt Zone dt/dt Figure 4-4. Calculated values of dt/dt for the UC-Davis reformer at the entrance region Figure 4-5. Measured temperature distributions of the UC-Davis steam reforming test rig operating at varying space velocities and varying conversion efficiencies

50 38 Conservation of Specie The continuity equation for each of the component gases in the control volume can be written as Equation rate of increase mass flow of j mass flow of j rate of production of mass of j in into the control out of the control of j inside the control = + the control volume volume volume volume or m t j = m j, in m j, out + m j, gen (4-11) The component gases in the steam reformer products are assumed to be methanol, water vapor, carbon dioxide, carbon monoxide, and hydrogen. The mass generated of component j inside the control volume can be estimated using the Nakagaki correlation (Equation 4-3) and chemical equilibrium. Thus, for a control volume at a certain temperature and pressure, the rate of methanol reacted is known and, consequently, the rate of generation of the product gases is also known. The rate of mass of specie j entering the control volume is generally known from boundary conditions. However the rate of mass escaping the control volume must be calculated. For purposes of the model it was assumed that all species flow at the same axial velocity as the bulk flow. The bulk flow velocity exiting the control volume can be expressed as Equation 4-1. u out m& out = ρ A out cs = m& outrt PA cs AcsP T dx m + & RT t = PA cs in RT = m& inrt PA cs L 1 + T 0 T T dx = u t j, out (4-1) Consequently, the mass flow rate of the exiting gas j is given by Equation 4-13.

51 39 P A m & = RT & j cs j, out = u j, out x jmout (4-13) In the above equation, P j is the partial pressure of gas j inside the control volume. The term x j is the molar concentration of gas j, and is defined as Equation n j x j ( t = to ) = N (4-14) n j= 1 j t= t o Chemical Model The chemical reaction for steam reforming of methanol being considered is shown in Equation CH3 OH + xho ah O + bco + cco + dh + ech3oh (4-15) In a control volume a certain rate of methanol can be reacted (r M m c ). The remaining methanol and water that enters the control volume is not reacted (Figure 4-6). It was assumed that the water-gas shift reaction is not the time limiting step in the reformer reaction and thus proceeds to equilibrium rather quickly. For modeling purposes the water gas shift reaction was assumed to reach equilibrium concentrations, while the reforming of methanol was bound by the rate of reaction constraints of the control volume. Similar approximations have been made in literature and the work of Rabou et al. [0] is an example.

52 40 Methanol (mol/s) Water H CO CO CV reacted (mol/s) Not reacted (mol/s) H CO CO Methanol (mol/s) Water H CO CO Figure 4-6. Schematic representation chemical reaction model in the control volume. The chemical equilibrium fractions for the water gas shift reaction can be calculated in the following way. The specie balance of the complete steam reforming reaction (e and f are equal to zero) leads to a = + x d (4-16) b = d (4-17) c = 3 d (4-18) The solution for d can be obtained from the water-gas shift reaction (Equation 4-19) H + + (4-19) O CO H CO The equilibrium constant for the water gas shift reaction can be approximated with the Moe (196) correlation [1] (Equation 4-0) ln K p = (4-0) T where T is the temperature of the reaction in Kelvin. Given this,

53 41 [ x Kp xkp( Kp 4) + Kp + 4Kp + 4 xkp 5Kp + ] d = (4-1) ( Kp 1) With these equations the number of moles of each chemical component per mole of methanol reacted can be obtained as a function of temperature. The number of moles of methanol reacted can be obtained from the Nakagaki correlation (Equation 4-3). The equilibrium heat of reaction and products as a function of temperature for the methanol steam reforming reaction is given in Figure 4-7. Gas Composition [kmols/ kmol of Methanol reacted] Equilibrium Reformate Composition Magnitude Depicted in Secondary Axis Temperature (K) CO H CO Hr [kj/kmol of MeOH] Figure 4-7. Equilibrium products of the steam reforming reaction (CH 3 OH + H O ah O + bco + cco + dh ) with respect to temperature Finite Difference Solution The control volume analysis easily leads to a finite difference solution of the heat transfer and fluid flow problem presented in the previous sections. A grid was set up where each element has a temperature (T m,n ), in which m represents the node index in the Heat of Reaction [kj/ kmol of Methanol reacted]

54 4 x-direction and n represents the node index in the r-direction (as shown in Figure 4-8). The derivatives of the catalyst bed temperature were approximated to be T x T r T T T T r m+ 1, n m, n+ 1 m, n+ 1 + T T r ( x) + T m 1, n m, n 1 ( r) m, n 1 T T m, n m, n (4-) Similarly the time derivative of the temperature could be written as T t T l+ 1 m, n T t l m, n (4-3) where l is the time index. R r 0 x r m,n+1 x r m,n x r m+1,n x Figure 4-8. Separation of the reformer volume into annular elements with constant gas properties in space for the finite difference solution method used Given these equations, Equation 4-5 could be rewritten as

55 43 ( ) ( ) ( ) c l m n M r l m n l n m l n m c c l m n l m n l m n l m n l m n l m n l m n C r E x T T T C k r T T T r T T r t T T 1,, 1, 1,, 1, 1, 1, 1., 1, = K K ρ α α (4-4) Similarly the boundary conditions can be rewritten (Equation 4-5). i m n l n M l n M in g l n l N m l N m w l m T T T T T T T T T T = = = = = 0,,,, 0,,,,0 (4-5) The term is unknown and it is a function of the finite volume temperature and methanol concentration. At first, the temperature equation can be solved assuming a temperature and a methanol concentration. For the computer program developed for this study, T ( ) 1, l+ m n E r r M m,n l+1 was initially assumed approximately equal to T m,n l. Similarly the concentration of methanol inside the volume, x CH3OH,m,n l+1 was at first approximated to be equal to x CH3OH,m,n l. Once an approximate solution for the temperature was obtained the conservation of specie equation was solved. The finite difference expression for 4-10 is given by 1,, 1,, 1 1,,,,,, = l gen m n l m n j l n m j l m n j l l m n j m m m t m m & & & (4-6) The expression for the mass flow of specie j leaving the finite volume can be expressed as

56 44 l+ 1 l+ 1 l l+ 1 l ( ) ( ) l+ 1 x j, m, n Acs P x Tm 1, n Tm 1, n Tm, n Tm, n m& = j, m, n 0. 5 (4-7) l+ 1 l+ 1 R t Tm 1, n Tm, n Note that as with the temperature solution, x l+1 j,m,n is unknown. In the computer program this value was initially assumed to be equal to x l j,m,n. From the solution of the content of all the species in the finite volume using Equation 4-4 and 4-7, for a given temperature field, a new value for the molar concentration of each component can be obtained (Equation 4-8). l+ 1 m j, m. n MW l+ 1 j x = j, m, n (4-8) N l+ 1 m i, m. n i MWi This new, revised solution for x l+1 j,m,n was used to solve for the finite volume temperature and mass contents. This process was carried out until satisfactory convergence was achieved. FEMLAB Solution Besides using a finite difference scheme, the system of differential equations was solved using finite element analysis using FEMLAB Multiphysics modeler software. The FEMLAB software is designed to solve systems of coupled partial differential equations using proprietary solution schemes based on finite element analysis. The system of partial differential equations for transient heat equation (including the convection term), transient conservation of species, transient continuity, chemical equilibrium and the Nakagaki correlation were programmed into the FEMLAB solver. A triangular element grid was used to cover the two-dimensional reformer space. Appendix B specifies details

