Energy-efficient method for thermal processing of Utah tar sands

A modification of the existing University of Utah process for thermal recovery of oil from tar sands is described and tested. In the first step of the process, bituminous sand is fed into a fluidized pyrolysis bed, where the bitumen is cracked and vaporized at temperatures of 430°C to 550°C. In the second step, energy required to operate the fluidized bed is generated by burning coked sand, a by-product of the pyrolysis reactor, with air in a second fluidized bed reactor located directly below the pyrolysis reactor. The coked sand is transferred from the pyrolysis to the combustion reactor by gravity, and energy is transferred from the combustion reactor to the pyrolysis reactor by using a potassium-containing heat pipe as the heat exchanger device. In the third and final step, the synthetic oil produced and some of the energy from the exiting streams are recovered. Main modifications of the existing thermal process include fluid dynamic decoupling of the fluidized bed reactors by

AN ENERGY-EFFICIENT METHOD FOR THERMAL PROCESSING OF UTAH TAR SANDS by Raschid Jose Bezama A dissertation submitted to the faculty of The University of Utah partial fulfillment of the requirements for the degree of Doctor of Philosophy Department of Chemical Engineering The University of Utah June 1983 © 1983 Raschid Jose Bezama All Rights Reserved THE UNIVERSITY OF UTAH GRADUATE SCHOOL SUPERVISORY COMMITTEE APPROVAL of a dissertation submitted by Raschid J. Bezama This dissertation has been read by each member of the following supervisory committee and by majority vote has been found to be satisfactory. I I Chainna~ - D. S a er 7 ~. A. G. Db 1 ad THE UNIVERSITY OF UTAH GRADUATE SCHOOL FINAL READING APPROVAL To the Graduate Council of The University of Utah: I have read the dissertation of Rasch; d J. Bezama 'n Its final form and have found that (1) its format, citations, and bibliographic style are consistent and acceptable; (2) its illustrative materials including figures, tables, and charts are in place; and (3) the final manuscript is satisfactory to the Supervis­ory Committee and is ready for submission to the Graduate School. Date Chairperson, Supervisory Committee Approved for the Major Department A. La on yler Chairman I Dean Approved for the Graduate Council James L. Clayton I Dean of The Graduate School ABSTRACT A modification of the existing University of Utah process for thermal recovery of oil from tar sands is described and tested. In the first step of the process, bituminous sand is fed into a fluidized pyrolysis bed, where the bitumen is cracked and vaporized at temperatures of 430°C to 550°C. In the second step, energy required to operate the fluidized bed is generated by burning coked sand, a by-product of the pyrolysis reactor, with air in a second fluidized bed reactor located directly below the pyrolysis reactor. The coked sand is transferred from the pyrolysis to the combustion reactor by gravity, and energy is transferred from the combustion reactor to the pyroly­sis reactor by using a potassium-containing ~eat pipe as the heat exchanger device. In the third and final step, the synthetic oil produced and some of the energy from the exiting streams are recovered. Main modifications of the existing thermal process include fluid dynamic decoupling of the fluidized bed reactors by using an appropriate oxygen-free gas to fluid­ize the pyrolysis reactor, and rerouting of the different streams in a pattern that maximizes the global thermal efficiency of the process. In the modified process, the heat pipe network is designed to be able to transport all the energy required at the pyrolysis reactor from the combustion reactor. The modified process is studied from thermodynamic and dynamic points of view. The thermodynamic study indicates the feasibility of the design and the existence of an optimal design and operating conditions for the given process. The thermodynamic impact of using heat pipes as a method for transporting energy from the combus­tion reactor to the pyrolysis reactor is compared to other existing methods such as recycling hot sand and preheating the pyrolysis reactor fluidizing gas. The dynamic behav­ior of the proposed process was studied experimentally by running a laboratory-scale unit and analytically with the help of mathematical models specially developed for the simulation o£.the proposed process. The results of the simulation indicate that the heat transport between both fluidized reactors is limited, at low temperatures, by the sonic limit of the existing potassium heat pipe, and at high temperatures, close to the. design operating point, by the convective heat transfer inside the fluidized-bed reactors. v CONTENTS ABSTRACT . • • . • LIST OF FIGURES LIST OF T'ABLES NOTATION • • • ACKNOWLEDGMENTS Chapter I. INTRODUCTION II. OBJECTIVES III. BACKGROUND AND PROCESS DEVELOPMENT IN THERMAL PROCESSING OF TAR SANDS IV. Literature Review ...••...• Previous Developments of the University of Utah Thermal Process . • . • . • New Developments of the University of Utah Thermal Process • . . . . . . • THERMODYNAMIC ANALYSIS OF THERMAL RECOVERY . • . The Basic Equations • . • . • . • • . The Pyrolysis Reactor . •. .• . The Combustion Reactor • . . . • . Transferring Energy from the Combustion Reactor to the Pyrolysis Reactor V. DEVELOPMENT OF A DYNAMIC MODEL The Conservation Equations for a Heat Exchanger Stage . • . . . The Conservation Equations for the Reactor Stage . . • . . • . . . Five-Stage Dynamic Model .... The Heat Pipe Model • . • . . . The Solid Flow/Valve-Controller Model • . . . . . . . . . . Simplified Energy Model . . .• iv viii xii xiii xvi 1 6 8 9 15 22 27 28 34 42 45 59 60 68 73 76 81 84 VI. EQUIPMENT AND EXPERIMENTAL PROCEDURE VII. Description of the Laboratory Unit • • • Description of the Data Acquisition System .•••...•..•• Data Management . • • • • • • RESULTS AND DISCUSSION Dynamic Characterization of the Laboratory unit .•.•. Thermal Dynamic Simulation Using the Two-Stage Energy Model • • • Stability of the Proposed Process ••. Data Analysis Using the Five-Stage Model • . • . • . VIII. CONCLUSIONS AND RECOMMENDATIONS. Appendices A. THERMODYNAMIC DATA BASE FOR THERMAL PROCESSING OF TAR SANDS .•• B. FIVE-STAGE MODEL PROGRAM C. D. STEADY-STATE SIMULATOR PROGRAM FOR THE MAXIMUM HEAT FLOW IN A POTASSIUM HEAT PIPE • • . • • • • MASS FLOW SIMULATOR PROGRAM • E. TWO-STAGE ENERGY MODEL APPLIED TO THE UNIVERSITY OF UTAH IABORATORY UNIT •• F .. LISTING OF DATA ACQUISITION ALGORITHM • G. PROGRAM LISTING OF DATA PROCESSING AND DISPLAY • • • . . • • . REFERENCES vii 89 90 98 104 III 112 127 136 142 152 160 171 183 189 195 205 224 241 LIST OF FIGURES Figure 1. Optimal pyrolysis temperature for oil recovery. . . . . . . . . . • . . . . . . . • . . . . . . . . . . . . . • • . . . 14 2. University of Utah thermal process, Weeks (1977), with permission .•..•• ~......... 17 3. University of Utah thermal process, Seader and Jayakar (1979), with permission.......... 21 4. University of Utah thermal process (~983).... 23 5. Effect of the fluidizing gas condition over the total pyrolysis energy requirement for different reactor temperatures........... 37 6. Effect of the amount of fluidizing gas over the total pyrolysis energy requirement for different reactor temperatures........... 38 7. Effect of the amount of fluidizing gas over the total pyrolysis..................... 39 8. Energy available from combustion of coked sand for various coke yields from pyrolysis............................... 43 9. Energy available from combustion of coked sand at various reactor temperatures......... 44 10. Energy available from combustion of coked sand. Effect of preheating fluidizing air... 46 11. Three different modes of transferring energy from combustion to pyrolysis. a.- Preheating of fluidizing gas. b.- Recycle of hot sand. c.- Using heat pipes... 47 Ene'rgy balance for the combustion reactor. Energy transferred from combustion reactor to pyrolysis reactor by preheating the fluidizing gas................................ 48 13. Energy balance for the combustion reactor. Energy transferred from the combustion reactor to the pyrolysis reactor by recycle of hot sand.......................... 49 14. Energy balance for the combustion reactor. Energy transferred from combustion to pyrolysis by using heat pipes................ 50 15. Heat exchanger stage representation.......... 61 16. Reactor stage representation................. 62 17. Uni versi ty of Utah 'lprocess. Stage representation............................... 74 18. Maximum predicted heat transfer flow for the University of Utah heat pipe......... 80 19. Schematic representation of the solids discharge control system..................... 83 20. Basic material diagram of the unit........... 91 21. Process flow diagram of laboratory unit...... 92 22. Instrumentation diagram (computer excluded).......................... 96 23. Schematic representation of the equipment and computer....................... 99 24. Performance of the average filter algorithm over data from thermocouple no. 11........... 106 / 25. Data management. a.- Data acquisition flow chart. b.- Data display flow chart. Software shown in parenthesis................ 107 Reactor pressure drop history. (Data in file FILE91.)................................. 113 Reactor temperature history. Experiment number 7..................................... 114 Reactor temperature history. Experiment number 12.................................... 115 j 29. Reactor temperature history. Experiment number 13.,................................... 116 ix 30. Reactor temperature history. Experiment number 14.................................... 117 31. Reactor temperature history. Experiment number 15.................................... 118 32. Reactor temperature history. Experiment number 16.................................... 11 9 33. Expanded view of experiment number 14 ........ 34. Expanded view of experiment number 15 ........ 35. Expanded view of experiment number 16 ........ 36. Reactor temperature history. Experiment number 17 .................................... 37. Two-stage energy model prediction when no energy flow limit is imposed on the 122 123 124 125 heat pipe performance........................ 129 38. Two-stage energy model prediction when the energy flow through the heat pipe is limited by viscous forces at low temperatures. . . . . • •• • • • • . . • • • . . • . . • • • . . • • • . . . 130 39. Two-stage energy model prediction when the energy flow through the heat pipe is limited by sonic velocity at low temper a t ure s •..•••.......•••..•.....•.•...... 40. Two-stage energy model prediction using parameters from experiment number 7 . • • • . • • • . . 41. Two-stage energy model prediction using parameters from experiment number 13 ......... 42. Two-stage energy model prediction using parameters from experiment number 14 ......... 43. Effect of the initial heat pipe temperature on the numerical solution of the set of 131 133 134 135 equations. . . • . . . . . . . . . • . . . . . . . • . . . . . . . . • . . • . . 138 44. Stability of the system for constant total energy generation of 7500 Btu/hr............. 140 45. Stability of the system for constant total energy generation of 5100 Btu/hr............. 141 x 46. Reactor pressure drop. Pneumatic control of the solids discharge valves............... 143 47. Model prediction using parameters from experiment number 12......................... 148 48. Model prediction when feeding tar sand....... 150 49. Some specific heat capacities as predicted from data file............................... 162 50. Comparison of tar sand heat capacity from literature and data base..................... 163 xi LIST OF TABLES Table 1. Elemental composition of reactants and products in equations (4) and (5)............ 31 2. Characterization of reactants and products... 32 3. Parameter sensitivity analysis............... 41 4. Lost-work analysis at minimum energy self-sufficieny condition.................... 55 5. Lost-work analysis at maximum thermodynamic efficiency .•••..•.•.. -• . • . • . • • . • . . • • . . . . • . . • . • 57 6. Dimensionless parameter usage at each stage.. 77 7. Parameters describing the operating conditions of the experiments at start-up.... 120 8. Cyclic response of the reactor pressure drop to pneumatic control.................... 144 9. Hot sand-to-wall heat transfer coefficient... 147 10. Available thermodynamic data................. 161 A a B Cp dp e F f G G g H = = = = = = = = = = = = NOTATION Cross-sectional area Time constant for energy balance Availability Heat capacity particle diameter Fractional deviation from reference (Equation 42) Solid mass flow rate Function Gas mass flow rate Dimensionless gas flow rate (Equation 35) Dimensionless energy generation Enthalpy = Convective heat transfer coefficient = Integration time (Equation 43) = Control constant (Equation 43) = Conduction heat transfer coefficient = Stage height = Heat pipe perimeter = Solid flow perimeter = Wall perimeter LW M,m Q q r s rxn s t v w z z a a l y '6H r = = = = = = = = = = = = = = = = = = = = = = = Lost work Mass Control pressure Electrical power Energy input Dimensionless energy input Reactor effective radius Reaction rate Entropy Temperature Time Velocity Heat pipe active area fraction Axial direction Dimensionless axial direction Dimensionless gas-wall heat transfer coefficient (1 - Ep)a Dimensionless gas-solid heat transfer coefficient Dimensionless sensible heat flow Heat of reaction Dimensionless radiative heat transfer coefficient Reactor particle-phase fraction Wall emissivity xiv = = = = = = = = Subscripts: av = g = gs = gw = HP = = j = max = :r = ref = s = surr = sw = w , = = Perimeter ratio Dimensionless solid-wall heat transfer coefficient Dimensionless temperature Dimensionless electrical power Density Stefan-Boltzmann constant Time constant for energy balance Dimensionless time at wall energy balance Average Gas phase Gas-solid Gas-wall Heat pipe Reactor number Stage number Maximum Reactor Reference state Solid phase Surroundings Solid-wall tiall Ambient xv ACKNOWLEDGMENTS I wish to express my sincere gratitude to Professor J. D. Seader for his stimulating guidance and encouraging support throughout this work. I also wish to acknowledge my appreciation to all my professors and fellow graduate students for their assistance on this project and other related subjects. Special thanks are gratefully given to my friends Dr. Pedro A. Aylwin and Dr. Osvaldo A. Bascur for our many valuable discussions. Partial financial support by the Award DE-ATLO­SOLO 10331 and 2 DOE/Laramie, Wyoming and the University Research Committee is also acknowledged. Grateful appreciation is extended to my mother and for their encouragement during my schooling. Finally, a very special word of thanks to my wife, love, motivation, and understanding. CHAPTER I INTRODUCTION Production of petroleum and natural gas in the United States is decreasing. oil production peaked in 1970 at 3.5 billion barrels, and natural gas peaked at 21.7 trillion ft 3 in 1973. Since then, both petroleum and natural gas production have been on the decline in r~aation to the continuously increasing domestic energy consumption. These trends reflect the fact that domestic oil and gas are becoming more difficult and expensive to find and produce, as development is moving to the exploi­tat. ion of deeper wells and the exploration of areas of di£ficult access like the Alaskan area or the ocean floor. This increasing demand for energy and the simultaneous tendency to decrease dependence on unpredictable foreign energy sources have justified an increasing effort for studying and developing technology for the production of synthetic crude oil from alternative sources like coal, shale oil, or tar sands. As a source of synthetic crude oil, tar sand deposits are characterized by a very high viscosity bitu­. en impregnated in consolidated sand: in particular, States deposits of tar sand have a higher viscosity 2 bitumen than other known deposits due mainly to a higher molecular weight (Bunger, 1979). United States reserves of tar sand are estimated between 25-30 billion barrels of bitumen, with more than 90% of this total located in Utah: therefore, it seems economically attractive to process this source in spite of its high viscosity. Current work is being done on the development of technology to process Utah tar sands at two levels of pro­duction. At laboratory and bench scale, basic research is being conducted at the University of Utah on pJ:'.ocesses like hot-water extraction, solvent extraction, and thermal recovery of tar sands; simultaneously, at pilot-plant scale, a 50-100 barrels per day plant has been designed, constructed, and operated by Enercor in Salt Lake City utilizing University of Utah data and technology on hot­water processing of tar sands. Advantages of the hot-water extraction technique, like simplicity of the process and low operating tempera­ture (below 250°F), are easily counterbalanced by high energy requirements, significant water consumption, and very high product viscosity. This reasoning seems to justify the increasing interest in thermal recovery techniques that can produce a lower vi S.CQSi ty product wi th a l most no water consumption and no external energy supply. On the other hand, thermal processing of tar sands inv-olves higher operating temperatures (IOOOoF or above), higher technological risk, and higher capital require­ments. 