Kinetics of Propane Cracking Related to Its Use as a Heat-Transfer Fluid

Jan 16, 2015 - The overall rate constants were consistent with an extrapolation of results ... A combination of the propane kinetics with simple heat ...
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Kinetics of Propane Cracking Related to Its Use as a Heat-Transfer Fluid Alan K. Burnham,*,† Gregory J. Turk,†,‡ James R. McConaghy,§ and Leonard H. Switzer, III† †

American Shale Oil, LLC (AMSO), Rifle, Colorado 81650, United States Antero Engineering, LLC, Salida, Colorado 81201, United States

§

S Supporting Information *

ABSTRACT: The kinetics of propane cracking at high pressure were measured to evaluate its suitability as a heat-transfer fluid, in either a closed loop or directly injected into the formation, to retort oil shale in situ. Rate constants were measured in batch reactors at isothermal temperatures from 450 to 540 °C and at constant heating rates of 1.5 and 3.6 °C/min. Rate constants were also measured in a flow loop for isothermal temperatures ranging from 440 to 473 °C. The lowest temperatures in the batch autoclave experiments showed evidence of autocatalytic kinetic behavior, but the higher temperature batch experiments and the flow loop were more nearly first-order. The overall rate constants were consistent with an extrapolation of results from higher temperature measurements. Product selectivity changed as a function of conversion, with low conversion products rich in C4+ products and high conversion products predominantly methane. A combination of the propane kinetics with simple heat balance calculations shows that more than enough propane is supplied by the retorting operation to balance the consumption by cracking, making the use of propane for the heat-transfer fluid self-sustaining.

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INTRODUCTION A variety of true in situ oil shale retorting processes are being researched,1−3 including various methods of delivering or generating the heat of retorting down into the oil-shale formation. Heat delivery methods include resistive electrical circuits, closed-loop hot pipes immersed in a boiling oil pool, and hot gas injected into heating wells. The initial permeability of the oil shale is usually negligible; therefore, a base case is to heat the formation by thermal conduction from the heater wells. However, 20−30% porosity and associated permeability is generated by kerogen removal, which might provide for significant convective heat transfer into the formation. The advent of super-insulation materials4 makes generation of heat on the surface and injection underground more attractive than previously. To carry that heat to the retort in either a closed loop or direct injection, a heat-transfer fluid with both a high heat capacity and high thermal stability is desired. One proposal by EGL Resources5 initially used steam because of its high latent heat and then switched to an industrial heattransfer fluid, such as Therminol VP-1 (Solutia-Eastman) or Dowtherm A (Dow Chemical), to raise the formation to a pyrolysis temperature of 350 °C. Given that the driving force for heat deposition is proportional to how much hotter than 350 °C the heat transfer fluid is, it is highly advantageous to use a heat-transfer fluid above 450 °C, which is above the working temperature of all currently available industrial heating fluids (with the exception of molten salts). Although steam, N2, and CO2 are very stable, they have relatively low sensible molar heat capacities compared to light hydrocarbons. Ethane is about 50% greater; propane is twice as large; and butane is nearly 3 times higher. However, a preliminary screening indicated that the thermal stability of butane was too low;6 therefore, propane was selected as the best compromise between thermal stability and heat capacity. © XXXX American Chemical Society

The current work explored the suitability of propane for use as a heat-transfer fluid in this application. Although many papers have studied propane cracking for one reason or another,7,8 most are at higher temperature or lower pressure than for this application, and the most relevant study6 was at significantly higher temperatures. Given that the activation energy at elevated pressure may be higher than near atmospheric pressure, the available literature was insufficient for a reliable extrapolation to our conditions. In addition, the existing literature does not ordinarily cover the synthesis of products larger than propane because of alkylation reactions, which are a particular interest in our application. Also, the high-pressure and low-temperature conditions studied in this paper are relevant to petroleum geochemical applications.