57 45 of the number of elements and number of degrees of freedom for each of the reformers modeled. The results of the FEMLAB finite element solution and the finite difference solution were similar and compatible. However, it was found that the finite difference solution resulted in lower reformate CO concentration. This was due to problems with the inherent lower magnitude of CO concentration in the reformate gas as compared to water, hydrogen and carbon dioxide concentrations. This difference created problems with convergence and minimum error search. The FEMLAB solver uses sophisticated minimum error search schemes which include parameter step annealing and magnitude parity. Because of this computational error reduction feature, the FEMLAB solution was assumed to have a higher degree of accuracy than that of the finite difference scheme. Steam Reformer Model Validation Validation of the steam reformer model results were done through comparison of the performance of modeled steam reformers with the UC-Davis reformer. Experiments were conducted using a small-scale laboratory steam reformer located at the University of California Davis (See Appendix C). Data from this experimentation was used for model validation. The UC-Davis reformer was built and instrumented by Dr. Paul Erickson and his students. Experiments for this study were conducted by UC-Davis students under the direction of the author of this dissertation. The experimental plan was also developed by this author for the purposes of this study. Various peer-reviewed publications have been based on research conducted using the UC-Davis reformer. These include Liao et al. [], Yoon et al. [3], and Betts et al. [4] (the latter regarding some of the results of this study).

58 46 Data obtained from the UC-Davis reformer cannot be used as a complete model validation tool. The reformer suffers from loose wall temperature control and gas sampling limitations that inevitably lead to discrepancies between the magnitudes of the model results and the data collected. Nonetheless, important physical phenomena, such as transient CO spikes, CO concentration dependence on reformer temperature, and steadystate and transient reformate composition trends are captured in the data obtained. Agreement between the model and the UC-Davis reformer were expected to be approximate. General Description The UC-Davis reformer is a catalyst filled cylinder with a radius of 0.47 inches (0.01 m), and an axial length of 1 inches (0.533 m) (Figure 4-9). The reactor temperatures are controlled using eight electrical band heaters located on the surface of the reactor. Thermocouple readings of the internal wall temperature of the reactor and the centerline temperature are used to gauge the level of chemical activity in a certain region of the reformer and are also used as feedback controller inputs for the heaters. The reformer heaters operate in on or off mode, as triggered by computer controlled relays. A simple feedback control loop, programmed in Labview, was used to maintain a constant wall temperature. Before entering the reformer water and methanol are vaporized and superheated using a series of pipes containing cartridge heaters. The temperature of the reactants entering the reformer is recorded and is controlled via a control and data acquisition computer, which is capable to turning on or off the vaporizer and super-heater heaters. Water and methanol fuel flows are metered using positive displacement gear pumps. The

59 47 gear pump RPM was used to measure flow. A scale, on top of which the reactants sat, was also used to measure methanol and water flow into the reformer. Reformate gas analysis was done using a NOVA Analytical Systems, Inc. gas analyzer (Model 7904CM). This analyzer was designed for simultaneous analysis of hydrogen, carbon dioxide, methane, and carbon monoxide. The instrument recorded carbon dioxide, carbon monoxide and methane using a Non-Dispersive Infrared (NDIR) detector with temperature and pressure compensation. Hydrogen was detected using a temperature controlled thermal conductivity (TC) cell. Before being analyzed for composition, the reformate gases were cooled to an almost freezing temperature (i.e. ~0 o C) and the condensed liquid was extracted. The UC- Davis reformer was designed with the capability to capture all condensed liquids before gas sampling to perform conversion studies. Due to the condenser volume, a short time response delay exists in the gas analysis data. Although the exact time response delay is a complex computation problem that falls outside the scope of this study, a general estimation can be performed. The length of the condenser tube is 18ft (5.5m). Reformate enters the condenser at approximately 300 o C and leaves at approximately 0 o C. If the mass flow rate of fuel entering the reformer is assumed to be the same as that entering the condenser, then the reformate volume flow entering the condenser can be calculated. For purposes of estimation the mass flow rate leaving the condenser can be said to be equal to the reformate mass entering it. Based on these assumptions the time it would take the time it would take to have a complete condenser volume gas change (i.e. retention time) for a premix flows at STP ranging from 5 ml/min to 30 ml/min would range between 1.3 to 0. seconds. Additionally, for all premix fuel flow rates, the flow through the

60 48 condenser can be estimated to be mostly laminar due to low Reynolds number. This means that a well organized concentration boundary layer is expected to exist within the condenser pipe. Therefore, although the reformate gas composition may change, traces of previous reformate gas could still be trapped in the boundary layer, close to the walls. This may lead to a decrease in the sharpness of the data obtainable from the gas analyzer. The NOVA gas analyzer manufacturer-quoted response time is 3 seconds to 90% response. The gas analyzer takes samples at a rate of 1 L/min through a positive displacement pump. At this flow rate the condenser unit would have a retention time in the order of 1 s. Since the response time is in the order of seconds and the measurable effects (that is reformate gas concentration changes) occur in the order of minutes, these effects are recordable using the UC-Davis hardware configuration. In terms of time response, passage through the vaporizers may be the limiting factor. For the model, the fuel flow rate boundary condition is at the inlet of the reformer. This is not the case for the UC-Davis reformer. If it were assumed that only liquid fuel passes through the vaporizers, then the residence times of the vaporizers would be as shown in Table 4-1. The estimated residence times presented in this table are maximum values. For lower fuel flows the space velocity is over-estimated because vaporization occurs earlier in the vaporizer tube, which increases the volume flow rate of the fuel. Also, the degree in which vaporization occurs within the vaporizers is not immediate and this will result in a ramped increase in fuel flow into the reformer instead of a step increase as in the model. This should result in smoother, drawn-out, transient gas composition data for the UC-Davis reformer.

61 49 Table 4-1. Estimated maximum times for vaporizer volume flow change. Fuel Flow Rate (cm 3 /min) Vaporizer residence times 5 5 min 10 3 min 15 min min 5 1 min 30 ~50 s Experimental Method The reformer catalyst bed was maintained idle for storage completely sealed and pressurized with carbon dioxide. This process ensured minimum catalyst deactivation due to oxidation. For experimentation, the reformer was first heated using the reformer band heaters under a CO gas purge. Once the appropriate reformer temperature was reached (50 o C), vaporized and superheated water and methanol were introduced. The experiment maintained the reformer wall temperature at a nominal 300 o C with fluctuations due to heater binary control. Water-methanol flows between 5 and 30ml/min were passed through the reformer in 5ml/min intervals. Changes in methanol-water flows were executed every 10 minutes. The data sampling rate was two seconds. UC-Davis Reformer Data The data obtained from the UC-Davis reformer shows the following trends: There is a temporary increase in the CO concentration in the reformate gas leaving the reformer when there is an increase in fuel flow rate into the reformer (Figure 4-10). The CO concentration in the reformate is very closely related to the catalyst bed temperature (Figure 4-11). Catalyst bed temperature fluctuations in the UC-Davis reformer were mostly because of band heaters binary operation and imposed control scheme. The steady-state CO concentration in the reformate lowers with increased fuel flow rate into the reformer (Figure 4-10 and Figure 4-1).