3 Thermal processing of tar sand is characterized by two simultaneous reactions occurring at two separated places in the process. In the first reaction, bitumen is pyrolyzed at 450-550oC and synthetic crude is obtained together with gases and coke. This reaction is endother­mic where more than 80% of the required energy is used to heat the sand to reaction temperature. The second reac­tion involves the combustion of a by-product of the first reaction if available, or the combustion of a fuel in I general, in order to obtain the energy required for the pyrolysis reaction. Different types of thermal processes ar~ originated by the way that energy is transferred from the combustion reactor to the pyrolysis reactor. Thus, recycle of large amounts of hot sand was investigated by Gishler (1949) in a cat-cracking type process and by Rammler (1970) on a variation of the Lurgi-Ruhrgas proc­ess. Also, at the University of Utah, preheating of the pyrolysis reactor fluidizing gases has been tested by Weeks (1977) as a way of transferring the required energy from the combustion reactor to the pyrolysis reactor. Later, the same process was modified by the addition of a potassium heat pipe in order to increase the energy flow reactor to the pyrolysis reactor; this process, known as the University of Utah thermal process was developed by Seader et ale (1979). 4 A somehow different thermal recovery technique is described by Chalmers et ale (1980), where the tar sand is agglomerated into discrete pieces that are treated on a circular traveling grate and subjected to sequential treatment with hot gases. The heat required is obtained in a coke burn-off zone from combustion of the residual coke by a recycle of the gas obtained from the distilla­tion zone though the hot material. Another process, developed by Audeh and Chen (1979), utilizes recycle of hot sand to transfer the required amount of energy from the combustion reactor to the pyrolysis reactor, and operates the pyrolysis reactor at 600-8S0oF and ±60 Ib/in2 pressure, and the combustion reactor, located above the pyrolysis reactor at lS00-2000 oF. The work presented here involves the improvement and dynamic simulation of the University of Utah thermal process described by Jayakar (1979): also, a thermodynamic simulation and comparison of three alternative techniques of transferring energy from the combustion reactor to the pyrolysis reactor is included. The dynamic simulation of the University of Utah thermal process is carried out at two levels of complexity. A five-stage model is developed tq simulate transient behavior of the University of Utah's existing unit. Finally, a two-stage model is used to 5 study control strategies and stability of the unit in gen-eral terms. CHAPTER II OBJECTIVES The objectives of this research program can be classified in five groups: 1. Thermodynamic simulation. Examine different existing and proposed alternatives for transferring energy from the combustion reactor to the pyrolysis reactor from a thermodynamic point of view. Determine limiting fac­tors and possible new alternatives. Process design. Modify and improve the currently existing University of Utah thermal process experimen­tal unit with emphasis on product recovery and energy efficiency. 3. Dynamic modeling. Develop a dynamic model to simUlate the University of Utah modified thermal process in order to study and interpret real-time data obtained from the experimental unit. Investigate the effect of the different transport phenomena involved in the different sections of the equipment on the overall behavior of the modified unit. 4. Use a modified version of the dynamic model to study control strategies from a general point of view, and assess the impact of using heat pipes on the control­lability and stability of the proposed process. 7 5. Automate the modified unit for computerized data acquisition, to provide the information required for items 3 and 4 above, and set the guidelines for future digital control of the existing experimental unit or any larger-scale unit. The first two objectives are somehow related and are add'ressed in the following two chapters. The third fourth objectives are treated together in Chapter V. IChapters VI and VII are dedicated to objective 5 and to and discussion of the results. The last a summary of the results and experience gained during the course of this research in the form of CHAPTER III BACKGROUND AND PROCESS DEVELOPMENT IN THERMAL PROCESSING OF TAR SANDS The rapid increase in the price of petroleum since the OPEC oil embargo of 1973 has activated interest in the ' development and exploitation of Utah tar sands with a oonsequent increase in literature and research programs on this resource. However, the future of a commercial oper­ating plant is still not clear and economic feasibility depends basica.lly on government regulations and the unpre­di. ctable behavior of the world oil market, which has wit­nessed recent decreases in crude oil prices. Commercial processing of tar sands is presently carried but in Canada by GCOS Ltd. and Syncrude Canada Ltd. with a process that involves hot-water extraction of the bitumen followed by delayed coking to produce a synthetic crude. Due to the different nature of the bitumen contained in Utah tar sands compared to that in the Canadian deposits, direct application of Canadian technology has been unsuccessful. Because of this reason and the high consumption of energy when processing tar sand by hot-water extraction techniques, interest has been focused on developing 9 alternative methods of processing tar sand such as sOlvent extraction and thermal recovery methods. In-situ recovery techniques are necessary for deposits with large Overburden-to-payload ratios. This chapter presents the development regarding the thermal recovery methods. In particular, special " attention is given to the development of the University of utah ' fluidized-bed thermal process. Prior to examining I the most recent developments in thermal recovery technol­iogy, a review of the literature is presented in order to provide the reader with sources of information regarding aspects of tar sands like nature, geology, and reserves, as well as technical aspects such as product quality, process technology, and currently active proc-esses. Review Since the oil embargo of 1973, information' about characterization, geology, and recovery technology of Utah tar sands has been increasing steadily. Most information comes from publications covering progress of the Univer­sity of Utah and the United States Department of Energy research. The reader who is interested in information sand deposits located outside the United States other publications (Gutierrez et al., 1977; Phizac'kerley and Scott, 1967; and Carrigy, 1967). Geology and reserves of the Utah tar sand deposits are presented in the work of Ritzma (1975). 10 An extensive compilation of information about processing and utilization of tar sands can be found in the works of Jayakar (1979) and Venkatesan (1979). Their reviews cover most of the already existent technology for .recovering oil from tar sand deposits by different proc­esses like hot-water extraction, solvent extraction, and thermal recovery processes until 1979. Also, literature surveys on tar sand recovery and extraction are adequately covered by the work of Bunger (1979). Therefore, the literature survey presented here will cover main develop­ments on thermal processing of tar sands published after 1979. The previously mentioned references are recommended to the investigator interested in earlier developments in the technology, characterization, and geology of tar sand resources. Since 1979, at least three patents have been issued for different methods of processing tar sands by thermal pyrolysis. The first is an innovative method developed at the University.of Utah by Seader and Jayakar. process consists of fluidized-bed pyrolysis of tar £'rom ambient-temperature feed followed by fluidized­combustion of the resulting coked sand by-product of olysis, with 85 to 95% of the required energy for the lysis reactor transferred from the combustion reactor 11 by multiple heat pipes. The use of heat pipes allows a lower operating temperature in the combustion reactor compared to a fluidized-bed pyrolysis/fluidized-bed com­bustion process, where the thermal coupling is by the flue gas only. Thus, the Seader-Jayakar process is an improve­ment on the thermal efficiency of the overall process as qiscussed later. The second method, developed by Audeh and Chen (1979) for Mobil Oil Corporation, is a redesign of the process first proposed by Peterson and Gishler (1951). The process utilizes low-temperature distillation bf bitumen, in the range of 600 to 850°F, in the presence of relatively low recycle of hot sand to provide the required pyrolysis reactor energy. The pressure of the distillation reactor is maintained in the range of atmos­phf; 1!ric to 100 psia. The hot sand recycle is produced from a fuel generating reaction between coked sand, steam, and air operating at temperatures above lS00 oF. The limiting factor of this proposed design is that rich tar sand, with a bitumen content of 10 wt % or above, is required to satisfy the energy balances (the designer proposes to use bitumen). The thir.d thermal process, developed by for A. G. McKee & Co., consists of a circular traveling grate where the tar sand, previously a«J91o;merated into discrete pieces, is subjected to sequen­& 1 treatments with hot gases passed upwardly or down­through a relatively deep permeable bed of the 12 pieces. All of the energy required at the pyrolysis section of the path is obtained in a coke burn-off zone from combustion of coke that remains in the material on the grate after the pyrolysis of the bitumen. The heat produced at combustion is tranferred to pyrolysis by recycling part of the hot solid and by recycle of hot gases. Drawbacks of this design are the large residence times required for heat-up and the need for tar sands with high bitumen content in order to reach proposed operating temperatures. All the thermal processes examined can be easily applied to the recovery of bitumen from Canadian tar sand deposits because of the high bitumen content of this resource. On the other hand, the adaptability of these p~ocesses to the recovery of Utah tar sand deposits is strongly limited by the thermal efficiency of the proc­esses examined because of the low bitumen content of the Utah tar sands. Therefore, the production of coke and subsequent combustion is a critical factor in the develop­ant of a self-sufficient thermal process for the recovery oil from tar sands (Oblad et al., 1976). When designing a reactor to process tar sand by ,....~440. of pyrolysis, or thermal processing, it is of funda­importance to know and understand the effect of variables such as temperature, pressure, and reten­time. Also, the nature of the feedstock will greatly 13 influence reactor performance. These variables and others such as energy recovery, product characterization, product recovery, and material handling are currently being studied at the University of Utah with laboratory-scale equipment. Although most of the efforts have been focused on fluidized-bed pyrolysis (Weeks, Hanks, Jayakar, Venka­tesan), some studies have been made on diredt coking, visbreaking, and catalytic cracking (Bunger, 1979). As shown in Figure 1, results indicate the existence of an optimal pyrolysis temperature, to maximize yield of syn­crude, within the range of 425°C to 550°C depending initial bitumen content of the tar sand processed. Above 550°C, oil yield is decreased by increasing gasifi­cation (C1 to C4 production). Below 425°C, the reduced e~tent of pyrolysis of the bitumen decreases the oil yield arid increases the amount of unreacted bituminous material adhered to the sand particles. Another important variable of the thermal process­ing technique, the yield of coke in the pyrolysis reactor, is meing consistently reported within the range of 16 to 24 wt % of bitumen, almost independent of temperature or retention time for temperatures between 4500 to 550°C and retention times above 16 minutes. Then, it seems reason­able to assume that a constant supply of coked sand will in the pyrolysis reactor, to be later s::: (l) 3 60 +J 'r-! .Q 4-4 o 40 20 o 400 DeEosit Bitumen Ref. o Sunnyside 8.5 Venkatesan o Sunnyside 10. Hanks o Tar Sand Triangle 4.5 Venkatesan \} Tar Sand Triangle 4.7 Jayakar 450 500 550 600 650 Pyrolysis Temperature (oe) Fig. 1. Optimal pyrolysis temperature for oil recovery. 