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EXPERIMENTAL METHODS

Experiments were conducted in both a batch reactor and a continuous flow loop. Diagrams of the two apparatuses are given in the Supporting Information. The batch experiments were conducted in a model 4652 500 mL Parr pressure reactor with a 1500 W heater and controller. A 50 g charge of propane was added to the vessel, and it was then heated according to the desired thermal history. The thermal histories varied substantially in character to provide a rich data set for robust model calibration. Isothermal runs were conducted from 450 to 540 °C; stepped thermal histories were conducted at 480−430−480 °C and 510−455−510 °C (with the middle temperature held for ∼3 h); and constant heating rates were conducted at 1.5 and 3.6 °C/min. Temperatures were measured in the fluid, and exact thermal histories were recorded for use in the kinetic calculations. Pressures were Received: November 20, 2014 Revised: January 15, 2015

A

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Energy & Fuels typically about 10 MPa. Samples were withdrawn periodically and analyzed by gas chromatography to determine the amount and composition of the product gases. Quantities were adjusted to reflect the amount of sample withdrawn for analysis. Ordinarily, nominally pure propane (0.18% ethane and 0.02% butane) was used in the tests, but one experiment used commercial-grade HD-10 propane (1.9% ethane, 1.5% isobutene, and 0.5% n-butane) for comparison. The circulating flow loop experiments were conducted at temperatures ranging from 440 to 475 °C. Liquid propane was pumped at a controlled mass flow rate from a reservoir containing approximately 7 kg of nominally pure propane into the reaction train. The reservoir operated typically at 18 °C and 1.36 MPa, and the reaction pressure was typically 8−9 MPa. The first vessel in the reaction train was a preheater to convert propane into gaseous form, and the second vessel was a coil through a furnace to heat propane to the desired reaction temperature. The hot propane was then circulated at a rate of 9 kg/h through a stainless-steel pipe reactor (5 cm inner diameter × 1.6 m long) for a residence time of a couple minutes, which is similar to what is expected in practice. After exiting the pipe reactor, the propane reaction was quenched at pressure using a simple water cooler and then into a condenser and flash vessel at 80 °C and 1.5 MPa, which was designed to condense any significant quantities of C5+ products. Propane was then cooled again and flashed into the propane reservoir.

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no analytical solution is available for these equations, a solution is available for n and m equal to 1 (basic Prout−Tompkins11,12 limit)

α = 1/(e−kt (1 − α0)/α0 + 1)

where α0 is the integration constant that provides an initial fraction reacted to get the reaction started. The solution to eq 5 is a sigmoidal function in time. Some reactions are sigmoidal in logarithmic time rather than linear time, but one can obtain equally good fits using the linear-time sigmoidal model with an apparent reaction order of n ≈ 2 in eq 4. The circulating flow reaction data was first-order at the low conversions measured, and rate constants were determined by linear regression. The Parr reactor data showed acceleratory behavior at low temperatures; therefore, more sophisticated analysis methods were used. One method was Friedman’s isoconversional method, which involves applying a simple Arrhenius analysis at various fixed extents of conversion. This results in apparent first-order A and E values as a function of conversion for what is basically an infinitely sequential reaction, i.e., no competitive reactions that change the pathway of the reaction. A second method was fitting the parameters on the righthand version of eq 4 to the conversion data by nonlinear regression using Kinetics0513 to determine values of A, E, m, and n for various fixed values of q. Equation 4 was numerically integrated over the exact thermal history, which included a significant nonlinear heating phase.

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KINETIC ANALYSIS METHODS

RESULTS Batch Reactor. Thermal and pressure histories for four experiments are shown in Figure 1. The pressure varies

The temperature dependence of chemical reactions is ordinarily described well by the Arrhenius equation, where the rate constant k increases according to the formula

k = Ae−E / RT

(5)

(1)

where A is a frequency factor, because it has units of reciprocal time in addition to other units, depending upon the kind of reaction, E is the activation energy for the reaction, R is the universal gas constant, and T is the absolute temperature. If the reaction is first-order in the concentration or amount of reactant, the rate constant can be derived from a logarithmic plot of rate or fraction remaining as a function of time, and the frequency factor and activation energy are then derived from a plot of the logarithm of the rate constant versus 1/T. Hydrocarbon cracking is usually not described well by a single firstorder reaction over wide ranges of conversion. Instead, hydrocarbon cracking is governed by a free-radical chain reaction consisting of initiation, propagation, and termination reactions. Although a variety of detailed mechanisms exist, the objective of the current work is to derive relatively simple global kinetic expressions that are consistent with the concept of a chain reaction. The autocatalytic reaction formalism considers the reaction to be governed by two successive reactions: initiation and propagation9 A → B;

B + A → 2B

(2)

which can be described mathematically by the relation dA/dt = −k1A − k2AB. In this case, A would be propane and B would be products, including radicals that then attack other propane molecules. The differential equation cast in terms of fraction of the reactant (propane) that has been consumed, α, is

dα /dt = k1(1 − α)n1 + k 2α m(1 − α)n2

Figure 1. Thermal and pressure histories for four selected batch reactor experiments.