62 50 Figure 4-9. Experimental steam reformer developed at UC Davis 16 Data obtained at U.C. Davis on /5/004 3 Methanol - Water Mix Flow Rate Increased levels of CO are registered when there is an increase in flow rate CO Spike Due to Change in Flow Rate Run Time (min) Steady state CO levels are lowered with increased flow rate Pump feedback (ml/min) CO concentration (%) CO Concentration (%) Figure Example of CO spike in the reformate of the UC-Davis reformer after an increase in premix fuel flow (transient condition). Also shown is a decrease of the steady-state CO concentration in the reformate with increased premix fuel flow rate in units of ml/min.

63 Relationship between CO production and Catalyst Temperature Temperature ( o C) Magnitude is depicted in secondary axis CO Concentration (%) Run Time Reformer Exit Centerline Temperature CO concentration (%) Average Centerline Temperature (oc) Figure Reformate CO composition relationship to changes in temperature in the reformer catalyst bed 1 H Gas Concentration (%) Premix Flow Rate (std. ml/min) CO Spike CO Spike CO Gas Concentration (%) Run Time (min) Feb. 5th, 004 Test Run at UC Davis 0 Pump feedback (ml/min) H concentration (%) CO concentration (%) Figure 4-1. CO and H concentrations in the reformate gas in the UC-Davis reformer 1 The drastic decrease in CO concentration at the end of the graph comes as a result of shutting down the premix flow into the reformer.

64 5 It was a major goal of this investigation to be able to analytically reproduce the general operating characteristics of actual reformers. For this purpose an analytic version of the UC-Davis reformer was developed using the heat and chemical rate equations previously outlined. The virtual version of the UC-Davis reformer exhibited all of the major performance characteristics of the actual reformer to varying extents. The constants used for the virtual reformer are shown in Table 4-1 and were matched as best as possible, to measured or inferred UC-Davis reformer constants. Table 4-1. Values constants used to develop virtual UC-Davis reformer Name Value Used How it was determined Catalyst Bed Heat Capacity (C c ) 900 J/kg-K Inferred Catalyst Bed Thermal Conductivity (k eff ) 5.00 W/mK Estimated from Data Catalyst Bed Density (eρ c ) kg/m 3 Measured Length (L) m Measured Radius (R) 11.9 x 10-3 m Measured # of Radial Elements (N R ) 5 Number of Axial Elements (N L ) 150 Time Step ( t) 0.5 s * The catalyst heat capacity was assumed to be equal to an average of the heat capacities of copper and zinc. ** The catalyst heat capacity was assumed to be equal to an average of the heat capacities of copper and zinc. *** The void factor used was 50% and the catalyst density was measured via a water displacement method for a single pellet. To determine the value for catalyst bed thermal conductivity (k eff ), an average was taken of the measured thermal conductivity of the UC-Davis reformer catalyst bed at different points during the warm up stage, when no fuel flow exists. Figure 4-13 shows an example of the data used a band heater zone in the UC-Davis reformer. The resulting steady-state carbon monoxide composition of the virtual UC-Davis reformer are shown in Figure In arrows are listed the average steady-state CO

65 53 compositions for the actual UC-Davis reformer. In this figure, note that greater flow rate results in lower steady-state CO concentrations. Error between the UC-Davis reformer steady-state CO concentration and that of the virtual reformer is potentially due to the existence of un-reacted methanol in the sampled gas. Although before sampling the reformate gas was cooled to extract condensate, the difference between methanol and water freezing temperature only permitted a reduction in temperature to almost 0 o C. The existence of unreformed methanol in the sampled gas is detectable through the methane (or hydrocarbon) signature detected by the NOVA gas analyzer (Figure 4-15). Temperature (oc) Zone Thermal Conductivity Average Conductivity = 4.87 W/(K-m) Calculated Thermal Conductivity (W/(m-K)) Time (mins) 0 Zone Surface Zone Exit Wall Zone Exit Centerline Effective Conductivity Figure Estimation of catalyst bed thermal conductivity using UC-Davis reformer data

66 CO (mole fraction) E-0 1.7E E % UC-Davis 1.68E % UC-Davis 1.9E E E-0 1.5E E E E E Premix Flow (std.ml/min) x CO (dry) x CO (dry w/ MeOH) Figure Virtual UC-Davis reformer steady-state CO composition with varying premix fuel flow rates 1 10 Signal Output Run Time (min) Figure Hydrocarbon signal output obtained during UC-Davis reformer run.

67 55 Figure 4-16 shows a representative CO spike at the point in which the premix flow rate changes from 10 to 15 ml/min. The change in flow rate was produced at time zero. This figure is meant to be a parallel to Figure Figures 4-16 and 4-17 show the temperature distributions of the modeled reformer for 5 ml/min and 30 ml/min flow rate, respectively. The temperature distribution is similar to those recorded for the UC-Davis reformer (Figure 4-4), where the minimum temperature is recorded at the reformer entrance region. These results support the observations made from data obtained from the UC-Davis reformer and therefore demonstrate the validity of the assumptions made in the model presented Virtual Reformer CO Spike (10 to 15ml/min) CO Mole Fraction Time After Point of Flow Rate Change (s) Figure Virtual UC-Davis Reformer CO spike from 10 to 15ml/min premix flow rate change

68 56 Figure Virtual UC-Davis reformer temperature distribution (color plot) and methanol concentration (contour plot) for premix flow rate equal to 5 ml/min Figure Virtual UC-Davis reformer temperature distribution (color plot) and methanol concentration (contour plot) for premix flow rate equal to 30 ml/min

69 57 Reformer Design Currently, most reformer design is done based on a space velocity basis. This method is not consistent with the complex nature of the reforming reaction and the virtually infinite number of possible reformer designs. As part of this study, a set of nondimensional reformer performance variables are introduced in order to provide generality to the results obtained from analytical modeling and experiments. This section generally describes some parameters that affect reformer performance. In particular the question of how reformer length, catalyst loading and fuel flow rate affects the reformate gas composition is investigated using these suggested non-dimensional parameters. Design Generalization It has already been established that heat transfer and chemical rate characteristics are major determinants of complex reformer performance. Equation 4-5 can be rewritten as Equation 4-9. ρεc c T t = k eff r T r + k eff T r + k eff T x ερ E r c r M (4-9) In order to generalize the design of reformers, it is useful to define non-dimensional variables. For a reformer tube of a certain radius (R), the volume of catalyst in the reformer (V c ) will determine the length of the reformer (L). According to Nakagaki and other correlations it could be said that the mass of the reformer has a potential reformation capacity, Ω. The potential reformation capacity of the catalyst could be defined as the theoretical hydrogen power (P H ) that a certain mass of catalyst could produce if the reformer were maintained exactly at the wall temperature, and with the inlet methanol concentration. This can be easily calculated by using the Nakagaki