15 combusted with air in a second reactor in order to gener-ate part or all of the energy required by the process. Results presented by Bunger (1979) indicate that product distribution is more sensitive to operating temperature than pressure of pyrolysis: also, yields of atmospheric coking are comparable to catalytic cracking. Therefore, a direct incentive does not appear to exist for operating the pyrolysis reactor at pressures much above ambient conditions. Effect of retention time on oil yields and product characterization has been reported by Venkatesan (1979), 1 working with Sunnyside and Tar Sand Triangle samples in a pyrolysis reactor. The reported data ~ndicate that oil yields over 80 wt % can be achieved with retention times in the range of 12 to 16 minutes. Although no data are reported for retention times below 12 minutes, a lower oil yield is expected because of increased unreacted feed. Similarly, extended retention times decrease oil yield with an i9creased production of ~revious Developments of the ' University f AUtah Thermal Process The first equipment designed for thermal process University of Utah to recover oil from tar sands ~isted of a two-staged fluidized bed. General features this process, as presented by Weeks, are shown in 16 Figure 2. Bituminous sand is fed directly into the coking bed, where bitumen is cracked and vaporized at 4S0oe to SlOoe. Products of this section are vaporized oil, gases, and coked sand. The coked sand is transferred by gravity to the combustion bed where it reacts with air, generating the energy required at the pyrolysis reactor. The energy dissipated in the combustion reactor is transferred as sensible heat with the hot gases product of combustion to the pyrolysis reactor. Provisions are ' included to add propane or any combustible gas directly into the combus-tion bed with the purpose of generating enough energy to satisfy pyrolysis demand and to eliminate by combustion any trace of oxygen remaining after combustion. Some of the clear advantages of this process with respect to other non-thermal processes are: reduction of water consumption to a minimum, ease in handling solids, . clean by~products, higher quality product, and energy self-sufficiency depending on the content of bitumen and the selection of the operating variables. Advantages of secondary importance are near-isothermal operation, high heat transfer rates, and no hot sand recycle, keeping the size of the reactors to a minimum for a given flow of Disadvantages of this processing scheme are a very operating t ,emperature in the combustion reactor and a arge flow of fluidizing gas in the pyrolysis reactor, Upper Expansion Section Combustion Gases Tar Sand Feeding Section ': \ .~'. _._._-- . ..~.. - :!~t1t~(%:' '~j~:~~~: ~:?}:} :--'.~' . "'.'" Combustion Gases and Products to Recovery Section Pyro1y~is Section Coked Sand Central Expansion Section Combustion Section Propane Air if:/~~j~~i ]'~:>/~:/':~"'< .~ ,;.,:; .... ·::.'tA ..; Burner ___ •• :::1 Fig. 2. Spent Sand University of Utah Thermal Process, Weeks (1977), with permission. 17 18 without considering the typical drawbacks of using fluid­ized beds such as possibly large pressure fluctuations into the beds, dust pollution caused by attrition of the l'articles, difficult scale-up, bed pressure drop, and unavoidable high coke combustion temperature. In particu­lar, just to keep constant the ratio of pyrolysis fluidiz­ing gas to tar sand feed rate in order to maintain the required transfer of energy from combustion reactor to pyrolysis reactor when scaling-up the design, the design ratio of fluidizing velocity to minimum fluidizing veloc­ity must be increased proportionately to the reactor height, thus limiting the maximum size of the fluidized bed. other 'observed disadvantages of this process are difficulties in the flow of solids through the central expansion section due to a pressure gradient opposite to the direction of the flow and the dilution of the gaseous products with the large amount of combustion gases making more difficult the process of product separation. Also, in order to reach the required operating temperatures in the combustion reactor, tar sand with bitumen content of 11% or above is required if energy self-sufficiency is For these reasons, which were well understood by Weeks, the process as shown in Figure 2, was never oper­ated in a coupled manner. The only runs made were with separate operation of either the pyrolysis section or the combustion section. 19 At this point, it is important to notice that the above methoa for transferring energy from the combustion zone to the pyrolysis zone, by using the sensible heat of the hot gases produced in the combustion zone, was pro­posed and tested on a laboratory scale at the University of Calgary by Moore et ale (1979). Their proposed process consists of fluidized-bed pyrolysis of tar sand followed by combustion of the coked sand with a mixture of air, oxygen, nitrogen, and steam in a moving-bed reactor. As expected, the data presented indicates that' the process is capable of transferring the necessary amount of energy f~om combustion zone to pyrolysis zone because the ratio of fluidizing gas to tar sand feed is sufficiently high for the difference in temperatures between the two reac­tion zones utilized in their experiments. The major drawback of this process is encountered when scaling-up the throughput of the unit because the maximum value of the' above-mentioned flow ratio decreases with an increase in fluidized bed height if operating conditions are to be maintained (such as reactor temperature and solid reten­tion time). Then, for large-scale operation of this process, very high combustion temperatures are required in order to provide the necessary amount of energy transport the pyrolysis reactor. 1 To solve one of the basic problems of thermal rocessing, the potentially high operating temperature of 20 the combustion reactor, the first design was modified by Jayakar (1979) to permit thermally coupled operation. The modified design makes use of heat pipe technology to transfer most of the energy required in the pyrolysis reactor from combustion of the coked sand in the combus-tion reactor, as shown in Figure 3. Detailed information on the proposed process is given in u.s. patent 4,160,720. operation of the unit is similar to that described for the first process by Weeks, with the addition of the behavior of potassium heat pipes, which are recommended for the range of temperatures of the patented process (possibil-ities of using sodium or cesium as an alternative working fluid instead of potassium is also mentioned). The modi-fied unit clearly improves the efficiency of the process by decreasing the minimum content of bitumen required to achieve energy self-sufficiency to 8-9% from 11% required for the first process. Efficiency is also improved by including production of steam by heat transfer from the hot sand after preheating the combustion air. Combustion o opera~ing temperature is lowered to the 550-600 C range, lowering reactor costs. Major drawbacks of the second design are similar the first design with exclusion of high combustion rating temperature which has been decreased by the use heat pipes. Scale-up procedures are also eased by the heat pipes in the sense that geometry will 21 Tar Sand Feeding Section Product and Gases to Recovery Section Combustion Gases Heat Pipes Pyrolysis Section Combustion Section Air Preheating Section Steam • Spent Sand Fig. 3. University of Utah Thermal Process, Seader and Jayakar (1979), with permission. 22 not change appreciably from pilot-plant to commercial unit. In essence, the heat pipes serve to compartmental-ize a large fluidized bed into a number of smaller paral-leI beds. On the other hand, the introduction of heat pipes in the design increases 'the technological risk of the proposed process. New Developments of the University of Utah Thermal Process After demonstrating the technical feasibility of operating heat pipes in the range of temperatures and ehergy flows required for thermal recovery of oil from tar sands, it is a direct consequence to suppose that, for some specific design conditions, the heat pipes will be capable of transferr,ing, from the combustion reactor, 100% of, the total energy required at the pyrolysis reactor. ;, Thus, it is no longer necessary to use the sensible heat of the combustion gases at the pyrolysis reactor and the'refore any appropriate and available gas can be used to fluiClize the pyrolysis reactor in the manner shown in I i Figure 4. For practical purposes, i.e., industrial scale, possible choices of a fluidizing gas for the pyrolysis eactor are reduced to two: (1) steam, which can be sily produced on a large scale from energy recovered om the hot sand, and easily obtained at the laboratory l~1 and (2) light ends, a mixture of C1-C4 compounds Pyrolysis Section Heat Pipes Combustion Section Make-up Fuel (Coal, Gas) 23 Tar Sand Feeding Section Spent Sand Products to Recovery Section ~ Fluidizing Gas (Steam, Light Ends) Combustion Gases Air Steam University of Utah Thermal Process (1983). that is obtained as a by-product of the pyrolysis reac-tion. Both gases can be easily separated from the syn-thetic oil produced in the pyrolysis reactor, and recycled \ in the required amount. In both cases, perfect separation of the recycled fluidizing gas and the products is not required, thus simplifying the design. Immediate advantages of the process shown, in Figure 4, with respect to the previous proposed process in 3, are: Increased product concentration by using an amount of fluidizing gas to the pyrolysis reactor that is close to the minimum required for fluidizing the bed. Indirectly, the heat transfer coefficient can be increased to its optimal value when operating the fluidized bed at low fluidizing velocities, as pointed out by different authors (Grace, 1982; Botterill, 1975; Richardson and Shakiri, 1979). Simplified pyrolysis reactor design by fixing the amount of fluidizing gas to be used. The flow rate ratio of fluidizing gas to solid feed required for minimum fluidization decreases with an increase in the height of the solids inside the reactor; on the other hand, transferring energy to the reactor by using the sensible heat of the fluidizing gas tends to fix the value of the flow rate ratio mentioned above. Then, 25 if use of sensible heat is proven possible at laboratory-scale, its use in large-scale fluidized beds will be extremely difficult, if not impossible. Increased controllability of the process by decoupling fluid-dynamically both reactors. The previous thermal process consisted of a system of two fluidized beds thermally coupled and fluid-dynamically coupled with thermal time constants above 30 minutes and fluid­dynamic time constants, depending on the reactor heights, below one minute. This situation presents difficult problems to the control designer because of the short fluid mechanic time response of the fluid­ized bed compared to the thermal response. The prob­lem is simplified by fluid mechanicallydecoupling both reactors. Decreased reactor size. Preheating of the fluidizing air is accomplished by indirect heat exchange with the hot combustion gases that leave the combustion reac­tor. Hot sand is used for the sole purpose of produc­ing low-pressure steam, thus simplifying the process and increasing the global thermal efficiency of the proposed design. Increased thermal effi6iency of the energy recovery system by preheating the air with the hot combustion gasas. Both streams are of similar flow magnitUde and very close heat capacity. Efficiency is also 26 increased by decreasing the required compression power used to move the fluidizing streams. 6. Operating pressure at the bottom of the combustion • l : reactor is decreased by a significant factor because it is no longer necessary to fl~idize both beds with the same stream. Thus, in essence, both beds can now operate at the same pressure. Of all these advantages, items 2, 4, 5, and 6 are strongly plant-size dependent, i.e., the advantage will increase with increased size of plant. Advantages 1 and 3 already can be observed at the laboratory scale. The main disadvantage of this proposed design as compared to the one of Jayakar (1979), is the loss of the use . of sensible heat in the fluidizing gases fed into the pyrolysis reactor: but, as will be shown in the next chapter, this available sensible heat decreases with the size of the unit, being less than 1% of the total energy required at the pyrolysis reactor for a large plant when operating the pyrolysis reactor at 450°C, the combustion reactor at 550 °C , and processing at 10 wt % bitumen-rich This minor drawback is easily counterbalanced clear advantages mentioned previously. CHAPTER IV THERMODYNAMIC ANALYSIS OF THERMAL RECOVERY OF TAR SANDS • t Because of dwindling natural resources and ques-tionable availability of conventional liquid fuels, considerable emphasis is currently being placed on the development of alternative energy sources and processes, and the improvement of existing ones. Search for new energy resources has originated numerous studies of uncon­ventional sources like shale oil, oil from tar sand, and other heavy oils. Improvement of existing technology includes development of more efficient alternative processes, increased recovery of energy from waste or by­p bducts, and efficient resource utilization. - When modifying an existing process or developing a newf one, the first and second laws of thermodynamics can be applied to characterize each process by computing ~inimum or maximum work produced or required and energy consumption or production. Then, comparison between alternative processes can be carried out by total thermodynamic irreversibilities or thermo­Two different methods of using the of thermodynamics are presently 28 available. The first method, introduced as availability analysis by Keenan (1951), estimates the maximum amount of work that could be reversibly produced by bringing a unit of mass to thermodynamic equilibrium with a properly defined surrounding. The second method, developed by de Nevers and Seader (1980), calculates the difference between actual work and thermodynamic reversible work of a specified process as the actual lost-work of the system: consequently, this method is known as lost-work analysis. I Advantages and disadvantages when using either method are discussed by de Nevers (1981). Thermodynamic lost-work analysis has been chosen for this study because this me~hod simplifies the selection of reference states and eases the task of comparing alternative processes where no efficiency definition exists. The first step in the synthesis of an optimal thermal method for processing of tar sands is. to define a consistent thermodynamic data base. Then, the different alternatives are analyzed in general terms, searching for general rules and limiting factors. Finally, lost-work applied to compare the proposed alternatives. uations Application of the first and second laws of rmodynamics to an open, steady-flow, steady-state requires knowledge of enthalpy and entropy of each 29 component present referred to standard chemical species. At low pressures, these can be obtained by proper integration of the specific heat capacity Cp, Perry , (1973): ,- H{T}J' = ~Hf{T f}' + TfT CPJ·{T}dT@constantP re J ref ( 1 ) S{T}J' = ~Sf{Tref}J' + TfT Cp.{T}dT (2) ref J T @ constant P ~Hf{Tref} and 6S f {Tref} are the enthalpy and entropy of formation from the standard chemical species per unit of mass of component j at reference temperature T ref. Pressure effects have not been considered in this study because pyrolysis of tar sands and consequent combustion of coked sand are carried out at pressures very close to atmospheric level. Heat capacity CPj{T} is obtained by using least correlation of proper data with the equation: N. J = 1: a. ,b.{T} i=l ~J ~ b.