(3)

The empirical reaction orders are added for flexibility. In this case, n1 is the usual reaction order and m is a growth dimensionality, which might be considered analogous to a branching ratio for chain reactions. If we assume that n1 = n2, then

somewhat from run to run and over time for a given run; our kinetic analysis does not consider pressure explicitly. There is a significant time prior to temperature stabilization, which was accounted for in the kinetic analysis by numerical integration of the rate law over the measured thermal history. Preliminary inspection of the data found that the conversions of the commercial and pure propane samples were within experimental precision of each other. Some coke was deposited on

dα /dt = k 2(1 − α)n (α m + z) ≈ k 2(1 − α)n [1 − q(1 − α)]m (4) where z = k1/k2 ≈ 1 − q. The z and q factors are basically a way of starting the reaction in the absence of an explicit initiation reaction. Equation 4 is known as the extended Prout−Tompkins equation.10 A more detailed derivation is given in the Supporting Information. While B

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Energy & Fuels the reactor, but it was too small to quantify. A maximum conversion of about 95% was achieved at 540 °C. The isoconversional analysis can only be conducted up to the lowest conversions for all experiments being considered. Given that the 457 °C experiment achieved only 4% conversion, it was not included in some of the analyses. The isoconversional results varied with the selection of experiments included, but generally, the initial activation energy was about 250 kJ/mol and dropped to about the 170−210 kJ/mol range after several percent of conversion. The extended Prout−Tompkins method was more useful for analyzing the Parr reactor data. Again, however, the optimal kinetic parameters depended upon the experimental range analyzed. An example of the fit to all data is shown in Figure 2.

Figure 3. Fits to the entire temperature range and lower temperature range of the isothermal Parr data. The lines represent fits to the extended Prout−Thompkins model (eq 4).

Figure 2. Comparison of experimental and modeled data for propane cracking at various thermal histories. The lines represent fits to the extended Prout−Thompkins model (eq 4).

conversion ranges. The higher activation energy for the low temperature data is consistent with the higher isoconversional activation energy at low conversions mentioned above. It is possible that a full autocatalytic model might be able to reconcile this discrepancy by having a higher activation energy for the initiation reaction than the propagation reactions, as would be expected, but that possibility was not explored. The product selectivity at isothermal and stepped conditions changed with conversion, as shown in Figure 4 for some examples. Similar trends, except for the isobutane/n-butane ratio, were observed for ramped heating, as shown in the Supporting Information. Products larger than propane are formed primarily by alkylation reactions, with some contribution from termination reactions. The ratio of 1-butene/nbutane followed a trend similar to ethene/ethane. The ratio of 2-butene and isobutene to 1-butene initially increased with conversion, with the former becoming constant at higher conversions and the latter declining. This change in product selectivity emphasizes the point that the nature of the reaction or, more specifically, the mix of reactions changes with conversion; thus, there is no reason to expect the global reaction to be first-order or the apparent activation energy to be

The fits are equally good at isothermal, stepped heating, and linear heating, which shows the versatility of the model. There was negligible difference in the fits with q = 0.99 and 0.999, because the initiation−propagation ratio (1 − q) can be compensated by the growth dimensionality (m). The derived parameters for these two cases are given in Table 1. The large value of n is indicative of being sigmoidal in logarithmic rather than linear time. Further insights into the kinetics come from fits to the isothermal data over the entire range and the lower temperature range. Those results are shown in Table 1 and Figure 3. The quality of the fit for all isothermal experiments improves when not considering all thermal histories. This is probably more of a reflection of the limitations of the kinetic model than anything else. This conclusion becomes more evident when considering only the low conversion data at the lowest three temperatures. The fit to those experiments becomes evident, and the activation energy becomes more similar to literature values, which are normally derived over smaller temperature and Table 1. Kinetic Parameters Derived from the Parr Reactor Data case all data all data all isothermal 457−484 °C a

q (fixed)

A (s−1)

0.999 0.99 0.999 0.999

× × × ×

1.71 2.19 5.44 2.12

8

10 108 109 1014

E (kJ/mol)

m

n

residualsa

177.2 177.4 201.5 263.8

0.57 0.71 0.46 0.61

1.96 2.10 1.75 3.41

0.240 0.246 0.055 0.009

Normalized per point. C

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Figure 4. Examples of changes in product ratios as a function of conversion. Step 480 is the 480−430−480 °C sequence, and step 510 is the 510− 455−510 °C sequence. C2/C1 is the ratio of ethene plus ethane to methane; i-C4/n-C4 is the ratio of isobutane to n-butane; and (C4 + C5 + C6)/C1 is the sum of all C4−C6 hydrocarbons, both alkanes and alkenes, to methane.