70 58 correlation. For example, the potential reformation capacity of the UC-Davis reformer can be calculated in the following way. The catalyst mass in the UC-Davis reformer is [ ] = 0. kg kg ( πr L) = 1983 ( m) ( m) mc = ερ cv = ερc π (4-30) 3 m If the wall temperature were to remain constant at 300 o C, then r M,max and the potential rate of methanol reacted ( Equation n& CH 3OH, reacted,max )can be calculated as shown in r M,max n& ( T, x ) w CH 3OH, reacted,max CH 3OH, inlet = r M = m c 4 = kmol sec kg 5 kmol sec (4-31) Since the theoretical chemical reaction for the reformation of methanol yields three moles of hydrogen per mole of methanol, the potential reformation capacity (Ω) can be obtained as shown in Equation 4-3. kj 5 kmol Ω = 3LHVH n& CH 3OH, reacted,max = = kW kmol sec (4-3) Also a non-dimensional temperature, θ, can be defined as T θ =, (4-33) T w and a non-dimensional length and radius can be defined as * r = * x = r R x L (4-34) (4-35)

71 59 Given these formulations, Equation 4-9 can be non-dimensionalized to yield VρcεCcT Ω w θ VTwkeff = t R Ω 1 * r θ VTwkeff + * r R Ω 4 θ πr T + * r VΩ w θ ερcerrm V * x Ω Equation 4-38 is made up of various important non-dimensional groups. A nondimensional conductivity, k *, can be expressed as (4-36) k * VT k w eff = R Ω The non-dimensional conductivity can be construed as a measure of the ratio (4-37) between the potential radial conduction heat transfer into the reformer and the potential rate of energy absorbed by the reforming reaction. It is important to point out that the non-dimensional conduction is not actually a function of catalyst length. The nondimensional conduction could be expressed as k * = T w k eff 3LHV ρ εr H c (4-38) For the UC-Davis reformer the non-dimensional conduction is equal to The last term in Equation 4-36, is a measure of the ratio of the actual rate of energy absorbed due to the reformation reaction and the potential rate of energy absorbed by the same reaction in terms of hydrogen production. The potential reformation capacity, Ω, is actually a measure of the potential hydrogen production of a certain mass of catalyst at a certain temperature. Thus it would be convenient to express the last term in Equation 4-36 in terms of hydrogen production and not methanol consumption. In practical terms the hydrogen production is easier to measure than the methanol converted.

72 60 With this formulation a catalyst effectiveness, λ, can be defined as the ratio of the actual hydrogen power output obtained from a certain amount of catalyst (P H ) and the potential reformation capacity of the catalyst (Ω) (Equation 4-39). λ = P H Ω (4-39) The catalyst effectiveness is a measure of the degree of reforming activity that a certain mass of catalyst undergoes. High catalyst effectiveness is desirable because it means that the reformer is operating closer to its idealized potential hydrogen yield. High catalyst effectiveness occurs in short reformers. On the other hand, the larger the reformer, the lower the catalyst effectiveness will be. From Equation 4-40, it is also notable the time constant term, τ, VρcεCcT τ = Ω w = C T c w 3LHVH rm,max (4-40) The transient response in a reformer would be strongly determined by τ. A low value of τ would lead to very fast temperature changes in the reformer tube leading to a decrease in prominence of the CO spikes. For the purpose of fully expressing the heat equation in non-dimensional terms a non-dimensional time (t * ) is defined (Equation 4-41). t * tω = Vρ εc T c c w (4-41) The non-dimensional heat equation is θ = * t k r * * θ + k * r * 4 θ πr T + * r VΩ w θ * x E r r M 3LHVH rm,max (4-4) It is also convenient to non-dimensionalize the fuel flow into the reformer as

73 61 flow ND = LHV fuel m fuel Ω (4-43) where LHV fuel represents the lower heating value of the fuel, and m fuel represents the mass flow rate of the fuel. The non-dimensional flow is a ratio between the rate of fuel energy introduced into the reactor and the reactors reformation capacity. The nondimensional flow can be expressed in terms of space velocity (Sv) as shown in Equation flow ND S ρ v fuel, in = (4-44) 3r M,max ερ c Also useful is the definition of reformer efficiency, η ref, η ref = PH P fuel (4-45) The chosen reformer efficiency definition neglects the power required to promote and sustain the reforming reaction. This definition only deals with issues of conversion, since a detailed analysis of reformer heating design is outside the scope of this study. In Equation 4-4, P H represents the rate of hydrogen chemical energy obtained from the reformer, and P fuel represents the rate of fuel energy introduced into the reformer. The maximum reformer efficiency for a methanol reformer is given by = 3LHV H η ref,max (4-46) LHVCH 3OH The non-dimensional flow can be re-written as flow ND 1 = η ref,max n& n& CH 3OH,, in CH 3OH,, reacted,max (4-47)

74 6 Given these definitions, the reformer model was used to obtain characteristic reformer efficiencies versus catalyst effectiveness curves for reformers with the same non-dimensional conduction (0.315) as the UC-Davis reformer at various fuel flow rates and different lengths. These curves are shown in Figure Additionally Figure 4-19 shows the corresponding non-dimensional fuel flow rates as a function of catalyst effectiveness. 7 Efficiency (%) Magnitude represented on secondary axis Catalyst Effectiveness Efficiency (5ml/min) Efficiency (10ml/min) Efficiency (15ml/min) Efficiency (0ml/min) Flow ND (5ml/min) Flow ND (10ml/min) Flow ND (15ml/min) Flow ND (0ml/min) Non-dimensional Flow Figure Reformer Efficiency with varying catalyst effectiveness at varying fuel flow rates From Figure 4-19 certain observations can be drawn. First, for non-dimensional fuel flow values less than 1, the variations in reformer efficiency versus catalyst effectiveness and non-dimensional fuel flow versus catalyst effectiveness, collapse into a single representative performance curve. Second, as mentioned previously, lowered catalyst effectiveness results in higher reformer efficiency. In other words, longer

75 63 reformers will exhibit higher steady-state efficiencies at the cost of ever increasing catalyst mass. Therefore, there are diminishing benefits to a long reformer, which will have a similar hydrogen yield as a reformer designed with catalyst effectiveness equal to 0.1. As the fuel flow rate of a reformer approaches infinity, the catalyst effectiveness will approach 1 and the reformer hydrogen yield will approach zero. These results can be used to scale a reformer using non-dimensional variables. For the purposes of this study, three reformer designs were considered: an infinite reformer with catalyst effectiveness equal to 0.05 at maximum power; a mid-size reformer with catalyst effectiveness equal to 0.1 at maximum power; and a small reformer with catalyst effectiveness equal to 0. at maximum power. The terms small, medium and large are used as visual clues to the general aspect ratio that these reformers would have in comparison with each other if the diameter were to remain the same. All reformers were designed to produce a maximum 60 kw of hydrogen (lower heating value based). If a reformer tube with the same non-dimensional conduction as the UC-Davis reformer were to be scaled to produce 60 kw of power, its length and steady-state operating characteristics would be as shown in Table 4-, based on steady-state catalyst effectiveness and reformer efficiency analysis. Table 4-. Steady-state design of 60 kw reformer with an 11.9 mm radius and a catalyst bed density of 1983 kg/m 3 Catalyst Effectiveness (λ) Reformer Tube Length (m) Steady Sate Reformer Efficiency (%) Premix Power to Produce 60 kw of H (kw) D. Davieau [16] studied the use of space velocity and aspect ratio as major similitude parameters for reformer design. His conclusions are that these two parameters

76 64 are not sufficient to fully characterize reformer steady-state conversion and reformate quality. These conclusions are supported by the analysis presented in this study. The space velocity and aspect ratio, although important, do not capture the governing physics of the reformation process.