{T} is any specified function of temperature T, ~ constant coefficient independent of the t;t:mperature. The reactions involved in pyrolysis of tar sand combustion of coked sand considered for this study ( 3 ) 30 involve the use of composite components, called products, due mainly to the complexity of the components involved. Hence, in general terms, the main pyrolysis and combustion reactions can be represented by mass-balanced equations: bitumen ~ oil + light ends + coke @ pyrolysis, (4) Barbour et ale (1976) + air ~ combustion gases @ combustion. ( 5) Average elemental composition of the compounds ~articipating in the pyrolysis reaction have been obtained from the work of Bunger (1979). Composition of flue gas is obtained by total combustion of coke with the stoichio-metric amount of air. Average compositions are presented Table 1 as weight fractions of the total product. In order to characterize thermodynamically each from Equations (4) and (5), data obtained from different sources have been used. In particular, hydro-carbon data have been obtained from API (1966) tables, sand and gas properties have been obtained from JANAF (1960) tables, and bitumen and oil properties have been estimated by using an atomic contribution method with the ch~racteristic parameters shown in Table 2 (Wuithier, Once ~nthalpy and entro~y of a component j have computed, availability B· can be readily calculated J by applying the equation: Table 1 Elemental composition of reactants and products in equations (4) and ( 5) . Product Carbon Hydrogen Nitrogen Sulfur Oxygen Argon Utah bitumen 84 - 85 10 - 12 0.8 1 - 2 1 - 2 oil 86 - 87 11 - 12 0.4 1 - 2 Light ends 66 - 78 8 - 27 1 - 2 o - 3 Coke 87 - 88 2 - 3 2 2 Air 75 - 76 23 1 Flue gas 8 0.2 70 0.1 21 1 Table 2 Characterization of reactants and products. Characteristic H/c ratio density sp. gr. Average BP Characterization Factor K Average MW H/c ratio density sp. gr. Average BP Characterization Factor K Average MW H/c ratio Approximate composition: Average MW H/c ratio Value 0.130 by wt. 4 - 12 °AP! 0.986 - 1.04 580 - 620°C 11.2 - 11.8 560 - 600 0.133 by wt. 17 - 24 °AP! 0.91 - 0.95 400 - 420°C 11.3 - 11.8 340 - 360 0.243 by wt. 432122 wwwttt %%% CCc12 5 wt % c3 23.4 4 0.03 by wt. 32 B {T}. = H {T}. - T S {T} . J J surr J where Tsurr' the surrounding temperature, is taken as 298 K for this study. 33 (6) From the combined first and second laws of thermo-dynamics, a relation, derived to compute the thermodynamic lost-work of a steady state flow process, is presented by de Nevers and Seader (1980) as: JI dLW = ~ B.dm. + L (1 - j J J k where dLW (always greater than or equal to zero) repre- ( 7) sents the total thermodynamic lost-work of a differential control element of the system under study, L B.dm. j J J represents the total net flow of availability into the ,) differential control elemeht, ~ [1 - (Tsurr/Tk)]dQk represents the total flow of energy into the differential control element as Carnot-equivalent work, and E dW 1 l represents the total shaft work obtained from the differ-ential control element including electrical or mechanical and excluding the injection or flow work that has considered in the availability term. Finally, the data collected have been organized data bank consisting of 11 compounds that are most when studying thermal processing of tar sands. From data bank, H., S., and Cpo can be easily obtained by J J J 34 direct call to subroutines HT, ST, and CP, respectively, all of which are included into a file &THRMI. A detailed description of this data bank and a listing of the file &THRMl is given in Appendix A. Pyrolysis Reactor In order to estimate the levels of energy required by the pyrolysis of tar sands it is necessary to make an energy balance around the pyrolysis reactor. For this purpose, some justifiable assumptions can be made: 1. Tar sand feed at ambient temperature. Although pre­heating of tar sand feed seems attractive in the sense of decreasing the need for high level heat sources for the pyrolysis reactor (similar to Lurghi-Rurhgas process), the increased agglomeration and consequent difficult handling of the feed make this alternative very questionable. The coked sand and the fluidizing gas leave the reac­tor at the reactor temperature. This assumption seems reasonable in the case of the coked sand product the high residence time of the solid (above 10 minutes), but the fluidizing gas is more likely to leave the reactor at a temperature between the reactor temperature and the gas inlet temperature because of the low heat transfer coefficient between fluid and particles/wall and a characteristic short residence 35 time of the fluid in a bubbling fluidized-bed. On the ,other hand, product released from the reacting bitumen at pyrolysis temperature will help the fluidizing gas to reach reactor temperature, as will the well-mixed condition in the bed. 3. Coke to bitumen ratio of 0.20. As previously stated, observed coke yields indicate little dependence of pyrolysis temperature for reactor temperature above 425°C. Typical values of coke yield reported are included in the range of 18 to 22 wt % of the initial bitumen content. Oil to bitumen ratio of 0.70. Reported values for this ratio are included in the range of 0.50-0.85 depending on pyrolysis temperature, solid retention , time, initial bitumen content, and experimental recov- .\ ery technique used. In spi te of the fact that the value of this parameter considerably affects the design of a tar sand processing plant, little effect ~s observed when calculating the energy load of the . pyrolysis reactor. Other parameters to be considered are the pyroly­reactor temperature and the initial bitumen content. reactor temperature is limited to the range of 425 to 525°C to optimize oil production as previously stated. bitumen contents of 4 to 12 wt % are typically found in Utah deposits of tar sands. Both reactor temperature and initial bitumen content are important parameters when determining the energy load of the pyrolysis reactor. 36 Independent variables are the feed rate of tar sand, the flow rate of fluidizing gas, and temperature and composition of the fluidizing gas. Practical choices of fluidizing gas are steam and light ends (mixture of C1 -C4 gases available as a by-product of pyrolysis). Reasonable upper bound to the temperature of the pyrolysis fluidizing g'as is given by the temperature of the combustion reactor since higher temperatures could be reached only at the cost of burning extra fuel or part of the oil product. Results of calculations made around the pyrolysis reactor, as shown in Figures 5, 6, and 7, indicate that 450-500 kJ/kg tar sand are required to operate the reactor over the specified range of temperatures. Approximately 85% of all this energy corresponds to that required for heating the tar sand from ambient level to reactor temper­ature: thus, the range of energy requirement given is very much insensitive to changes in oil yield and coke yield. Variations of bitumen content and temperature of pyrolysis indicated a small, but not negligible, change in the total requirement at pyrolysis as shown in Figures 6 and An increment of approximately 10 kJ/kg tar sand corre­to each one % increase in the initial bitumen !>i 01 ~ <ll s:: ~ Ul or-! 4 2 Ul ~ rl 0 o ~1 0.01 t:l. Initial Bitumen Fraction Pyrolysis Temperature Oil Yield Coke Yield Fluidizing Gas 0.05 Fluidizing Gas Inlet Temperature (OC) 0.08 450 0C 70 wt % 20 wt % Light Ends 0.10 Fluidizing Gas/Tar Sand Flow Ratio 0.50 1.00 Fig. 5. Effect of the fluidizing gas condition over the total pyrolysis energy requirement for different reactor temperatures. w ...... Pyrolysis Temperature (oC) 6 475 450 - 425 4 Initial Bitumen Content = 8 wt % 2 Fluidizing Gas = Light Ends Temperature Gas ( Inlet) = T (pyrolysis) - 25 (K) Oil Yield = 70 wt % Coke Yield = 20 wt % (J) -r-! ~ 0 ~ ______________ ~ __ ~~~ __ ~~~ __ ~ ________ ~ __________ ~ __ ~~~~~~ M o 0.01 0.05 0.10 0.50 1.00 ~ :>t ~ Fluidizing Gas/Tar Sand Flow Ratio Fig. 6. Effect of the amount of fluidizing gas over the total pyrolysis energy requirement for different reactor temperatures. w (X) ~ cO +J bI "~I-J .~ +J ~ Q) S Q) ~ .,..~ ~ tJ1 Q) ~ ~ tn H Q) ~ fxl Ul o,-f U) ~ ,.....f 0 H ~ AI Initial Bitumen Fraction 6 0.12 4 Reactor Temperature = 450°C 2 Fluidizing Gas = Light Ends Temp~rature Gas ( Inlet) = 425°C Oil Yield = 70 wt % Coke Yield = 20 wt % 0 0.01 0.05 0.10 0.50 1.00 Fluidizing Gas/Tar Sand Flow Ratio Fig. 7. Effect of the amount of fluidizing gas over the total pyrolysis energy requirement for different initial bitumen contents. 40 content of the tar sand being processed or, for each 100e increase in the pyrolysis temperature. The influence of the fluidizing gas condition (temperature and flow ratio) are dependent on the fluidiz­ing gas to tar sand flow ratio, as shown in Figures 5, 6, 7. In particular, Figure 5 shows an important influ­ence of the fluidizing gas to tar sand flow ratio when the temperature of the fluidizing gas differs by more than lOOoK from the pyrolysis temperature, and the flow ratio is above 0.10. Both factors are important only in laboratory-scale units, where the fluidizing gas to tar sand flow ratio is high (depends on the inverse of the height of the bed), but this effect is negligible for industrial-size reactors where the flow ratio is decreased because of 20-30 ft-high beds compared to 1-3 ft-high laboratory units. Average sensitivity of the total energy require­ment of the pyrolysis reactor is calculated indirectly by the value of the partial derivative with respect to the variable in question. Typical values of other variables are used. Sensitivity results are ':BnLown in Table 3 as % of change in the total amount of ergy required in the pyrolysis reactor, when the vari­le indicated is changed by the amount shown. 41 ( , Table 3 Parameter sensitivity analysis(1). Parameter Parameter % Change Parameter Value Change of Q/F1 (2} Reactor temperature(3} 47SoC +5 OC +4.4 Bitumen content 12% -2% -4.9 Oil yield(4) 70% -5% -0.05 Ooke yie1d(4) 20% -5% +0.9 Feed temperature 25°C +5 OC -0.9 AHo (pyrolysis) 200 Btu/lb +20 Btu/lb +1.2 (1) Analysis made with F2/F1 = 0.70 and T(F2 } = 540°C (2)% Change = (New value - Old value}*(100/01d value) (3) Oil and coke yields assumed constants (4) Change is counterbalanced by change in gas yield 42 The Combustion Reactor All the energy required for the pyrolysis reactor can be obtained by burning the residual coked sand in the combustion reactor where air is used as a fluidizing gas. The following is a study of the effect of different design variables on the total amount of energy available from the combustion reactor. The most important independent variable that affects the total amount of energy produced in the combus­tion reactor is the initial bitumen content of the tar sand being processed. Therefore, this parameter is included as the main variable in these studies. Other parameters to be considered are: Coke yield. This parameter is predetermined by the extent of the pyrolysis reaction and fluctuates between 18% to 22% according to pyrolysis temperature, initial bitumen content, and solids retention time. The effect of this parameter and the initial bitumen factor on the total amount of energy available from the combustion reactor is shown in Figure 8. Combustion temperature. The range of this parameter is limited by the pyrolysis temperature at the lower end and by mechanical properties at the upper end. The effect of this parameter on the total energy available at combustion, as shown in Figure 9, is N I o lO~------------~----------~------------~ M X '0 s:: ctS en s... ctS +J 8 O'l ~ "- tj ~ '0 (l) :> o @ 2 p:: Q) ,Q o +J >t O'l Pyrolysis Temperature Combustion Temperature Air Inlet Temperature = 450°C = 550°C = 525°C Coke Yield from Pyrolysis (wt %) ~ 0 s:: ~--~--------~------------~----------~ ~ 0.00 0.05 0.10 0.15 Tar Sand Bitumen Fraction 43 rig. 8. Energy available from combustion of coked sand for various coke yields from pyrolysis. N I 0 ~ -X rso:: co fJ) S-I co +J 0'1 ~ "'- t-J ~ S-I 0 +J t) co (l) t:t: s:: 0 .r-! 4J fJ) ~ ~ 5 0 C) 5 0 S-I ~ ro (l) > 0 5 (l) t:t: (l) ~ 0 +J >t 0"1 S-I \l) c !::tl 10 8 6 4 2 0 Pyrolysis Temperature = 450 (oC) Air Inlet Temperature = T (Combustion) - 25 (oK) Coke Yield, @Pyrolysis = 20 (wt %) Combustion Temperature (oC) 0.00 0.05 0.1.0 Tar Sand Bitumen Fraction O.lS 9. Energy available from combustion of coked sand at various . reactor tempera~ures. 44 45 explained mostly by the change of enthalpy of the sand. 3. Air inlet temperature. This design variable has a small effect on the total energy available at combus-tion when compared with the previous two parameters, as shown in Figure 10. This small influence is explained by a small heat capacity for air, and a small flow rate of air relative to the sand flow rate. Finally, it is necessary to indicate that all the calculations have been made considering the stoichiometric amount of air relative to the total amount of coke pres-ant. The total amount of energy available at combustion will be decreased slightly if an excess of air is used . because of an increase in the amount of unreacting gases, and decreased strongly if a stoichiometric ratio less than unity is used because of unreacted coke. Transferring Energy from the Combustion Reactor to the pyrolysis Reactor To transfer the' energy required in the pyrolrsis from the combustion reactor, three possible alter-tives, shown in Figure 11, are studied here. Depending the alternative, the additional energy required at Combustion is plotted in Figures 12, 13, and 14. For a f lue of zero, energy is in balance in that the combustion eactor produces the exact amount of energy required at N I 10 o M ~ -ro ~ ~ Ul ~ ~ 8 t1\ "~t-:) -~ 6 4 2 Pyrolysis Temperature Combustion Temperature Coke Yield at Pyrolysis = 450 (OC) = 550 (oC) = 20 (wt %) P = T{combustion} - T{air, inlet} (oK) 0.00 0.05 0.10 0.15 Tar Sand Bit~men Fraction Energy available from combustion of coked sand. Effect of preheating fluidizing air. 46 Tar Sand Air Tar Sand Products Fluidizing Gas Flue Gas Air Heat Recovery Spent Sand Tar Sand Products Hot Sand Recycle Flue Gas Air Recovery Spent Sand Products Pipes Heat Recovery Spent Sand Fig. 11. Three different modes of transferring energy from combustion to pyrolysis. a.- Preheating of sluidizing gas. b.- Recycle of hot sand. c.- Using heat pipes. ~ o .~ +.I Ul ~ ~ 6 ~ 4 .... 1 ~o Or-! Jjx - roro Q)~ H cd 2 .~ Ul ~ tJiH (l)cd ~-IJ ~tJ\ tJ\~ H'-.. Q) t-) 0 ~~ ril-r- i cd ~ o .~ +.I .~ ro ro ~ -2 0.1 Pyrolysis Temperature = Coke Yield = Oil Yield = Initial Bitumen Fraction _In Balance 0.5 1.0 5.0 10.0 Fluidizing Gas/Tar Sand Flow Ratio Fig. 12. Energy balance for the combustion reactor. Energy transferred from combustion reactor to pyrolysis reactor by preheating the fluidizing gas. ~ 00 ~ o -M +J Ul ~ ~ o u 4 °~0 2 H~ ~ ~ '0- O)rcJ ~ ~ -.-I cd ~ (J) tJ1 0) ~ 0 f:t:::ttl +J :>t 0'0' ~~ 0)"-..... ~t-) tLl~ ~ ttl ~ -.°-I +J -.-I '0 '0 ~ -2 -4 1 In 5 Pyrolysis Temperature Coke Yield Oil Yield Initial Bitumen Fraction 0.06 0.08 0.10 0.12 0.14 10 Hot Sand Recycle Ratio = 4500 C = 20% = 70% 50 100 Fig. 13. Energy balance for the combustion reactor. Energy transferred from combustion reactor to pyrolysis reactor by recycle of hot sand. ~ ~ Pyrolysis Temperature Coke Yield Oil Yield Bitumen 500 = 450 0C = 20% = 70% 550 Combustion Temperature (oC) Energy balance for the combustion reactor. Energy transferred from combustion to pyrolysis by using heat pipes. 50 51 the pyrolysis reactor. A positive value indicates that additional energy must be released in the combustion reactor from another fuel source. A negative value indi-cates that an excess of energy is avajlable. 