independent of conversion. The corresponding issue is what range of conditions one can reliably use a simple global kinetic model for engineering purposes. Circulating Flow Loop. The circulating flow loop examines other aspects of the decomposition kinetics. The buildup of products was relatively small because of the large reservoir mixed in with every cycle. In practice, conversions were too low to yield a C5+ condensate, and the product concentrations in the propane reservoir never exceeded 5 wt % because of dilution by the large propane inventory. The reaction rate in the flow loop is most easily considered as a reactor with a constant inventory of reactant in which the products are continuously and selectively removed. The total fraction of propane cracked in the system ranged from 0.2% at 440 °C to 4% at 473 °C, and the system inventory dropped by as much as half because of sampling. The amount of reactant at each temperature was calculated from the reactor volume (2.9 L) and the propane densities, which ranged from 103 to 82 kg/ m3 for 440 to 473 °C. The ratio of cracking products to reactor content was roughly linear with time, as shown in Figure 5. The slope of the line at each temperature is the rate constant. An Arrhenius plot of the rate constants thus calculated is shown in Figure 6. This results in A = 8.55 × 1013 s−1 and E = 274.7 kJ/ mol. The A and E values from the circulating flow reactor are similar to the low-temperature values from the Parr reactor sans the nucleation characteristics. To eliminate the complication of the heating time in the Parr reactor, both sets of kinetic parameters were used to calculate the extent of reaction as a function of relevant times and temperatures, as shown in Figure 7. The first-order parameters are faster for the first few percent of reaction, and then the nucleation−growth parameters (e-PT) predict greater conversion. Product selectivities in the circulating flow loop were similar to those at low conversion in the Parr reactor. The average molar ratio of C4+ products to methane was 0.36; the average molar ratio of C2 products to methane was 0.41; and the

Figure 5. Generation of propane cracking products as a function of time at various temperatures with a more constant composition of reactants in the reaction vessel because of the flow circulation and large reservoir.

Figure 6. Arrhenius plot of the rate constants determined in the continuous flow loop reactor.

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that paper is at relatively high temperatures compared to our interest. Hautman et al.16 show a more complete review of old data along with their measurements, which might suggest that the activation energy increases as temperature decreases. There is also evidence, both experimental17,18 and calculated,14,19 that the activation energy of hydrocarbon cracking increases with pressure and that the role of pressure changes from acceleratory to deceleratory as the temperature declines from industrial cracking temperatures to natural petroleum formation and destruction temperatures. This change is due to shorter radical propagation chain lengths relative to initiation and termination reactions as well as increased alkylation reactions. This change is also consistent with the apparent crossing of reactivity of the atmospheric pressure expression by Hautman et al. with the data by Hepp and Frey. Consequently, a simple linear extrapolation of the atmospheric pressure kinetics predicts a few times faster cracking rate than what we observe at a high pressure. In our application, it is desirable to inject the heat-transfer fluid at a temperature substantially above the retorting temperature of about 350 °C to minimize the required flow rates and pumping costs. However, the degradation of organic heat-transfer fluids increases exponentially with temperature; thus, there is a trade-off between pumping costs and heattransfer fluid costs. Moreover, our intent is to supply all or most of the propane from our product stream, which means that the amount of propane decomposed to heat a certain volume of rock cannot significantly exceed the amount of propane that is generated from that rock. The trade-off between increased heat delivery and increased propane cracking is summarized in Figure 9. This simple,

Figure 7. Comparison of the reaction extent calculated as a function of time at various temperatures for the Parr reactor (e-PT) and continuous flow loop (first-order) kinetic parameters.

average molar ratio of ethylene to ethane was 2.7. This implies that, even though relatively high conversions were obtained in the reactor, the constant removal and dilution of products meant that selectivities were typical of low conversions. In fact, they were a little richer in C4+ and olefins than in the Parr reactor.