77 CHAPTER 5 RESULTS In this chapter, the primary chemical and heat transfer conditions that produce variations in hydrogen and carbon monoxide levels in the reformate are presented and discussed. Special attention is given towards the benefits and detriments associated with reformer size and catalyst loading in the context of fuel cell system incorporation. For this purpose three general reformer length with the same non-dimensional conduction were compared based on their reformate quality. Case 1 - Infinite Reformer Transient Efficiency Figure 5-1 shows predicted transient changes in reformer efficiency for various changes in non-dimensional fuel flow rates for the infinite reformer. This figure shows how increasing fuel flow results in an immediate drop in reformer efficiency followed by a gradual approach to the steady-state efficiency. Figure 5- shows the same effects on the reformer efficiency in terms of the efficiency difference variable, ~ η ref, defined as ~ η ref = η η (5-1) ref ref, S. P. The variable η ref,s.p. is the set point efficiency, which is the steady-state efficiency at a certain flow rate. In other words, for any steady flow rate the difference efficiency is equal to zero. The magnitude of the initial efficiency overshoot is summarized in Table 5-1. It is notable that the magnitude of the efficiency-overshoot increases with increased change in 65

78 66 non-dimensional flow. However, if the change in non-dimensional flow remains constant, changes between smaller non-dimensional flow quantities results in less overshoot. The magnitude of the drop in efficiency before reaching steady-state is also recorded. This is called an efficiency undershoot. High overshoot magnitudes reduced the extent of the efficiency undershoot. Table 5-1. Case 1: Summary of the transient response of the infinite reformer to changes in fuel flow rate Flow ND Change in ND Flow Overshoot Initial Final % % % % % % Infinite Reformer Transient Efficiency 13.1 Efficiency (%) Run Time (s) ND Flow Key: 10kW: 1.10% 0kW:.0% 30kW: 3.31% 40kW: 4.4% 50kW: 6.63% to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW 10 to 60kW Figure 5-1. Case 1: Infinite reformer operation when undergoing step variations in fuel flow rate

79 Infinite Reformer Transient Efficiency Run Time (s) ND Flow Key: 10kW: 1.10% 0kW:.0% 30kW: 3.31% 40kW: 4.4% 50kW: 6.63% 10 to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW 10 to 60kW Figure 5-. Case 1: Reformer efficiency changes due to change of fuel flow rate for the infinite reformer Case - The Medium Sized Reformer As with the infinite reformer, the medium sized reformer was subjected to transient variations in fuel flow. The changes in efficiency of the reformer are depicted in Figure 5-3 and 5-4. Figure 5-4 shows these changes in terms of the efficiency difference variable. Table 5- summarizes these results. As with the infinite reformer, large changes in non-dimensional flow result in large overshoots, and changes between lower nondimensional flows results in lower overshoots. The magnitude of the efficiency undershoots for the medium reformer were significantly lower than those of the infinite reformer. In addition, as changes in nondimensional flow rates occurred between higher non-dimensional flow rates, the efficiency undershoot was decreased and even disappeared.

80 68 Table 5-. Case : Summary of the transient response of the medium size reformer to changes in fuel flow rate Flow ND Change in ND Flow Overshoot Initial Final % % % % % % Efficiency (%) Medium Reformer Transient Efficiency ND Flow Key: 10kW:.% 0kW: 4.43% 30kW: 6.65% 40kW: 8.87% 50kW: 11.08% 60kW: 13.30% Run Time (s) 10 to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW 10 to 60kW 100 Figure 5-3. Case : Transient efficiency of the medium size reformer under transient load changes Case 3 - The Short Reformer As with the infinite and medium sized reformer, the short reformer was run under transient load conditions as shown in Figure 5-5 and 5-6. In contrast with the larger reformers, the short reformer did not exhibit efficiency undershooting, except between step changes in non-dimensional flow between 4.44% and 8.88%. Besides this effect, the

81 69 results follow a similar pattern to those obtained for the infinite and medium reformers. The results of the transient run of the short reformer are given in Table Medium Reformer Transient Efficiency ND Flow Key: 10kW:.% 0kW: 4.43% 30kW: 6.65% 40kW: 8.87% 50kW: 11.08% 60kW: 13.30% Run Time (s) 10 to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW 10 to 60kW 100 Figure 5-4. Case : Reformer efficiency changes due to change of fuel flow rate for the medium size reformer Table 5-3. Case 3: Summary of the transient efficiency response of the short reformer to changes in fuel flow rate Flow ND Change in ND Flow Overshoot Initial Final % % % % % %

82 70 Reformer Efficiency (%) Short Reformer Transient Efficiency Run Time (s) ND Flow Key: 10kW: 4.94% 0kW: 8.88% 30kW: 13.31% 40kW: 17.75% 50kW:.19% 60kW: 6.63% 10 to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW 10 to 60kW 100 Figure 5-5. Case 3: Transient efficiency of the short reformer under transient load changes Carbon Monoxide Concentration Generally, temporary increases in CO concentration in the reformate gas occurred during periods of increased premix flow rates. This phenomenon occurs because lower fuel flow into the reformer leads to higher average steady-state temperature. This increase in temperature results from less fuel being reformed, thus lower overall associated endothermicity. High temperatures promote greater CO and water formation in the water gas shift reaction. High temperatures typically occur near the exit of the reactor. When there is an increase in fuel flow into the reformer, the average temperature of the reactor does not drop immediately and this contributes to better transient conversion (i.e. increased efficiency), and higher transient CO concentrations. As time progresses the average temperature of the reactor drops because of the increased endothermicity

83 71 associated with the increased level of methanol reacted. Eventually the lower average reformer temperature results in a lower steady-state CO level at higher fuel flow rates. An off-steady-state CO concentration is defined as [ CO] Off steady = [ CO] [ CO] steady (5-) Short Reformer Transient Efficiency ND Flow Key: 10kW: 4.94% 0kW: 8.88% 30kW: 13.31% 40kW: 17.75% 50kW:.19% 60kW: 6.63% Run Time (s) to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW Figure 5-6. Case 3: Reformer efficiency changes due to change of fuel flow rate for the short reformer Figures 5-7 through 5-13 show the transient CO concentrations and off-steady-state CO concentrations for the infinite, medium, and short reformer triggered by step changes in fuel flow into the reformer. The CO spike is defined as the initial increase in CO concentration that occurs after a step increase in fuel flow occurs. This spike occurs before there is a gradually decreasing approach of the transient CO concentration towards the steady-state CO concentration. In all instances, the larger the non-dimensional flow change the greater the

84 7 CO spike. However, between equivalent changes in non-dimensional flow, larger CO spikes occurred between lower non-dimensional flows. For example, in the infinite reformer, a CO spike 38 ppm larger is recorded for a change in non-dimensional flow from 1.10% to.1% as compared to a change from.1% to 3.31%. CO spikes, as defined here, are a measure of the transient deviation of the CO concentration as compared to the steady-state condition for the same flow. This deviation was highest for small reformers, which in order to obtain a certain steady-state hydrogen flow rate required a large change in non-dimensional flow. However, if Figures 5-8, 5-10 and 5-1 are examined it can be seen that the off-steady-state CO concentration also includes the unsteady, smooth drop in CO concentration associated with the increased fuel flow into the reformer. Changes in flow between low non-dimensional flows caused greater disruptions to this smooth drop. In a fuel cell system these flow changes may cause variations in fuel cell performance and a degree of CO poisoning during the transient event if appropriate actions are not taken. From this standpoint, fuel flow variations in reformers operating with high non-dimensional flows (i.e. small reformers) should exhibit greater linearity in power output during load changes. The short reformer, the change of non-dimensional flow from.19% to 6.63% exhibited a very mild CO drop disruption. In Figure 5-13, the CO concentrations associated with a similar change in reformate hydrogen power output, namely from 40 kw to 50 kw (lower heating value based), is depicted. For this case, the short reformer produced a lower steady-state and transient CO concentration change than the larger reformers. This occurs because the catalyst