1. Preheating of fluidizing gas. This alternative considers the use of heat exchangers in the combustion reactor to transfer the required amount of energy from the combustion reactor to the fluidizing gas used in the pyrolysis reactor. Considering the amount of fluidizing gas used as an independent variable and the initial bitumen fraction as a parameter, Figure 12 shows an optimal energy design for fluidizing gas to tar sand flow ratio of 1.8 to 2.0, independent of the bitumen content. Also, energy self-sufficiency can be achieved when processing rich tar sand, with bitumen content of 11% or above and high combustion tempera-o tures above 600 C. A problem not considered here is . the design of a reactor capable of handling a very high fluidizing gas to tar sand flow ratio of 1.8-2.0 without increasing the pressure of the reactor much ! above atmospheric limits. of hot sand. As shown in Figure 11-b, a of hot sand can be used to transfer enough energy from the combustion reactor to the pyrolysis Hot sand recycle ratio defined as the ratio 52 of the hot sand recycled to the hot sand leaving the system is used as a design-independent variable together with the initial bitumen fraction. The results, shown in Figure 13, indicate an optimal hot sand recycle ratio of 15-20, independent of the ini­tial bitumen fraction of the tar sand being processed. Also, energy self-sufficiency is achieved if tar sand with a bitumen content higher than 8.5% is used. The clear advantage of this mode of transferring energy over preheating the fluidizing gas used in the pyroly­sis reactor is the reduced requirement of bitumen content to reach energy self-sufficiency. This advan­tage is counterbalanced by a huge increase in the size of both pyrolysis and combustion reactors necessary to maintain solid retention times. Use of heat pipes. In this design, energy self­~ ufficiency is a strong function of initial·bitumen and combustion temperature with a minimum initial bitumen content of 7.3 wt % if combustion temperature equals pyrolysis temperature, for which an infinite area of heat exchange would be required. Any increase in the combustion temperature pyrolysis temperature decreases the required ar~a of heat transfer and increases the minimum bitu­men content limitation. A reasonable minimum bitumen of 8.5 wt % gives a temperature driving force 53 o of approximately 110 F for heat transfer purposes. other values can be obtained from Figure 14. The major drawback in the use of heat pipes as heat trans-fer media, compared to the two modes of heat transfer mentioned previously, is that the technology is new and needs to be demonstrated on a large scale. In modeling all three modes of transferring an approach temperature of 15°C was used in every /. heat exchanger design: also, energy recovery from any ,'stream was considered only if the stream temperature is above lSOoC. With the exception of the first mode of tr'ansferring energy, the amount of fluidizing gas used in th~ pyrolysis reactor was kept at a low value, but above the minimum fluidizing amount. In order to compare thermodynamically the three modes of transferring energy from the combustion reactor to the pyrolysis reactor, lost work was computed using the basis previously described, considering an ambient temperature of 25°C as the lowest temperature heat sink of • , 0 Also, steam generat10n at 140 C was used possible. Two possible situations were consid-reference or comparison purposes. The first self-sufficiency on the all process, while the second reference basis consid-maximum thermodynamic efficiency measured as minimum ~ -work within the feasible range of the variables. 54 Table 4 shows that, in order to operate each mode of transferring energy, the bitumen content of the feed-stock must be greater than a minimum value, which depends on the operating mode selected. Thus, a tar sand with 11 wt % bitumen is required to operate an energy-balanced thermal process that transfers energy from the combustion reactor, at 8l0 oK, to the pyrolysis reactor, at 723 0K by reheating the fluidizing gas. The hot sand recycle and US~ of heat pipes modes requires the use of tar sands with 'at least 8.5 wt % bitumen in order to generate enough e'nergy in the combustion reactor. The amount of energy recovered from the hot spent sand depends mainly on oper-ating temperature of the combustion reactor. From this energy, enough work can be obtained to compress the gas and the air used to fluidize each reactor when using the ~ot sand recycle mode. Only part of the obtainable work . , is required to compress the fluidizing gas and air when using heat pipes. AQ,ditional work is necessary when operating under the first mode, i.e., preheating the ~luidizing gas. Finally, under conditions of an energy-balanced process, the use of heat pipes for transferring the combustion reactor to the pyrolysis reac-shows a clear thermodynamic advantage over the other modes presented. For each mode of transferring energy considered, is an operating condition that minimizes the total Table 4 Lost-work analysis at minimum energy self-sufficiency condition. ' En~rgy Transfer Preheating of Mode Fluidizing Gas pyrolysis temperature (K) 723 Bitumen fraction , (wt) (minimum value) 0.11 Combustion Hot sand recycle ratio Fluidizing gas/ t ,ar sand ratio 1 lost work (k,J /kg tar sand) 810 o 2.0 373 739 Hot Sand Recycle 723 0.085 750 18 0.35 317 539 55 Use of Heat Pipes 723 0.085 785 o 0.05 352 423 56 lost work under the constraint of an energy-balanced process. This operating condition corresponds to the maximum thermodynamic efficiency of the particular mode as discussed by de Nevers and Seader (1980). The maximum thermodynamic efficiency condition is obtained by relaxing some of the constraints imposed when searching for an energy-balanced process. This efficiency condition is found by increasing combustion operating temperature at the expense of higher minimum bitumen content for the first and second modes, and by decreasing operating combustion temperature at the expense of increased heat transfer area when using heat pipes. Operating conditions that simultaneously satisfy an energy-balanced process and maximum thermodynamic efficiency for each mode of energy transfer are shown in Table 5. This analysis is not limited to the three configu­rations already described. Material and ener~y balances used to simulate each mode of energy transport can be associated to other processes. For example, the firt mode of transferring energy presented here (preheating the pyrolysis fluidizing gas), is very closely related to the first process studied at the University of Utah, previous to the introduction of heat pipes, where the energy pro-uced in the combustion reactor \vould be transferred to the- pyrolysis reactor by fluidizing the pyrolysis reactor ith the hot combustion gases. Similarly, the second mode 57 Table 5 Lost-work analysis at maximum thermodynamic effici~ncy. Energy Transfer Preheating of Mode ~luidizing Gas Pyrolysis temperature (K) 723 Bitumen fraction (wt) (minimum value) 0.12 Combustion temperature (K) Hot sand recycle ratio Fluidizing gas/ tar sand ratio Energy recovered from hot sand (kJ/kg tar sand) Coke yield (%) Total lost work (kJ/kg tar sand) 942 o 0.70 520 20 70 670 Hot Sand Recycle 723 0.09 817 5.0 0.10 388 20 70 469 Use of Heat Pipes 723 0.073 723 o 0.05 291 20 70 372 58 of transferring energy (recycle of hot sand), is somewhat similar to the Lurgi-Ruhrgas process as described by Rammler (1970). Finally, the results presented in this chapter related to the use of the heat pipe can be gener­alized to some extent to other processes that use other media of transferring energy indirectly from combu.stion to pyrolysis. CHAPTER V DEVELOPMENT OF A DYNAMIC MODEL In order to understand and control the process developed at the University of Utah for thermal processing o,f tar sands, a mathematical model of the system was develope4. In the first step, the process was divided into five stages that were classified as heat exchanger or reactor type. Next, the conservation equations were derived for each type by writing balance equations over stationary vQlume elements following the Eulerian contin­uum , approach. When deriving the conservation equations, space-average properties were defined within a volume elex,nent. In particular, the particle .phase was consid­ered to be a continuously distributed phase provided that the reactor volume element considered was large compared the particle size. Because of the geometrical nature of the process modeled, cylindrical coordinates were used for the erivation of the conservation equations, and angular aymmetry was considered when simplifying the resulting tions. Then the differential conservation equations the assumption of complete mixing of particle phase in radial and axial directions and 60 complete mixing of the fluid phase in the radial direc~ tion. General schematic representations of the heat I exchanger and reactor stage types are shown in Figures 15 and 16. Both stage types exchange energy with a heat pipe and a wall that acts as a first-order lag between the stage and the surroundings. The wall also exchanges energy with a heating coil and walls from other stages. -Inside the heat exchanger stage, fluidizing gas exchanges with the solid flow by countercurrent contact. the reactor stage, the fluidizing gas exchanges with the solid hold-up of the stage. The primary concern of this model development was simulate the process thermal behavior: therefore, main effort was spent in developing the energy conservation equations. Mass conservation equations were developed as needed to formulate the energy conservation equations. Momentum equations were not developed here, but are easily implemented if the fluid dynamic behavior is to be Conservation Equations a Heat Exchanger stage Conservation of mass over a volume element nr s 2dz , the particle phase, is given by: frate of massl _ rrate ofl _ [ra te of 1 LaccumulationJ - Lmass inJ mass outJ • (8) Gas Out Solids In Heating Coil G. J ~ I Wall Stage j T, h • I ----=-r----- • ~ • Twlz+ dz T/vl z+dz • -------1------ T/vl z ----~----- . . ·T I \t I · . • . z . t•- -T- w-lz ----......-.. • , To Other Walls Solids Out Gas In Fig. 15. Heat exchanger stage representation. - z+dz z To Other Walls Solids In Gas Out- F. G. J J ---------. ----. -- . ,(T '~)~' (~' v) ~I z Solids Out Gas In Fig. 16. Reactor stage representation. z+dz z 63 The rate of particle mass accumulation within the volume element is nr s 2dz (a/at)(£p ps ). The rate of mass in is nr 2£ p v I , and the rate of mass out is s p s s z nrs2£ppsvslz+dZ. Hence, substituting terms from Equation (8), dividing by nr 2dz , and taking the limit as dz s approaches zero, (9) Similarly, the continuity equation for the fluid phase becomes = - -aaz [p g v g (I - £p) ] • (10) Conservation of energy over a volume A dz, for the w wall, is given by: The rate of energy accumulation within the volume is Awdz(a/at)(pwCPwTw). The rate of energy in is Lw d zh gw(Tg - Tw ) + dz(Pw / L) + Aw [ -k w (aT w/ az)JI z + ~pdz£wcr(THP4 Tw4 ). The rate of energy out is Lwdzhoo(Tw - Too) + Aw[-kw(3Tw/az)Jiz+dZ . 4 4 L~, dzEWcr{Tw - Ts ). Then, dividing by Awdz, and taking limit as dz approaches zero, 64 / . . LHP 4 4 + (A-w) £ w0 (TH P - T w) · (12) The rate of energy accumulation for the gas phase is given by Agdz(a/at)(PgCPgTg ). The rate of energy in is L s dzh gs (T s - T 9 ) + A9 [-k 9 (aT 9 /az)]1 z + LHpdzh gw (THP - T 9 ) + A (p H v )1 . The rate of energy out is 9 9 9 9 z LWdZhgw(Tg - Tw) + Ag[-kg<aTg/az)]lz+dZ :of Ag(PgHgvg)I~+dz. Then, dividing by AgdZ, and taking the limi,t as dz approaches zero, a2T L = k ~ - ~(p H v ) + (s)h (T - T ) 9 az2 az 9 9 g Ag gs s 9 (13) Similarly, the rate of energy accumulation for the phase is As dz(a/at)(p s Cp s Ts ). The rate of energy in As [-k s (aT S /az)]1 z + AS (p s Hs vs) 1 z + L s dz£ wo (Tw 4 - T s 4). rate of energy out is As[-ks(aTs/az)]lz+dz As(psHsVs)lz+dz + Lsdzhgw(Ts - Tg). Hence, dividing by and taking the limit as dz approaches zero, 65 LS 4 4 + (--)€ cr{T - T ) As w W s· (14) Further simplifications of the mass conservation equations can be made by neglecting the accumulation term in both particle and gas phases. Then Equations (9) and (lO) are reduced to: (15) (16) Energy conservation equations can also be simpli-fied by making different assumptions for each phase. In particular, the wall energy conservation equation can be simplified by assuming a uniform temperature and then ' j including the boundary conditions (energy exchange with other walls) into the resulting equation. Then, consider-ing constant average-space properties, integrating over stage length L, dividing by L wh ~Too' and rearranging: P T ( w ) (1 w Lw L h 00T0 0 + - ~0)0 (17) 66 Introducing the dimensionless variables, ex = TX/(Too - I), LW = (Lwhoot)/(AwpwCPw)' n w = P w/ (L wL h 00T 00 ), a-.] = Q] ./(Lw L h 00 T 00 ), z = Z/L, a = h gw /h00, r 0 = (€wOT003)/hoo' ~s = Ls/Lw' and ~HP = LHP/Lw in the above equation, the following form of the energy conservation for the wall is obtained: '" de w cr:r = w I + ~ 0 J [( e + 1) 4 - (e + 1) 4] dz • s 0 s w (18) In a similar manner, the emerging conservation for the gas phase can be simplified~by assuming constant average-space properties, no accumulation, and no axial conduction. Then, introducing the new dimensionless variables S = hgs/hoo' ~s = Ls/Lw' and Yg = (GjCPg)/(LLwhoo) into Equation (13), e can be obtained as a function of Z: 9 aSg = ez y~[s~s(es - 6gl + a~HP(8HP - egl + a(sw - 8g l ]. (19) Also, the energy conservation equation for the rtiC71e phase can be simplified by assuming no axial Then, defining the dimension-ess parameter Y = (F.Cp )/(LL h oo )' Equation (14) can be ~ 5 J S W - 67 reduced to the following form: ass ~ 4 4 ~-z = ~[a(8 - 8 ) + o[(8w + 1) - (8 + 1) JJ. a Ys g s s (20) Equations (18) to (20) represent the set of dimen-sionless energy conservation equations for one heat exchanger-type stage. In order to solve this system of equations, the output conditions form the adjacent stages are required: therefore, the set of equations resulting for a five-stage system must be solved simultaneously. Considering the temperature of the wall to be at a pseudo-steady state, Equations (19) and (20) can be inte-grated simultaneously by using a shooting algorithm. The integration of the proposed system of equations can be by use of the following integral functions: (21) ) Then, by replacing these two functions into Equa- (18), the following expression is obtained for 68 + ~ cF2[~(e (0) + 1)4 + ~(e (1) + 1)4 - (e + 1)4J • s ssw (23) All implicit integrations are solved by using the ' trapezoidal formulation given by (24) This method of integration was selected because of ite A-stability and small h 3 local truncation error (Lambert, 1973) • . . The 'Conservation Equations fQr the Reactor Stage Conservation of particle mass over a ~olume ele­nr s 2dz , for the particle phase, is given by: the generation term is a negative that corresponds gasification of bitumen. The rate of accumulation within the volume element Wt ' 2dz is nrs2dz(a/at)(spps)· The rate of mass in is s , wrs2~ppsvslz' and the rate of mass out is wrs2Eppsvslz+dZ· 69 (26) Similarly, the continuity equation for the fluid phase becomes ,}t [ (1 - E:p ) P g ] = - '}z [ (1 - E:p ) P g v g ] + E:p P s r xn • (27) The continuity equation for the reactor stage can be .developed from the general formulation given by Equa-tion (11). Then, for the wall, the energy accumulation within the volume element Awdz is Awdz(a/at)(pwCpwTw). The rate of energy in is L wd z(l - EP: )hg w (Tg -wT ) ~, L dZE: h (T - T ) + dZ(Pw/L)+ Aw[-kw( aTw/az)] Iw p sw s w z + Lw d ZE: wo(T s4 w- T 4 ). The rate of energy out is Lwdzhoo(Tw - Too) + Aw[-kw(aTw/az)]lz+dz. Then dividing by Awdz, and taking the limit as dz approaches z~ro, a2T p L a ( C T) w at Pw Pw w = k w_ _2W _ + ~~L+ (Aw) h0 0 (T0 0 - T w ) az L L + (~) (1 - E:p)hgw(Tg - Tw) + (A W ) E:phsw(Ts - T ) A w W W L 4 4 + (A~w) E: wa (T s - Tw ). (28) Similarly, the rate of accumulation for the gas within the volume element Ardz is The rate of energy in is 70 LHpdz(l - £p)hgW(THP - Tg> + Ar dZ[(6hgs £p)/dPJ(Ts - Tg } +A [( 1 - £ ) p H v J 1 + Ar (1 - £ ) [-k (aT / az) J 1 • The r p 9 9 9 z P 9 9 z rate of energy out is L dz(l - E )h(T - T ) w P gw 9 w + Ar[(l - £ )p H v JI +d + A [(1 - £ )[-k (aT /az)JJI +d • P 9 9 9 z z r 9 9 9 z z Assuming that the chemical reaction occurs at the particle surface, the rate of energy production in the gas phase is zero. Then, dividing by Ardz and taking the limit as dz approaches zero, 6h £ - _a_[ (1 - £ ) P H v J + g$ P (T - T ) az p 9 9 9 dp S 9 L + (~) (1 - £}h (T - T ) Ar p gw w 9 L + ( HP)(1 - £)h (THP - T ). A P gw 9 r (29) For the particle phase, the energy accumulation is by Ardz(a/at)[ps£pCPsTsJ. The rate of energy in is LHpdz£phsw(THP - Ts) + LHpdz£wcr(THp4 - Ts4) + A~(£ppsHsVs)lz + Ar £p[-ks (3Ts /az)Jl z · The rate of energy out is L dz£ h (T - Tw) + LwdZ£wcr(Ts4 - Tw 4 ) f W P sw s Ar dz[(6hgs £p)/dPJ(Ts - Tg) + Ar(£pPsHsVs>lz+dz ArEp[-ks(aTs/az)]lz+dz. The rate of energy production Theref.ore, after dividing by Adz, r the limit as dz approaches zero, 71 L + (AW r ) Ephsw (TW - T s ) + E p Ps r xn (- 8H r .) • (30) The continuity equation can be simplified by assuming complete mixing of the particle phase: then, by integrating Equation (26) along the z-axis, multiplying by the axial area of the reactor, assuming constant average properties, and defining M. = ~r 2LE P : J s p s dM. ~ = dt (31) Also, the continuity equation for the -fluid phase can be simplified by neglecting the accumulation term. Then, integrating over z and rearranging gives: (32) The energy conservation equations can also be by assuming different approximations for each see In particular, the wall energy conservation equa-be simplified by assuming a uniform temperature 72 along the z-axis and then including the energy exchanged with other walls into the resultingequ~tion. Hence, considering constant average-space properties, integrating over the stage length, and. assuming complete mixing of the particle phase: A 1 Cp d(Tw/T ( w W w) ( 00 P T L h dt = ~LL_ wh _ T_ +. (1 - TW) W 00 Woo 00 00 + h (I-E) L T - T h T - T gw P I 9 w{ dz) sw{ s w) hOT L + Ep~ T 00 00 00 00 E aT 3 T 4 _ T 4 Q ' . + (w 00) (s w ) + ] h T 4 LL h T · 00 w 00 co 00 Introducing the dimensionless groups (1 . - Ep)(hgw/h oo ) and n = Ep{hsw/h oo ) into the set already defined in the previous section, the following (33) 8imensionless form of the energy conservation equation is de W dl'w 1 = n{t} - 8w + aloI (e 9 - ew)dz + n{es - 8w) + 8[(e s + 1)4 - (8 w + 1)4J + qJ'. (34) In a similar manner, the energy conservation for the gas phase can be simplified by assuming average-space properties, and no accumulation or conduction. Then, introducing the new dimensionless 73 variable 8 1 = (6Ar h gs € p )/(Lw h ~dp), G = G(z}/GJ. +1 , and H = Hg/(CPT~} into Equation (22) and rearranging: yl [ 8 I ( e s - e g) + a. I (ew - e g) + ~HP a. I (e HP - e ) ] • 9 g (35) Also, the energy conservation equation for the particle phase can be simplified by assuming constant average-space properties, and no axial conduction. Then, defining the time parameters L j = (MjCPs}/(LLwh~), and g. = (r n}(-~H )/Cp T~, and the dimensionless function 1 x p s tees) = [(FH)j - (FH}j+lJ/(FjCPs)(T j - Ts )' the energy equation of the particle phase can be reduced to: d(L]. e s ) 1 4 4 dt = 8 I 0 f (eg - e s ) dz + ~HP 0 [ (eHP + 1) - (e s + 1) ] (36) The University of Utah process was modeled by a of five thermally-coupled stages as shown in Figure The first, third, and fifth stages were modeled as exchanger-type stages, while the second and fourth es were reactor-type stages. Heating is provided with Ta~ Sand Gas + Products T STAGE 1 Entrainment Pyrolysis T-{J STAGE 2 Pyrolysis Reactor Coked . Sand A3 STAGE ·3 Entrainmen Canbustion Spent Sand STAGE 4 Corrbustion Reactor STAGE 5 Heat Exchanger Pyrolysis Heat Section Pipe combustion Section 74 Unive~sity of Utah Process. Stage representation. 75 to stages three and four, and heat walls is considered relevant only between four and five. Stages one to four are thermally by means of a potassium heat pipe. Tar sand is fed to stage one where it exchanges the stage walls and the gas that leaves stage the pyrolysis reactor. Stage two is fed with tar coming from s~age one and a fluidizing gas. Energy for the pyrolysis reactor is obtained completely heat pipe. Oil vapor and noncondensibles pro­pyrolysis leave the pyrolysis with the fluidizing gas, entering stage is exchanged with tar sand feed. The product of the pyrolysis reaction, leaves reactor and is fed into the combustion third stage is modeled as heat exchanger type sand exchanges energy with the walls of this with the combustion gases leaving the combustion At stage four, the coke carried by the sand is acts as fluidizing gas also. eleased by the combustion of coke is used to reactor at a given temperature and de the energy required at the pyrolysis section by ring this energy to the heat pipe. The last stage to model energy exchange between the hot sand 76 leaving the combustion stage and the fluidizing air fed to the combustion section. This five-stage system can be modeled using the equations developed in the previous sections by modifying the set of equations according to each stage situation. Thus, some equations are further simplified if one of the energy transfer elements is not present at a particular stage. Specifically, Table 6 lists the dimensionless groups obtained from the conservation equations and shows when each group becomes zero or nonzero at each stage. The dynamic simulation of the University of Utah unit is carried out by the program SYSIM, presented in Appendix B. The results are discussed in Chapter VII. The Heat Pipe Model As mentioned previously, energy is transferred from the combustion section to the pyrolysis section by means of a potassium heat pipe, but no details were speci­fied about the operation of this device. Basically, a heat pipe consists of a closed tube containing a suitable amount of working fluid (often under vacuum), where the inner walls are lined with a wick. The working fluid is vaporized at one end of the pipe and condensed at the Flow of liquid from one end to the other is facilitated by the wick. A detailed description of the at pipe operation can be found in the literature (Chi, q. J z a. a.' 8 ' Table 6 Dimensionless parameter usage at each stage. Comments energy exchange between walls axial position gas-wall convection gas-wall convection with particle phase gas-solid conveciion with wall gas-solid convection with particle phase gas energy ' transport solid energy transport radiation coefficient Stage #(*) 1 2 345 x X X X X X X X x x x X x x X X x X X X X X X X X X X X X X X X t;HP heat pipe exchange area X X X X e x 1T W solid-gas exchange area X X X solid-wall convection X X temperature X X X X X power to the wall X X wall time X X X X X X means that dimensionless group is nonzero at the given stage. 77 Dunn and Reay, 1978): therefore, this section will details concerning the development of a model simulates the heat pipe behavior. 78 Since vapor pressure of a fluid increases rapidly temperature, the temperature gradient along the vapor e is very small when working in the operating range of working fluid and below flow saturation: then, the pipe can be modeled as isothermal. When the heat average temperature is outside the operating range, temperature distribution can be modeled as a conduct­fin. Because of the large amount of energy that can transferred by a heat pipe, response to any disturbance fast and any position-dependent variable inside the pipe can be neglected. Flow saturation limits the maximum energy flow heat pipe can transfer at a given operating temper­Upper limits to the heat· transport capability of a pipe are set by one of the following factors: vis­forces at low temperatures, entrainment and burn-out high temperatures, sonic and capillary limits at any A subroutine, HPSYM, was developed to compute the potassium heat pipe at a given _mpetature by using the semi-empirical design equations procedure given by Chi (1976), and Dunn and Reay A computer printout and a brief description of 79 is presented in Appendix C, and a simulation, shown 18, of the actual heat pipe used was obtained by the dimensions and characteristic parameters to The total energy flow exchanged between the heat and a particular stage can be obtained from analysis the cons'ervation operations developed in the previous + Thus, defining QHP(j) as the heat pipe releases to stage j, the expressions relate QHP(j) to the properties and of s .tage j: LjLHP'€WO(THP = 4 - T 4) W· J J L· J LHP . f h gw . (THP - T )dz, where j = 1, 3 ; 0 g. J J J (37) LjLHP . €:wo (THP = 4 - T 4) s ; J J L. J + LHP . f ( 1 - e: )h (T - T )dz 0 p. gw. HP g. ] J J + L.L HP ] . €: p. h sw. (THP J J J ] T s . ), where j = 2, 4 • J (38) The total energy flow transported by the heat pipe Equation (32) as: 4 = j~lQHP(j), such that QHP(j) > O. (39) en +J +J ct:S ~ ~ 0 ~ ~ 4J ct:S (i.) ::c E ::::s E .~ x ct:S :E:: 10 3 10 2 400 Limiting Factor 1. Viscous forces 2. Sonic flow 3. Entrainment 4. Capillarity 5. Burn-out 500 600 Temperature ( o C) 700 80 Fig. 18. Maximum predicted heat transfer flow for the University of Utah heat pipe. 81 The average operating temperature of the heat pipe obtained by solving the followig equations for Tav: (40) (41) The validity of the heat pipe equations, i . e . , to (41), is given by the satisfaction of QHP(total)/QHP(limit at THP ) < 0.2 in order have isothermal operation of the heat pipe. According Cbi (1976), at higher values of the ratio mentioned the temperature drop along the heat pipe is not must be considered. lid Flow lue-Controller del In order to maintain constant residence times in pyrolysis and combustion reactors, the University of thermal process uses a Foxboro Model 40 proportional­; to control the discharge rate of from each reactor. Pressure signals from top and of each reactor are read by pressure gauges and fed differential-pressure cells, which transmit signals, to the pressure drop across the beds, to the The control signals are fed to tanks with seconds of time constant and the output 82 nals of the tanks are used to drive valves that control the reactors as shown in Figure 19. To model this control system, a PI form of the equation is used. Then, defining the error as a of the total span of the reactor, the control ssure applied by the controller can be computed from following set of equations: e. = 1 Pc,i M~. - Mr ef where i = 2,4 ; Mmax 1 t = K. (e. + fe.dt), where i = 2,4. 1 1 Ic,i o 1 The control pressure P . is fed to a system C,l (42) (43) the tube net, a dumping tank, and the valve. is system was modeled as a first-order lag with a negli-constant. Simultaneously, the pressure of is first-order lag system is used to drive the valve. erefore, the functionality between the pressure of the lag and the flow through the valve must be order to model the mass dynamics of the reac- Also, the time constant of the first-order lag must experimentally evaluated. Hence, the relationship tween the control pressure and the discharge rate can be Feed Pyrolysis 10--_ .·. . ..B ..e.. d ., . · . . Coked Sand . . . Combustion 1--_ Bed · . . . . . " . Spent Sand ~O dp Cell PI Control ~O dp Cell PI Control 19. Schematic representation of the solids discharge control system. 83 84 dP . ,.S ,,l.. ( dSt,1 .) = p . - P ., where i = 2,4 C,l. S,l. (44) (45) Equations (42) to (45) are solved simultaneously the mass conversation equations by the subroutine , presented in Appendix D. This system of equations dynamic model, and it can be solved pendently of the energy dynamic model. The tuning of the control parameters K. and I . l. e,l. be done by first finding the critical constants of the Then, the Ziegler-Nichols criterion can be used K. and I .. 1 C,l Ener Model In order to study stability and control strategies current University of Utah thermal process, a of the previously developed equations was chosen the development of a simplified energy model. Reason-le assumptions were then made, keeping in mind the sibility of generalizing the conclusions of this study r application to a larger-size plant. The assumptions possible range of validity are: Two-stage reactor representation with reduced wall effects. This model considers the use of two reactor-type stages, as defined previously, and neglects the the heat-exchanger type stages. By neglecting wall effects, thi, model better simulates larg-e size uni ts rather than a small one., Flow and level of solids at, steady-state. In this case, a solid flow or a solids level fluctuating around an average value will satisfy this assumption period of the oscillation is smaller than the 85 thermal time constants characteristic of this process. This assumption is also satisfied if appropriate , control of the flow and level of solids is chosen. Negligible contribution of the fluidizing gases to the thermal dynamic behavior of the reactors. This assumption covers three possible cases: (1) the temperature of the fluidizing gas is similar to the te~p~rature of the fluidized bed with a consequent small exchange of energy between the solid phase and g,as phase inside the reactor; (2) the heat transfer coefficient and time of contact between solid and gas and, (3) the heat capacity and flow c~pacity of the solid phase inside the reactor. Radiative heat transfer proportional to the difference o~ temperature between the bodies involved. This simplification allows the inclusion of radiation effects into the convective heat transfer coefficient, linearizing the heat transfer equations. This 86 assumption is justified because, in the range of operating temperatures (450°C to 550°C), the radiative energy flow amounts to from 24% to 28% of the total energy flow rate exchanged between the two bodies (coked sand and steel walls). With the above assumptions, the governing energy tions for each reactor can be derived from Equation = 1, S' = 0, 0 included into the n, and 6 1,s = 0 as: l' d6 l t:HP n l' (y) 1 dt = (-y-) 1 (6HP - e1 ) - 61 + (~) y 1 (46) d6 t.:Hpn l' (T) _2 = (-y-) 2 (6HP - 6 2 ) + ( 6 - 62 ) + (..:..9.) y 2 dt 1 y 2 (47) subindex 1 indicates the pyrolysis reactor and subindex 2 indicates the combustion reactor. The temperature of the heat pipe can be obtained and w. = ~ (37) to (40) under the following expres- = r L ·L E h J'--2 , 4 ] HP p]. SWJ. l: L . LHP e: h • J'--2 , 4 J p]. swJ. (48) (49) (50) 87 Then, Equations (46) to (48) can be solved simul-with the inclusion of the following constraints: ~HP 2 = { ~HP 1 , ~HP 2 f(Qlimit) T HP > T ac t'l va t'l on THP < Tactivatione f, The solution of the system of Equations (46) to (51) was obtained with an HP1000 computer. The respective and parameter determination is given in Appendix The results of this simulation are presented in Chap- VII. Application of this model to the data obtained the University of Utah thermal unit can be done if losses are included into the balance equationse losses are usually important for a small laboratory The solution of this system of three ordinary fferential equations presents extraordinary difficulties use of the temperature dependence of some of the In particular, the value of the time constant pipe equation is temperature dependent and 'reases with an increase in temperature. On the other the time constants for the first two equations (46 50-to-lOO times or more higher than the time the heat pipel