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DISCUSSION An initial autocatalytic phase as observed in the batch autoclave experiments is not usually seen in the hydrocarbon cracking literature, because most people find an apparent reaction order of 1.0−1.5. However, detailed mechanistic modeling of butane cracking14 saw such an acceleratory phase; therefore, it may be detectable only at a low temperature and high pressure, and the current experiments are unique in this characteristic. The first-order rate parameters from the flow loop experiments are easy to compare to others reported in the literature, as shown in Figure 8. The most relevant comparison is to Hepp and Frey,6 and an extrapolation of their results to our temperature range agrees well with our parameters. In a review on hydrocarbon cracking, Fabuss et al.15 state that the literature supports an activation energy in the 260−265 kJ/mol range. However, most hydrocarbon cracking work reviewed in

Figure 9. Fraction of propane generated from a given volume of rock that remains after heating the same volume of rock for various injection temperatures.

qualitative calculation involves only the mass fraction yield of propane from the oil shale, the heat capacities of the oil shale and propane, the residence time of propane for a single transit from surface to retort and back, the exit temperature of propane from the retort area, and the cracking rate expression. The propane yield from in situ retorting is about 5 g/kg of shale,20 and the heat capacity of propane at the relevant conditions is ∼3.5 kJ kg−1 K−1. For clay-rich oil shale relevant to AMSO,21 using enthalpy data from Camp,22 the energy required to retort the shale is estimated to be about 550 kJ/kg, including the energy needed to boil its water content and convert kerogen to gas and oil vapor. The transit time for a single pass through the system for a possible configuration is 2 min for a 1200 m long horizontal well and 900 m of overburden

Figure 8. Comparison of the flow loop first-order kinetics with previous work. E

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Energy & Fuels and surface piping to deliver and recover the heat-transfer fluid. If we assume that the exit temperature of the heat exchange fluid is 370 °C, at which the rate of cracking is negligible, the effective residence time for cracking for each loop is 1 min. From the relative heat capacities of propane and shale, the number of passes through the loop to heat a unit mass of shale ranges from 200 to 400 times, depending upon the input temperature. Figure 9 is then calculated using the first-order rate expression from the continuous flow loop experiment with the resulting total residence time for multiple loops. Details are given in the Supporting Information. The conclusion is that propane can be used effectively as a self-supplied heat-transfer fluid for temperatures up to 500 °C and possibly up to 550 °C, depending upon the details of the system and the heat exchange rate between the heater well and the formation. An uncertainty in this calculation relates to the differences in the Parr reactor and flow loop kinetics. Because a relatively small fraction of propane will be cracked over each loop, one could separate the cracking products from a slip stream of the appropriate magnitude and avoid the apparent acceleration of the cracking rate as heavy products build up. Such a strategy would also maximize the production of more valuable heavy products, which affects the economics of the total system.23