85 73 effectiveness of the small reformer is much higher than for the larger reformers, which leads to lower average catalyst bed temperatures. CO (wet fraction) Infinite Reformer Transient CO Concentration ND Flow Key: 10kW: 1.10% 0kW:.0% 30kW: 3.31% 40kW: 4.4% 50kW: 6.63% Run Time (s) to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW 10 to 60kW Figure 5-7. Infinite reformer transient CO concentrations for various for changes in power output CO Concetration (ppm) Infinite Reformer CO Spikes ND Flow Key: 10kW: 1.10% 0kW:.1% 30kW: 3.31% 40kW: 4.4% 50kW: 5.5% 60kW: 6.63% to 60kW 40 to 50kW Time (s) 30 to 40kW 0 to 30kW 10 to 0kW 10 to 60kW Figure 5-8. Infinite reformer transient off-steady-state CO concentrations for various changes in reformer power output

86 74 CO Concentration (wet fraction) Medium Reformer Reformate CO Concentration ND Flow Key: 10kW:.% 0kW: 4.43% 30kW: 6.65% 40kW: 8.87% 50kW: 11.08% 60kW: 13.30% Time (s) 10 to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW 10 to 60kW Figure 5-9. Medium reformer transient CO concentrations for various changes in power output CO Concentration (ppm) Medium Reformer CO Spikes ND Flow Key: 10kW:.% 0kW: 4.43% 30kW: 6.65% 40kW: 8.87% 50kW: 11.08% 60kW: 13.30% Time (s) 10 to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW 10 to 60kW Figure Medium reformer off-steady-state CO concentrations for various changes in reformer power output

87 75 CO Concentration (wet fraction) Short Reformer Transient CO Concentration ND Flow Key: 10kW: 4.94% 0kW: 8.88% 30kW: 13.31% 40kW: 17.75% 50kW:.19% 60kW: 6.63% Time (s) 10 to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 50 to 60kW Figure Short reformer transient CO concentrations for various changes in power output CO Concentration (ppm) Short Reformer CO Spikes ND Flow Key: 10kW: 4.94% 0kW: 8.88% 30kW: 13.31% 40kW: 17.75% 50kW:.19% 60kW: 6.63% Time (s) 10 to 0kW 0 to 30kW 30 to 40kW 40 to 50kW 10 to 60kW Figure 5-1. Short reformer off-steady-state CO concentration for various changes in reformer power output

88 Comparison of CO Concentrations for Different Size Reformers Operated Between 40 to 50kW Steady State Power Outputs CO Concentration (ppm) ND Flow ranges: Short Reformer: 17.75% to.19% Medium Reformer: 8.87% to 11.08% Infinite Reformer: 1.10% to 6.63% Run Time (s) Medium CO Short CO Infinite CO Figure Transient CO comparison of various size reformers operating between 40 and 50 kw hydrogen power output Hydrogen Concentration Transient and steady-state changes in hydrogen concentration for all reformers were such that the performance of PEMFCs or PAFCs should not be drastically affected. Even under the harshest transient conditions studied, in which a change in fuel flow was induced to generate a steady-state hydrogen yield between 10 kw and 60 kw, the reformers hydrogen yields did not deviate much more than 5% (Figure 5-14). The lowest hydrogen concentration was slightly below 60% wet molar concentration, and corresponded to the small reformer producing 60 kw hydrogen yield.

89 H Concentration ND Flow Changes: Infinite: 1.10 to 6.63 Medium:. to Short: 4.44 to Run Time (s) Infinite (10kW to 60kW) Medium (10kW to 60kW) Short (10kW to 60kW) Figure Comparison of transient changes in H concentration induced by step changes in fuel flow corresponding to hydrogen yields from 10 to 60 kw (lower heating value based) for various size reformers

90 CHAPTER 6 CONCLUSIONS Based on the experimental and modeling results presented in this study the following conclusions may be expressed. Steam Reformer Modeling A method of modeling a methanol steam reformer was proposed, demonstrated and compared to actual reformer performance The reformer model results matched general reformer behavior observed in actual reformers Based on this model formulation a new set of reformer similarity variables were proposed and the use of space velocity and aspect ratio as reformer design variables was demonstrated to be insufficient Steady-state Reformer Operation Non-dimensional flow, catalyst effectiveness, and non-dimensional conduction were introduced as potential reformer similarity variables. These variables effectively characterized the steady-state reformer performance of reformers. Based on these variables, the results showed that the steady-state CO concentration decreases with increased non-dimensional flows The steady-state H concentration decreased with increased non-dimensional flows Higher non-dimensional flows resulted in decreased steady-state reformer conversion Unsteady Reformer Operation Transient changes in fuel flow into steam reformers produced non-linear variations in hydrogen and CO concentration in the reformer outlet These variations are mostly caused by transient changes in reformer catalyst bed temperature 78

91 79 Step increases in non-dimensional fuel flow produced a transient increase in CO concentration followed by an eventual drop in the CO concentration This transient increase in CO concentration was greater when large changes in nondimensional flow occurred Step increases between large non-dimensional flows produced lower transient increases in CO concentration than the same step changes between lower nondimensional flows. The implications of the results and methodology presented in this study are that steam reformer chemical behavior, in the transient and steady-state case, has been expressed in mathematical, non-dimensional, terms. This leads to the proposal of similitude expressions that can be used to better characterize, compare, understand and design packed bed steam reformers. Also, this study demonstrates the mechanisms that produce carbon monoxide spikes during period of flow changes. Generally, from a design standpoint, there seems to exist a balance between the design of an infinitely long reformer that would have excellent efficiency (conversion) but would also have high outlet CO concentrations in the steady state and transient regime, and a short and stubby reformer with low efficiency but low transient and stead sate CO concentrations for the same fuel flow rate. The optimum point will depend on the fuel cell and its tolerance of CO. Design should be made such that during the highest changes in non-dimensional flow, the CO spike does not exceed a maximum concentration that would drastically reduce fuel cell performance through poisoning.