equation. Therefore, the system 88 The solution of the described system of equations integrating simultaneously the set of tions using the backward Euler method (see Lambert, 3), where each step is solved by using Newton's method convergence. Also, a variable-size integration step included in the solving algorithm in order to the changing value of the time constant of CHAPTER VI EQUIPMENT AND EXPERIMENTAL PROCEDURE This chapter presents a description of the experi­unit, data acquisition system, experi­and data management. The unit utilized modification of the unit fabricated, aIled, and operated by Jayakar (1979), Weeks (1977), Therefore, the information presented covers only those details where the units differ. details regarding dimensions and operating character­the unit can be found in the theses of Jayakar The data acquisition system consisted of: (1) ware: including an HP1000-F computer with different s of peripheral devices including a HP2631A printer, HP264SA CRT, and HP7906 hard-disk auxil­an HP22S0 measurement and control unit: software; including existing software for the HPIOOO-F r _ _ ,~~ter (part of the RTE-IVB operating system), and for HP22S0 interface (MCL/SO programming language), and tware specially developed to program the HP2250 inter­communications between both Also, special software was developed for data 90 preparation of plots using GRAPHICS 1000 The laboratory unit utilized in this study dif-in some respect from the proposed commercial plant, in Figure 4. Specifically, no heat exchangers were with the laboratory unit. However, a heat exchanger preheat the fluidizing air by cooling the flow leaving the combustion reactor. As is common in laboratory-scale equipment, no recycle of gaseous product was considered and the separation was carried by condensation with water as a A basic material flow diagram is presented in re 20, and a schematic representation of the labora-unit is shown in Figure 21. ion of the Laborator Unit As shown in Figure 21, the laboratory unit con­of a two-stage fluidized bed column thermally a potassium heat pipe, a tar sand feeding a product recovery section. Also, the unit an analog measurement and control section and a data acquisition and display section. Both reactor stages consist of a reaction zone structed from 2-inch, schedule 105, seamless pipe made 304 stainless steel with an entrainment zone con­ucted from 4-inch, schedule 105, seamless pipe made of Tar Sand Feed Gas Fluidizing Pyrolysis Product Synthetic Oil -- Gas Bed Recovery Dust Coked Sand I Combustion Air Combustion Gases Bed Spent Sand Fig. 20. Basic material diagram of the unit. ' ..' Py~~!~~~~ {,:r"~? Vent - Heat Pipe ~:::.~ ;~{: ~ Sand •. ~~ Dust Cooling Water Electrostatic Precipitator Condenser System Oil Oil Oil Oil Vent Oil Mist Eliminator 92 000 0 Stearn ________ N--=2, Air, CH4 Air ~ 0 00 Electrical Power Solid Control Valve Metering Valve 21. Process flow diagram of laboratory unit. 93 stainless steel. A 3/4-inch 00 by 7 ft potassium heat permits heat transfer from combustion reactor to lysis reactor when the operating temperatures are A l/2-inch flange made of 304 stainless both fluidized-bed reactions and simulta-support for the heat pipe and the stand that allows a flow of coked sand from the pyrolysis . a,C,;t;or to the combustion reactor. This discharge of reactor is controlled by pneumatically valves. A pair of Foxboro Model 40 proportional-pneumatic controllers provides the control signal solid flow valves by measuring the pressure drop each bed and comparing it to a previously estab-set~ point. Considering that the combustion gases longer used as fluidizing gas for the pyrolysis tor, the control of the amount of oxygen unreacted in combustion reactor is not necessary; then, the flow of r is estimated according to the amount of bitumen in the r 'sand utilized as feed and the expected production of the pyrolysis reaction. For all the experiments out in this study, utilizing clean sand or Sunny-ide tar sand, the total amount of air utilized was 0.68 lb/hr, which gives a superficial velocity of 0.38 ft/sec o an operating temperature of 550 c. This amount of air required for processing 4 lb/hr of tar sand with a bitumen content of 8 wt % and a coke yield of 18 wt % of This flow provides a superficial flow of 35 enough to fluidize the combustion temperature according to the data External sources of energy are provided by elec-wire wrapped around the combus- The total electrical power added to the reactor is manually regulated by changing the the heating wires. The only source of energy ilable to the pyrolysis reactor is the heat pipe: in order to reach the operating temperature temperature of the combustion reactor must be re~sed above 450°C to make the heat pipe operational. 94 shown in Figure 18 of the previous chapter, the maximum of energy through the heat pipe is below 200 watts an average heat pipe temperature below 400 ° C and energy flow for an average operating tempera- 350°C. Electriclal heating is used on the trq.inment section of the combustion reactor in order to nimize heat losses from the heat pipe and thus generate adiabatic zone in the middle section of the heat Electrical heating is also used on the entrainment tion of the pyrolysis reactor, the cyclone, and the ilter in order to eliminate condensation of the product Upstream of the product recovery section. 95 The pyrolysis reactor is fluidized with air when clean sand and nitrogen when feeding Sunnyside A superficial flow rate of air of 25 lb/hr ft2 2.5-3 times the minimum fluidization is used when operating the pyrolysis reactor at 450°C in order to optimize the transfer of energy heat pipe. Calibration of the gaseous flows is lished by a SINGER, DTM-115 meter from American Division, connected to the exit of the respective and operating the unit at ambient temperature. Details concerning the manufacturing of the dif-parts of the unit, dimensions, and operation char-excluding the data acquisition, are given in listed theses; therefore, the next section is to the detailed description of the data acquisi-control) system available. - As indicated in Figure 22, there are 16 thermocou-s (type-k, ungrounded, magnesium-insulated, stainless located throughout the laboratory unit. Thermocou-is located inside the high temperature filter and the temperature of the gaseous products leaving Thermocouple T2 is welded to the dust The temperature of the gases leaving the of the pyrolysis reactor is measured by Thermocouples T4 and TS are located the pyrolysis reactor, 18 inches and 8 inches above ,I [] Controller TS--+­T6-~~ Port power--, 11-2 -t.. .... T7 --+-...., Air To Product Recovery ITT""1===:::z- Filter Cyclone T2 PYROLYSIS REACTOR dp Cell • COMBUSTION d REACTOR Cell HEAT EXCHANGER , I I I I I- ---------------~ I T,6: Ambient Level Control Level Control ••• II .: Electrical Power ~ Sand Instrumentation diagram (computer excluded) • 96 97 distributor, respectively. The heat pipe temperature three thermocouples clamped onto the , T6 , T7 , and TS. Thermocouples T9 , TID and T12 are inside the combustion reactor at 30 inches, 15 and 44 inches above the distributor plate, respec­The temperature of the wall at the combustion is monitored by thermocouple TIl' and the tempera­the fluidizing air is measured by thermocouple T13 3 inches below the distributor plate. The temper­fluidizing gas used in the pyrolysis reactor by using thermocouple TIS located at the pyrolysis reactor, and below the distribu­Thermocouple T14 is used to monitor the temperature the discharged products from the cyclone used for pyrolysis fluidizing gas. Finally, ambient is given by thermocouple T16 . The tar sand feeding system consists of a feed a capacity of 60 lb of feed, a I-inch diameter rew conveyor, and a variable speed motor regulated by a type controller from u.s. Motors. The rota­the screw conveyor is detected by a micro­operates an LED lamp at the control panel. by the rotation of the screw Then, the rotation frequency can by measuring the period of the rotation. 98 In order to provide a means of studying the c response of the laboratory unit, characterized by ide range of time constants, digital data acquisition ipment was installed, calibrated, and operated. The ital system connsists of three sections, as shown in 23. In the first section, analog signals are from the laboratory unit and sent to the control for direct display. These signals include 16 ratures that can be displayed one at a time, 4 pres­e measurements, the flow of tar sand, and the average power applied to the different parts of the Manual and pneumatic control signals are sent from the control front panel. The second tion consists of an HP2250 intelligent interface unit re analog signals obtained from the control front panel and stored temporarily. This interface can programmed to make decisions according to the data also can be programmed to send control front panel on to the equipment. last section of the data acquisition system consists an HPIOOO-F computer and different types of storage and devices. The role of this third section is to store, and display the information digitized by HP2250 interface unit. Also, the HPIOOO-F computer is Tar Pyrolysis Bed Combustion Bed Air I II Heat Sand Front Panel Analog Input: t Therroocouples, Pressure Tr., Analog Filters. Others. - - Analog Output: Pressure Trd., SCR's. - Digital Input: switches. ..... Digi tal Output: Relays, Solenoid Valves. r-- - - I 'LABORATORY R00M f CRT - Display I Data, Run I RS-232 J computer....r - I - t - Measurerrent I I I , Microprocessor IEEE-48-8 HPlOOO - I COMPUTER , I Control I I - HP2250 · I INTERFACE I J I - t - DATA IDGGING: I Disc,Printer, I Plotter, Tape Fig. 23. Schematic representation of the equipment and computer. \.0 \.0 100 programming the HP2250 interface by running RAN IV or BASIC programs, which down-load to the interface unit. Both the HP1000-F computer and the unit accept real-time mUlti-tasking programming and work in a multi-user environment. The HP2250 interface is a real-time programmable designed for computer controlled automation tasks luding data acquisition and reduction, decision-making, execution of control algorithms that operate independ­of the controlling computer. This interface con-of an HP2l04 Processor Unit and a HP2251 Measurement Control Unit (actually, one HP2l04 Processor Unit can ge up to eight HP2251 Measurement and Control Units). HP2l04 Processor Unit consists of a card frame and er~l cards which control the operation of the HP2250 (1) receiving and compiling instruction MCL/50 language, (2) executing the set of ction-card commands generated, and (3) controlling the -488 protocol card (also HP-IB or GP-IB card) that a communication link between the HP2250 interface the HPIOOO computer. The HP2251 Measurement and Unit consists of a card frame and up to eight ction cards which provide capability of interfacing ifferent types of sensors and actuators. The capability for data acquisition and control of HP2250 interface unit depends directly on the type of 101 available in the HP2251 Measurement and At the present time, the following function s have been installed and tested: HP25501A l6-Channel High-Speed Analog Input. This card provides the basic analog input capability for HP2250 interface. It combines a high speed kHz) analog-to-digital converter with sample and amplifier and a programmable gain amplifier. The has 16 addressable input channels, which accept a the range of 0 to 10 volts and digitizes 14-bit resolution. HP25503A 32-Channel Low-Level Multiplexer. This card is used in combination with two HP25594A Thermocouple Reference Connectors to provide for 30 addressable channels of programmable-gain thermocouple inputs. Each Thermocouple Reference Connector provides a r~ference voltage for one to fifteen thermocouples. HP255l0A 4-Channel Analog Output. This card features ,four programmable 12-bit digital-to-analog converters. Output range is 0 to 10 volts with accuracy of 2.5 mV in unipolar mode or 4 mV in bipolar mode . • ' HP255l4A 16-Point Relay Output. This card provides 16 points of single-pole relay contacts, which are individually controlled. The relays are rated for volts AC at 3.0 A with operating time 102 The communication protocol used between the HPIOOO ter and the HP2250 interface is the IEEE-488 ard, a synchronous input-output interfacing that on the order of 1 megabit per second a distance of less than 4 meters between both (Stone, 1982). Because of a physical separation approximately 30 meters between the HPlOOO computer and the laboratory room where the HP2250 interface located, an HP37203L HP-IB Extender Unit has been The HP-IB Extender Unit allows communication using IEEE-488 standard for distances of to 1000 meters at a reduced transfer rate. In this rate has been reduced to a maximum of megabit per second. The HPlOOO computer used as controller for the interface unit is a 16-bit minicomputer with real­me, multi-user, multi-programming capabilities. The runs under the HP RTE-IVB operating system at a 1 megainstruction per second, and it is capable processing up to 200,000 single-precision floating int operations per second. High-performance software ailable includes floating point arithmetic, matrix rations, and trigonometric, logarithmic, and other ientific function in firmware making the HPIOOO computer for manipulation of large arrays of data. The is especially suited for simulation and control 103 The HP1000 computer manages the following resources, exclusive of the HP2250 interface HP7906 MR Hard Disk System. Approximately 19.6 M bytes are available in two disks, one removable and ~ne fixed. Both disks are characterized by an average time of 25 milliseconds and average transfer rate megabyte per second. This CRT (cathode-ray tube) t.erminal station has 12,288 by,tes of memory and two minicartridge tape units. Maximum transfer rate ' is 9GO char/sec. HP2631A Printer with a 128-character set and a speed of 180 characters per second with a 7x9 dot matrix Plotter. This device uses IEEE-488 standard to communicate with the computer, sharing the line the HP2250 interface unit. The available plot­area is 285x203 mm with an accuracy of 0.25 mm. RS-232C 8-Channel Multiplexer. This card a high-performance method for interfacing the computer to up to eight RS-232C compatible devices like CRT's or printers. Actually there are connected to this, one of which is located in the laboratory room to permit monitoring of Further information about capabilities and opera­the HP2250 interface unit and HP1000-F computer is in the following Hewlett-Packard manuals: No. 0-90011 HP2250 Measurement and Control Processor tem Introduction: No. 25580-90001 HP2250 Measurement Control Processor Programmer's Manual: No. 92068-90004 Programmer's Reference Manual: No. 92068-90002 Terminal User's Reference Manual: No. 92060-90023 FORTRAN IV Reference Manual: No. 92840-90001 Graphics Software User Manual: and No. 37203-90001 HP-IB Reference Manual. Detailed information concern­communication standards RS-232C and IEEE-488 is given by Stone (1982), and basic information about methods and procedures is explained in (1978). Handling and/or manipulation of data by digital iques can be divided into two areas, which generate of difficulties. The first area is the the data obtained on-line, which can of storage area, fast processing, high data transfer rates. For the unit studied, the 105 in this area is the large requirement of for the unprocessed data, which is generated of 1 megabyte per hour. The unit is character-for large response times for the reactors, in the 20 to 60 minutes, and short response time, of one minute, for the heat pipe and dynamic of the Therefore, the unit has to be run for long of time to observe transient behavior of the iors, usually from 12 to 48 hours, and the resolution to observe changes rated by 5-10 seconds or less. Data filtering and ta compression techniques are used on-line to reduce the age area required. Figure 24 shows the readings git~zed from thermocouple No. 11 compared to the same ta filtered by using an average type filter (Smith, The filtered data, shown in Figure 24, has been 3 0e higher than the origInal data in order to show It can be easily seen that a large amount of ta are gathered from this thermocouple or any of the The second area is related to f-line analysis, processing, and display of the data I Dedicated software has been developed for the and efficient use of the available resources. Figure 25 shows schematically the flow of data and software used to control the data manipulation or data when processing data on-line (25-a), or off-line TEMPERATURE vs TIME 2.4121~ __________________________________________________________________ ~ 2.38 l F,o.2S I N 0.15 I * 2.36 ~.