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American Chemical Society (ACS): Washington, D.C., 2010; ACS Symposium Series, Vol. 1032, Chapter 10, pp 185−216. (3) Burnham, A. K.; Day, R. L.; Hardy, M. P.; Wallman, P. H. AMSO’s novel approach to in-situ oil shale recovery. In Oil Shale: A Solution to the Liquid Fuel Dilemma; Ogunsola, O. I., Hartstein, A. M., Ogunsola, O., Eds.; American Chemical Society (ACS): Washington, D.C., 2010; ACS Symposium Series, Vol. 1032, Chapter 8, pp 149− 160. (4) Examples are Pyrogel XT from Aspen Aerogel, Northborough, MA; Izoflex and Microtherm thermal insulation from Microtherm, Inc., Sint-Niklass, Belgium; and Nanogel aerogel matting from Cabot Corp., Boston, MA. (5) Harris, H. G.; Lerwick, P.; Vawter, R. G. In situ method and system for extraction of oil from shale. U.S. Patent 7,743,826 B2, June 29, 2010. (6) Hepp, H. J.; Frey, R. E. Pyrolysis of propane and butanes at elevated pressure. Ind. Eng. Chem. 1953, 45, 410−415. (7) Volkan, A. G.; April, G. C. Survey of propane pyrolysis literature. Ind. Eng. Chem. Process Des. Dev. 1977, 16, 429−436. (8) Perrin, D.; Martin, R. The hetero−homogeneous pyrolysis of propane, in the presence or in the absence of dihydrogen, and the measurement of uptake coefficients of hydrogen atoms. Int. J. Chem. Kinet. 2000, 32, 340−364. (9) Burnham, A. K.; Weese, R. K.; Wemhoff, A. P.; Maienschein, J. L. A historical and current perspective on predicting thermal cookoff behavior. J. Therm. Anal. Calorim. 2007, 89, 407−415. (10) Burnham, A. K. Application of the Sestak−Berggren equation to organic and inorganic material of practical interest. J. Therm. Anal. Calorim. 2000, 60, 895−908. (11) Prout, E. G.; Tompkins, F. C. The thermal decomposition of potassium permanganate. Trans. Faraday Soc. 1944, 40, 488. (12) Brown, M. E. The Prout−Tompkins rate equation in solid-state kinetics. Thermochim. Acta 1997, 300, 93−106. (13) Burnham, A. K.; Braun, R. L. Global kinetic analysis of complex materials. Energy Fuels 1999, 13, 1−22. (14) Mallinson, R. G.; Braun, R. L.; Westbrook, C. K.; Burnham, A. K. Detailed chemical kinetic study of the role of pressure in butane pyrolysis. Ind. Eng. Chem. Res. 1992, 31, 37−45. (15) Fabuss, B. M.; Smith, J. O.; Satterfield, C. N. Thermal cracking of pure saturated hydrocarbons. In Advances in Petroleum Chemistry and Refining; Kobe, K. A., McKetta, J. J., Jr., Eds.; Wiley Interscience: New York, 1964; Vol. 9, Chapter 4, pp 157−201. (16) Hautman, D. J.; Santoro, R. J.; Dryer, F. L.; Glassman, I. An overall and detailed kinetic study of the pyrolysis of propane. Int. J. Chem. Kinet. 1981, 13, 149−172. (17) Dominé, F. Kinetics of hexane pyrolysis at very high pressures. 1. Experimental study. Energy Fuels 1989, 3, 89−96. (18) Jackson, K. J.; Burnham, A. K.; Braun, R. L.; Knauss, K. G. Temperature and pressure dependence of n-hexadecane cracking. Org. Geochem. 1995, 23, 941−953. (19) Dominé, F.; Marquire, P. M.; Muller, C.; Côme, G. M. Kinetics of hexane pyrolysis at very high pressures. Energy Fuels 1990, 4, 2−10. (20) Burnham, A. K.; McConaghy, J. R. Semi-open pyrolysis of oil shale from the Garden Gulch Member of the Green River Formation. Energy Fuels 2014, 28, 7426−7439. (21) Reeder, S. L.; Kleinberg, R. L.; Vissapragada, B.; Machlus, M.; Herron, M. M.; Burnham, A. K.; Allix, P. A multi-measurement corelog integration for advanced formation evaluation of oil shale formations: A Green River case study. Petrophysics 2013, 54, 258−273. (22) Camp, D. W. Oil shale heat-capacity relations and heats of pyrolysis and dehydration. Proceedings of the 20th Oil Shale Symposium; Colorado School of Mines Press: Golden, CO, 1987; pp 130−144. (23) McConaghy, J. R.; Burnham, A. K. Using liquefied petroleum gas in a hot circulating fluid heater for in-situ oil shale retorting. International Patent Application WO 2014/127045 A1, Aug 21, 2014.

ASSOCIATED CONTENT

S Supporting Information *

Schematic diagram of the batch autoclave system (Figure S1), schematic diagram of the circulating flow loop apparatus (Figure S2), product selectivity ratios for ramped heating experiments (Figure S3), background on the autocatalytic reaction equations, and details of the propane consumption calculation. This material is available free of charge via the Internet at http://pubs.acs.org.

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AUTHOR INFORMATION

Corresponding Author

*E-mail: [email protected]. Present Address ‡

Gregory J. Turk: Siluria Technologies, 409 Illinois Street, San Francisco, California 94158, United States. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors appreciate the assistance of Jeremy Huffman, Peter Steves, and the AMSO operations staff for constructing and operating the equipment. The authors also thank Genie Energy (Claude Pupkin, President of its AMSO subsidiary) and Total S.A. (Pierre Allix, Program Manager) for financial support and permission to publish this work.

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REFERENCES

(1) Ryan, R. C.; Fowler, T. D.; Beer, G. L.; Nair, V. Shell’s in situ conversion processFrom laboratory to field pilots. In Oil Shale: A Solution to the Liquid Fuel Dilemma; Ogunsola, O. I., Hartstein, A. M., Ogunsola, O., Eds.; American Chemical Society (ACS): Washington, D.C., 2010; ACS Symposium Series, Vol. 1032, Chapter 9, pp 161− 183. (2) Symington, W. A.; Kaminsky, R. D.; Meurer, W. P.; Otten, G. A.; Thomas, M. M.; Yeakel, J. D. ExxonMobil’s Electrofrac process for in situ oil shale conversion. In Oil Shale: A Solution to the Liquid Fuel Dilemma; Ogunsola, O. I., Hartstein, A. M., Ogunsola, O., Eds.; F

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