92 CHAPTER 7 SUGGESTED FUTURE WORK The analysis provided in this study is largely based on experimental results obtained from a single diameter reformer (UC Davis reformer). The numerical models developed generally matched the performance of this reformer. A greater body of experimental data, with reformers of varying lengths and diameters should be obtained to determine the practical limitations of the analysis presented here. In addition, a set of non-dimensional performance variables has been introduced to the field of reformer design and analysis. These variables worked extremely well at describing reformer performance based on the models developed. A study should be conducted to evaluate whether these variables (or a variation of these variables) are capable of determining similarity between distinct reformers. Expansion of this study into reformer controls design, fuel cell controls design, and overall fuel cell system controls design is a logical expansion of this body of work. Understanding the transient nature of fuel cell systems is critical to their effective control. It is hoped that this work forms the foundation for the development of controllers that are capable of producing fast transient load changes in future fuel cell systems with minimal efficiency penalties. 80

93 APPENDIX A COPY OF FEMLAB SOLUTION REPORT 81

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111 APPENDIX B PRELIMINARY LOAD FORECASTING STUDY The lose of efficiency due to the sluggish response of the fuel cell system when operating in a transient load environment may be mitigated by the capability to preempt the onset of large changes in load. For a vehicle this may be achieved by learning the driving style of the vehicle driver such that periods of increased acceleration may be anticipated. Figure B-1 shows a typical load profile for a drive cycle. This load profile was obtained by driving TBB- in actual traffic conditions Power (kw) Transient Load Profile for TBB- under the UFFL driving cycle Run Time (s) Figure B-1. Transient load profile for TBB- under the University of Florida Fuel Cell Lab driving cycle data obtained on

112 100 In order to test this hypothesis TBB- was driven under a standard route (UFFL cycle) and the transient load, L(t), was obtained. Given this load a power change index, PCI(t), was defined (Equation B-1). L( t) PCI ( t) = L( t) + L( t ) (B-1) t o In Equation B-1, t o is the sample interval in time. Thus L(t-t o ) represents the previous load data. For the data obtained from TBB-, t o equaled 1 second. Notice that PCI is equal to 0.5 if there is no load change and that its value is bound between 1 and PCI PCI for TBB- under UFFL driving cycle Run Time (s) Figure B-. Power Change Index (PCI) for TBB- under UFFL driving cycle data obtained on By looking at Figure B- it is evident that most of the time PCI oscillates slightly from 0.5, meaning that relatively small variations in load are occurring. Furthermore it is shown that large fluctuations in the value of PCI occur in a recognizable pattern. That is,

113 101 every large decrease in the value of PCI (from breaking or stopping), there is almost always a consequent large increase in the PCI value. Input Layer Hidden Layer Output Layer x (t-0) x (t-19) x (t-18) x (t-17) f(net) f(net) f(net) f(net) y 1 y f(net) f(net) y 3 x (t-0) f(net) f(net) y N Figure B-3. Multi-layer Perceptron. Out of ten runs PCI data from one of the runs was used to train a multilayer perceptron (MLP) whose input were 0 PCI data points. The MLP used had 10 output processing elements, whose output should approximate PCI(t+t o ), PCI(t+t o ), PCI(t+3t o ),, PCI(t+10t o ). Figure B-3 shows a diagram of the MLP used. The output of each processing element (PE) is the hyperbolic tangent of the sum of the multiplication of the inputs to a gain or weight (Equation B-). See Figure B-3. PE j = tanh xiwi, j (B-) i Therefore, the output of an output layer processing element can be written as shown below, where k is the output index, i is the input data index, and j is the hidden processing element index (Equation B-3).

114 10 PE output, k = tanh tanh xiwi, j w j, k (B-3) k i Training of the MLP was done using the error back propagation algorithm. An error was defined as the difference between the MLP output and the desired MLP output (Equation B-4). For the purposes of these experiments, the desired output was the future data, x(t+kt o ) that the k th output of the P.E.should approximate. E k, i x( t + kto ) PEoutput, k = (B-4) 5). From this error definition a mean squared error (J) can also be defined (Equation B- J = 1 N N M E k, i i= 1 k = 1 (B-5) In Equation B-5 N represents the number of data points introduced into the MLP (for our purposes N=0). M represents the number of output processing elements in the MLP (in our case 10). The mean squared error is used as a cost function whose minimum is found with respect to the weights using a gradient descent algorithm. The algorithm chosen was the error back propagation algorithm. The trained MLP was tested with new data obtained by driving TBB- under the same UFFL route but on a different day and under different traffic conditions. Figure B-4 shows the performance of the MLP as a load predictor. Generally, the trained MLP was able to predict large increases in load.

115 PCI MLP Performance Test Run Time Actual Run Prediction (t+10s) Figure B-4. The MLP performance as a 10 second PCI predictor.

116 APPENDIX C PHYSICAL DESCRIPTION OF THE UC-DAVIS REFORMER The following is a physical description of the UC-Davis reformer test rig. Part of the information presented in this Appendix was provided and prepared by UC-Davis students and Dr. Paul Erickson, director of the UC-Davis Hydrogen Production and Utilization Laboratory (HYPAUL). The UC-Davis methanol-steam reformer used for the experimental portion of this study can be divided into subassemblies as shown in Table C-1. Figure C1 shows a general process diagram of the UC-Davis reformer. Table C-1. UC-Davis reformer subassemblies Methanol Steam Reformer Elements Subassembly Premix Reservoir, Pump, Scale Pumping 3-Stage Vaporizer, Super-heater Vaporizer Steam Reforming Reactor Catalyst Bed Housing Condenser, Condensate Trap Condensing Unit General Description A premixed concentration of methanol in water is mounted on a scale. This premixed fuel is pumped through three stage vaporizers which ensure only vapors are introduced into the reformer tube at the user specified temperature. In order to ensure this a superheater is also inline. The superheater ultimately controls the fuel temperature before introduction into the reformer catalyst bed. 104

117 105 Figure C-1. General process diagram of the UC-Davis reformer After passing through the catalyst bed, the gaseous species were directed into two tubes via a system of valves. One route was for analysis and the other was for exhausting the reaction products, reformate. Both routes lead to the condensing unit where the reformate temperature is reduced in order to separate liquid water and unreacted methanol from the gas mixture. A condensate trap is used for this purpose. The generally dry reformate gas is then routed to the gas analyzer. The system also includes the capability to drain the catalyst bed through a series of drain valves and allow for carbon monoxide purging and preservation of the catalyst bed. An air purge was adapted after the catalyst section to ensure retrieval of all condensed species in the condensate trap.

118 106 Pumping Subassembly The pumping subassembly (Figure C-) is initially composed of a 4 liter (1 gal.) polyethylene carboy reservoir containing a user prepared molar concentration of methanol in water. For the experiments conducted a methanol concentration of 1.5 moles of water per mole of methanol was used. The tank was allowed to rest at room temperature. Appropriate premix concentration was verified through the use of hand-held specific gravity sensor (Anton Parr - model 35n with a resolution of g/cm3). The premix reservoir rested upon a 4100 gram-ohaus scale with a 0.1 gram resolution. The scale had a 9 pin bidirectional RS-3 port, which allowed the user to electronically record the scale reading during operation. The premix was then drawn out of the reservoir by way of a gear pump through a cm (0.5 in) I.D. vinyl tube. The gear pump head (model EW ) and driver (model A ) was manufactured by MicroPump and enabled the user a resolution of 0.1 ml/min with a premix flow rate range from.6 to 85 ml/min. The pump driver was equipped with a frequency output signal, which allowed the user to correspond a frequency (or gear pump RPM) to a flow rate. The user can then electronically record the instantaneous flow rate and control the pump with a voltage signal during operation. Calculations for the mass flow rate could be verified by both the recorded pump flow rate and by recording the change in mass of the scale divided by the time the experiment ran (both were recorded via the computer control program).