*... 2.34 * FIL.TERED DATA 2.82 U" 2.3121 • C) w n V 111 n: :J t- 2.28 2.24 2.22· AVERAGE FIL.TER ... OF POINTS- 2m THERMOCOUPL.E ... 11 FIL.TER-FIL.TER+3.m 2.20~~~ __ ~~ __ ~~~ __ ~~ __ ~~~ __ ~~ __ ~~~ __ ~~ __ ~~~~~~~ • 6~ .7121 .8121 .90 1. r2Il2J 1. 1121 < n: w n. I W ~_ ______T_ _I_ M_ _E_ __( _M_ _I __N_ ._ _) _________________________________________________________~ Fig. 24. Performance of the average filter algorithm over data from thermocouple no. 11. ..... o 0"1 (MCL/50 Main Task) 1=I ... ...---·-tl r:i:~ I.(~: 1~ 1( ~t FI~I:~fum a.- (FM:;R) t-iINI-CARTRIDGE CRT ..... t (SELEC) (SELEC) DISC (FMGR) HPlOOO .~ FILE:DATAnn COOPUTER ..... DISC DISC (SELEC) FILE:DATAFO ornATA) FILE :DATAf'l -- (PIm'Y, TITLE) h.- Fig. 25. Data management. a.- Data acquisition flow chart~ b.- Data display flow chart. Software shown in parenthesis. 108 Management of the data acquisition resources is by the program TRACQ (listed in Appendix F). initializes the HP2250 Interface by download­a series of tasks (written in MCL/50 language) and by locating storage space on the disk named DATAnn (where interactively specified at the beginning of each Then, the program reads the digital data generated HP2250 Interface, filters and compresses the data, on the CRT once a minute, and es the information in the disk at a variable frequency 0.003 Hz to 0.20 Hz. Real-time branching have been implemented in this program by two of the unused thermocouple channels as digital (binary 0 or 1). This branching capability allows r real-time options, two used to display the informa­in two different formats, one to read from the CRT station, and the last option for the program smoothly before its normal end time. obtained from the equipment can be printed on­off- line. Management of the stored data is done off-line by of programs (listed in Appendix G), which allow the of the data in different resources. Thus, the SELEC is used to examine a specified data file, locate a particular time of interest, and at the CRT. Then, if a hard-copy 109 is desired, the time and channels of interest are ferred to a different file (DATAF~) for posterior The program SELEC also has provisions for the information contained in the file DATAnn an interactively defined device like the printer or a The data selected and stored in the file DATAF~ is by a series of programs before a final graph is The first step of the sequence is done by ng the program XDATA, which acts on the file DATAF~ generates the parameters required for the proper of the plotter. These parameters and the infor-be plotted is then stored in a new file named The second step in the plotting sequence is to program PLOTY, which takes all the information the file DATAFl and sends the required information to The last step in the sequence is then to run program TITLE and add alphanumeric information to the like title, subtitle, axis, and comments. The programs X DATA , PLOTY and TITLE have been from the algorithm developed by Julian (1981) by ncreasing speed and efficiency of the data processing I i nvolved, and also, by modifying the input/output made from interactive data input to disk file input and includ-a token passing security algorithm. The increased 110 is necessary when plotting more than 100 points: in ticular, for the input of more than 100 points. CHAPTER VII RESULTS AND DISCUSSION Results from this study are presented in four The first section presents dynamic data ned from the existing laboratory unit under different . The second section presents calcu­results from the two-s,tage energy model and compares ~esults to the data presented in the first section. , stability of the proposed process is studied by " the two-stage energy model, with the results pre-the third section. Finally, fluid and thermal simulation of the existing laboratory unit is out by using a five-stage model, with the results ented in the fourth section.- The experimental results presented here were ~ned during Runs 7 and 12 to 17. Experiments 1 to 6 e used to calibrate the equipment and to test the data isition system. Experiments 8 and 9 were used to conditions for using tar sand, and 10 and 11 were unsatisfactory attempts in to fluidize the pyrolysis reactor. The low of success when attempting to use steam as the 112 for the pyrolysis reactor can be explained adequate insulation in the piping system. The fluid-dynamic behavior of the system studied by an oscillating response, as shown in This cyclic signal is generated by the hyster­response from the valves controlling the flow Amplitude and frequency of the response depend the control parameters used in the existing controllers (such as proportional band and The high frequency, low amplitude 1 superimposed to the main signal shown in Figure 26 cates the existence of bubbling inside the reactors. ling has been observed during all the experiments, frequencies between I and 3 seconds and amplitudes depend on reactor temperature and fluidizing gas . 1 The dynamic data presented here were filtered using rat-order lag filter constants of 0.1 and 0.05 for the olysis and combustion data, respectively. The thermal dynamic behavior of the system is by the nonlinear behavior of the heat pipe. 32 show the temperature history of the olysis reactor, combustion reactor, fluidizing gas, and air for the different operating conditions tUllltmarized in Table 7. Before turning on the electrical 1\ H ill n. v 0.. o rr o • ill ill lJJ n: n. 2.ae~ ____________________________________________________________________ _ COMBUSTION REACTOR 1. 20 PYROLYSIS REACTOR a.a~~~~~~~~~~--~~~~~~~~~~~~~~ __ ~ __ ~~~~~~~~~ 0 .. 1210 .60 • 90 1. 20 1. 50 1. B0 2. 1121 2.40 2.70 TIME (MIN) Fig. 26. Reactor pressure drop history. (Data in file FILE91.) 7.00 .. 2. a -I- a. a • 1•2 . • 1•4 • • 1•2 . SOLIDS FLOW (LB.I'HR) .. N S.01lJ COMBUSTION REACTOR I ~** 5.00 ri PYROLYSIS REACTO * " 4.fd0 U • L1 3.00 fvl FLUIDIZING AIR W 2.00 n: ] r ( 1. 00 It: W n. ~ fd. fd" W 0."" • 30 • 60 • 9121 1. 2121 1. 5121 1. 8121 2. ua 2.4121 2.7121 ... TIME <MIN:'> *1121**-:3 Fig. 27. Reactor temperature history. Experiment number 7 • ~ ~ ~ 7.21m~ __________ ~ __ ~~ __________ ~~ ____________ ~~ __ ~~~~~~~ ____ ~ 4.0 0.0 4.0 SOLIDS FLOW N B.00 I ~** 5.21ra rl * 1\ 4. ara U • (l W 3. 021· o v w n: 2.210 (LB/HR) COMBUSTION REACTOR FLUIDIZING AIR J I­< lr W n. I W t-a. 0a~~~~~~ __ ~ __ ~~~~~.-____ ~ __ ~~~~~~~~ ____ ~ __ ~ __ ~~~ s. ae • 2m • 40 • Bm • era 1. rara 1. 2ra 1.4ra 1. Bra 1. era 2. Ii! a TIME (MIN) Fig. 28. Reactor temperature history. Experiment number 12. 7.S2I __ ----------------------------------------------------------------------~ COMBUSTION REACTOR 4.121121 FLLiIDIZING AIR FLUIDIZING GAS 0.121121 S.SB 1. 121121 2.21121 3.121121 4.121121 5.Sm 6.00 TIME (MIN) .... 1121 ........ -2 Fig. 29. Reactor temperature history. Experiment number 13. t-' t-' 0'1 COMBUSTION REACTOR (KWATTS) • COMBUSTION REACTOR PYROLYSIS REACTOR FLUIDIZING AIR FLUIDIZING GAS a.BI2I~~~~~~~~~~~~~~~~~~~~~~~~~~~~ __ ~~~~~ __ ~.~ 121.121121 • 221 • 421 • 6121 • 8121 1. 21121 1.2ra 1.4fa 1. 6121 1. afa TIME (MIN) Fig. 30. Reactor temperature history. Experiment number 14. 7.ma _•_ -----4-.- I-I -----_-j_- --------2-1.- a- ~~----------_-j-_- 4-. -21- --S~O-L(-LI~DB~S/~H ~RF~L)~O ~W- ---_-, COMBUSTION REACTOR PYROLYSIS REACTOR 2.a0 FLUIDIZING AIR 1. aa FLUIDIZING GAS ~~--------------~'--------________________ -J a.aa~~~~ __ ~~~~~~~~~~ __ ~~~~~~~~~~~~~~~ __ ~~ __ ~~ 121.121121 .2121 .4a • em . aa 1.121121 1.2121 1.4m 1. sa 1. alZl 2.aa T I ME eM I N) Fig. 31. Reactor temperature history. Experiment number 15. I--' I--' 00 e.eB~ ________________________________ ~ ______________ ~ ________ ~ ________ -. s. " -I- 4.121 SOLIDS FLOW (LB/HR) 112J. fa • . .. 7.B0 6.0B ~_ ______- COMBUSTION REACTOR \ PYROLYSIS REACTOR --- ~ 4. 00 3.a0 FLUIDIZING AIR ~~~~~~~ FLUIDIZING GAS 1. (21(21 2.(21(21 4.(210 5. (21(21 6.0B TIME <MIN) Fig. 32. Reactor temperature history. Experiment number 16. describing the operating conditions Experiment number 7 12 13 Disk file DATA52 DATA58 DATA62 Solid retention (lb) pyrolysis 2.4 1.3 1.8 combustion 2.7 4.0 4.0 Solid flow (lb/hr) 2.0 4.0 4.0 Mass of the reactor walls (lb) pyrolysis 30 25 25 combustion 30 30 30 Electrical power (kwatts) 2.0 2.2 2.2 (ft 2 ) I Heat transfer areas heat pipe, pyrolysis 0.30 0.20 0.20 heat pipe, combustion 0.38 0.56 0.56 external, pyrolysis 6.3 4.2 4.2 external, combustion 5.0 7.3 7.3 of the experimeht~ at start~up. 14 15 16 17 DATA71 DATA72 DATA75 DATA76 1.3 2.4 2.4 2.4 4.0 4.0 4.0 4.0 4.0 4.0 3.0 4.0 25 30 30 30 30 30 30 30 1.5 2.3 2.3 2.3 0.20 0.30 0.30 0.30 0.56 0.56 0.56 0.56 4.2 6.3 6.3 6.3 7.3 7.3 7.3 7.3 I-' N o 121 r to the combustion reactor, clean sand is fed into h reactors and a quasi-steady state is obtained. Then, er electrical power is applied to the combustion reac-time delay generated by the heat pipe is observed energy is transferred to the pyrolysis reactor. s time delay depends mainly on the total electrical r applied to the combustion reactor. Tar sand is fed the equipment only if the temperature of the pyrolysis ctor is above 440°C. It has been observed that the h of a run with tar sands depends on the temperature pyrolysis reactor for temperatures below 480 °C ; 0, the size of the particles fed to the reactor affects performance of the discharge standpipe in the pyroly-reactor exit. A tar sand particle size from 100 to microns seems to be suitable for the existing unit. is particle size range is determined by the ' inside ameter of the pyrolysis reactor standpipe and solid flow trol valve; thus, by increasing the size of these two larger particle size can be used to feed the reactor. Figures 33 to 35, expanded views of Figures 30 to , and Figure 36 show the dynamic response of the experi-apparatus when feeding sunnyside tar sand. A decrease in the pyrolysis temperature is observed feeding tar sand, and is caused by the endothermic ture of the pyrolysis of bitumen. The increase in 8.2121~ ____________ ~=-~~~=-________________ ~~~ ______ ~~_~~=-________ ~ CLEAN · SAND _I_ TAR _I_ NO FEED • 6. 21121 _ COMBUSTION REACTOR ---------------------------------~ 5. rae. PYROLYSIS REACTOR ---------------------~~ S.00~~~~~I~ __ ~ •• __ ~~~I~~~~I~ __ ~~~I~~~~I~ __ ~.I __ ~~~I~~~~ 1. 40 1. 42 1. 44 1. 46 1. 48 1. 521 1. 52 1. 54 1. 56 1.58 TIME (MIN:> Fig. 33. Expanded view of experiment number 14. I-' N N N I ** N rt * f\ U • (l W o v W II J I­< n: w n. I W f- TEMPERATURE VS 8.BBr-----------------------------------------------------________________ ~ CLEAN SAND TAR SAND NO FEED 7.012J COMBUSTION REACTOR ----------------------------~------------------~ PYROLYSIS REACTOR 4.121121 3 .. 0" FLUIDIZING AIR 2.121121 FLUIDIZING GAS 1. Be 0.B0~_,~~,~,~,~, __ ~,~,.~~~~ __ ~~~ ____ ~~ ____ ~~ ____ ~ __ ~ __ ~~ ____ ~~_ 3.5121 3.7121 3.9" 4. 1121 4.312J 4. 5" 4.7121 4.9121 5. 10 5.3" T I ME eM I N) Fig. 35. Expanded view of experiment number 16. ~---~~------ 2.aa FLUIDIZING GAS 1. aa e.0a~~~~~~~~~~~~~~~~~~~~~~~~ __ ~~~~~~~~~~~~ a."" . 3" .60 1. 2(21 1. 50 1. 8(21 2. 10 2.40 2.7m TIME (MIN:> *1121**-2 Fig. 36. Reactor temperature history. Experiment number 17. t-' N U1 126 ustion temperature, caused by burning the coke pro­at the pyrolysis reactor, is delayed by a time ·""""..·,,:rtional to the solid retention time in the pyrolysis The decreasing temperature observed after 50 with Sunnyside tar sand in experiment r 15 is due to plugging of the feeder with a conse­decrease in reactor bed height. When running tar sand, the feed rate is maintained lb/hr. A higher flow rate increases the probabil­plugging, and a lower flow rate increases the ficulty of measuring the equipment response to tar sand Also, the electrical power supply is maintained 2.2 kwatts in order to achieve operating tempera-in both reactors. Experimental operating conditions were changed a run in order to provide information about the t under different situations.- Clean sand was used when unit was operating at conditions different from the itions mentioned above. Nitrogen was used to fluidize pyrolysis reactor when operating with tar sand, and was used to reach steady-state and to operate at low or other flow rates. Independent operating iables were tar sand flow, electrical power used for combustion reactor, and amount of fluidizing gas used the pyrolysis reactor. 127 Typical sources of troubles observed during a run plugging of the tar sand feed hopper due to the lomerating nature of the Sunnyside feedstock and feed used, plugging of the flexitube connecting the pyrol­s entrainment zone and the high-temperature cyclone due low local temperature and the presence of fines in the t gas stream, plugging of the pyrolysis solids-dis­valve due to low reactor operating temperatures 400°C), and plugging of the combustion solids­charge valve due to the small size of the discharge All these problems can be eliminated by using a size reactor and feed rate, and by' slight modi fica­mechanical design of the feeding system for a In particular, a shorter length and larger conveyor should improve feeder performance. Thermal dynamic simulation of the existing unit, using the two-stage energy model, was carried out with set of equations derived in Chapter V and developed in Numerical integration was done, with the help an HPIOOO-F computer, by the program SIMUL listed in Heat transfer coefficients inside the fluid­beds were calculated using the correlation presented and Cooper (1958). 128 The thermal impact of using heat pipes to transfer from the combustion reactor to the pyrolysis reac-can be seen by comparing Figures 37 to 39. Parame-rs for these simulations are obtained from the operating itions of experiment 12. In order to evaluate the el performance, the data obtained from experiment 12 e plotted together with the curves generated by the Figure 37 shows a simulation that considers no miting effects of temperature on the behavior of the pipe. In this situation, the flow of energy between reactors is limited only by the values of the heat ansfer coefficients inside the fluidized beds. This gure indicates clearly the necessity of modeling the o at pipe energy flow at low temperatures (below 450 C). high temperatures, the energy flow transferred from the ustion reactor to the pyrolysis reactor is limited by heat transfer coefficients inside the fluidized beds, no heat pipe limiting effect is expected to be t