119 107 Figure C-. Premix reservoir and gear pump Figure C-3. Vaporizer design The vaporizer subassembly, composed of a 3-stage vaporizer and a superheater, is shown in Figure C-3 and Figure C-4, respectively. The premix exited the gear pump through a cm (0.5 in) O.D. stainless-steel tube (0.14 cm (0.049in) wall thickness) and then entered the first vaporizer of the 3-stage vaporizer. All 3 vaporizers were built with identical dimensions. Each vaporizer was made of a 0.3 cm (8 in) stainless-steel pipe (nominal ½ Dia., schedule 40) threaded at both ends. Caps for the pipes were machined to adapt a cm (0.5 in) tube on the top end and a cm (0.5 in)

120 108 cartridge heater on the bottom end. The energy for vaporization was supplied from 10V cartridge heaters. The first vaporizer contained a 4.1 cm (9.5 in), 55W cartridge heater, while the last two stages contained 1.7 cm (5 in), 400W cartridge heaters. Each vaporizer was monitored for temperature by two, stainless-steel-sheathed, ungrounded K- type thermocouples. The temperatures of the cartridge heaters were monitored by a 0.05 cm (0.010 in) diameter-thermocouples. These thermocouples were located external to the flow, attached to the base of the cartridge heater (Figure C-3). The exit temperature of each vaporizer was monitored by a cm (0.065 in) thermocouple. These thermocouples were located internal to the flow, at the joint union between each stage (Figure C-3). At the exit of the second vaporizer a kpa (0-15psi) pressure gauge (identified as PT in Figure C-3) was installed to monitor the pressure near the beginning stage of the reformer. Several stainless-steel, high temperature quarter-turn valves were installed throughout the vaporizer subassembly. At the inlet to the first vaporizer, a valve was adapted to drain the condensate in the subassembly when necessary and at the exit of the first vaporizer a valve permitted an inlet for catalyst reduction gas. When the vaporized premix was not up to temperature, an exhaust valve at the exit of the 3-stage vaporizer prevented the flow from entering the superheater and catalyst bed housing, as shown in Figure C-4. A CO purge and check-valve were located at this position as well, to purge vaporized premix out of the reactor and lock oxygen out when not in operation. The superheater housing material was a 30.5 cm (1 in) stainless-steel pipe (nominal ¾ Dia., schedule 40) threaded at both ends. Caps for the pipes were machined to adapt to cm (0.5 in) tubes and could be removed. To prevent condensation build-up on the bottom of the superheater, which would funnel the liquid premix directly

121 109 into the reactor (thus damaging the catalyst); a recessed tube was inserted approximately 7.6 cm (3 in) inside the superheater. All pipes were sealed on the vaporizer subassembly by either welding or a sealant (Resbond 907GF), which was capable of withstanding temperatures up to 188 o C (350 o F). All tubes were sealed using Swagelok compression fittings. External heating was applied to the superheater using four nozzle band heaters (.5cm (1in) I.D., 5.1 cm (in) width), each with a 10V, 75W rating. To evenly increase the temperature distribution throughout the superheater, a highly thermal conductive aluminum tape was wrapped around the exterior. Three cm (0.065 in), stainless-steel-sheathed, ungrounded K-type thermocouples were strategically adapted to the superheater to monitor performance. Two thermocouples were integrated inside the superheater to monitor superheater internal and exit (just before the reactor) temperature. A third was placed externally on a nozzle band heater, permitting the user to monitor the temperature of the heat source. Insulation for both the 3-stage vaporizer and superheater composed of an alumina-silica material capable of withstanding temperatures up to160 o C (300 o F). Because the superheater utilized external heating, it required extra insulation composed of inorganic-volcanic rock fiber capable of withstanding temperatures up to 649 o C (100 o F). Catalyst Bed Housing Subassemblies The housing material for the reactor was a 61 cm (4 in) stainless-steel pipe (nominal ¾ in Dia., schedule 40) threaded at both ends, as shown in Figure C5. The upper cap was machined to adapt a cm (0.5 in) tube, as well as a cm (0.5 in) MNPT pipe on the side. The MNPT pipe created a pathway to a Kistler pressure transducer. The cap at the exit of the reactor was a specially machined reducing T fitting. One end of the T attaches to a cm (0.5 in) tube, while the other attaches

122 110 to the 1.9 cm (0.75in) pipe. External heating was applied to reactor A using eight nozzle band heaters (.5cm (1in) I.D., 5.1cm (in) width), each with a 10V, 75W rating. In a similar fashion as the superheater, a highly thermal conductive aluminum tape was wrapped around the exterior of the pipe to evenly increase the temperature distribution throughout the reactor. An array of seventeen 0.159cm (0.065in) Dia. stainless-steelsheathed, ungrounded K-type thermocouples was used to monitor the temperature within the reactor. The thermocouples were attached to the reactor housing using 0.3cm (0.15in) MNPT to 0.3cm (0.15in) pipe fittings by Swagelok. All fittings were sealed using a chemical resistant-teflon tape or by welding. To observe the temperature of the heat bands, eight 0.05 cm (0.010 in) Dia., ungrounded K-type thermocouples were placed between the heat bands and the exterior reactor wall. The reactor pressure was monitored using a kPa (0-15psi) pressure gauge and was located at the exit of the reactor (Figure C5). Insulation for the reactor was composed of a 3 thick calcium silicate material with a temperature tolerance of 649 o C (100 o F). In both reactors, a stainless-steel mesh (56 squares per inch, 0.015in wire diameter) was placed at the bottom of the reactor to provide a base for the palletized catalyst. The catalyst used in this study was a pelletized commercial-grade copper-zinc catalyst on an alumina substrate. This catalyst (FCRM-) was manufactured by Sud- Chemie and is recommended for an operating temperature range of 50-80oC ( o F). Stored in an oxidized state, the catalyst needed to be reduced before steam reformation could occur. Reduction was done by flowing diluted concentrations of hydrogen in nitrogen for a period of 10 hours. Successful reduction was accomplished

123 111 once temperature variations within the reactor were negligible. Information including the mass of catalyst loaded and the length of filled catalyst bed housing was recorded. The catalyst was cylindrical in shape and had dimensions consisting of 0.47 cm (0.187 in)- diameter and 0.5 cm (0.100 in)-thickness, as stated by the manufacturer. Schedule 40 SS Pipe 0.84 ID, ¾ Nominal 1 Length, Threaded Ends Figure C-4. Superheater design Condensing Unit Subassembly The condenser housing consisted of a modified 113liter (4.0 ft 3 ) refrigerator manufactured by Haier. The condenser was designed with two intakes, one for reactor exhaust and the other for analysis. To increase conduction and gas path length, the reacted species or exhaust entered into 0.635cm (0.5 in)o.d. copper tube coiled inside 10.cm (4in) PVC canisters ( canisters on the analysis side).

124 11 Figure C-5. The UC-Davis reactor design Here, water from an ice bath doused the coils via a submersible pump; lowering the gas temperature from 50 o C (48 o F) to 0 o C (3 o F). The decrease in temperature promotes a phase change, causing water, methanol and other relevant species to condense. The condensate was then trapped in the container. The condensate from the exhaust side was removed from the trap via a drain located on the outside of the condenser. To acquire the condensate from the reactant species for analysis, the container was removed from the unit. The dry product gas on the analysis side was then routed to the gas analyzer, while the exhaust gas was directed to the fume hood. Two 0.159cm