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Society ot Petroleum Engineers

SPE 30295Experimental and Numerical Simulations of a Novel Top Down In-SituCombustion ProcessR. Coates, S. Lorimer', J. Ivory, Alberta Research Council'SPE MemberCopyright 1995, Sooiety of Petroleum Engineers, Inc,This paper was prepared for presentation at the International Heavy Oil Symposium held in Calgary, Alberta, Canada, 19-21 June1995,Thispaperwas selectedfor presentation by an SPE ProgramCommitteefollowing review ofinformationcontained inan abstractsubmittedby the author(s), Contentsof the paper, as presented,have notbeen reviewed by the Society of Petroleum Engineers and are subjected to correction by the author(s), The material, as presented, does notnecessarily reflect any position oftheSociety of Petroleum Engineers, its officers, or members, Papers presented at SPE meetingsare subject to publication review by Editorial Committees ofthe Societyof Petroleum Engineers,Permission to copy is restricted to an abstract of not more than 300 words, illustrations may not be copied, The abstractshould contain conspicuous acknowledgment ofwhere and by whomthe paper is presented, Write Librarian, SPE, P,O, Box 833836, Richardson, TX 75083-3836, U,S.A. (Facsimile 214-952-9435),

ABSTRACTAnew in-situ combustion strategy, the top down process, is currentlyunder detailed laboratory study. The process, aimed at overcomingsome of the problems that have restricted the successful applicationof in-situ combustion in oil sand and heavy oil formations, involvesthe stable propagation of a combustion front from the top to thebottom of a reservoir, exploiting gravity drainage of the mobilized oilto a lower horizontal well.Operational parameters that have been investigated and presentedhere include: air injection flux, degree of pre-heating, internal steamflood pre-heating and injection of normal air versus injection ofoxygen enriched air.To compliment the experimental investigation, the thermal numericalsimulator STARS has been applied to the in-situ combustion processby incorporating reaction kinetics for Athabasca oil sand. Asuccessful history match of an experimental test is presentedaccompanied by a discussion of application of the model to fieldscale.INTRODUCTIONIn-situ combustion has long been recognized as having the potentialfor being an economical thermal oil recovery process in heavy oil andoil sand deposits. The energy required to supply heat to the reservoircompares quite favourably with steam. The estimated cost! to place

References and illustrations at end of paper

1 GJ of energy in a 7 MPa reservoir is $2.6-$4.4 using steam and$1.0 for in-situ combustion using air (assuming $2/GJ fuel cost,capital cost not included). In-situ combustion is not compromised bylarge heat losses to overburden and underburden in thin formationsor by high heat losses from the well bore to the overburden in deepformations as is the case with steam injection. Also in-situcombustion theoretically has important applications in reservoirscontaining bottom water and as a follow up process to waterfloodedand steamflooded forrnations2Previous in-situ combustion field projects, however, have been lesssuccessful than steam, primarily because of the difficulty incontrolling the combustion front advancement. The customary in-situcombustion operation of the past involved the injection of an oxygencontaining gas into a central vertical injection well surrounded by anumber of vertical production wells (typically as part of a largerpattern of injection and production wells). Combustion was initiatednear the injection well and horizontally propagated radially outwards,aiming to drive the mobilized oil towards the production wells. Theproblem frequently encountered was that the combustion frontstended to advance erratically with the vertical sweep constrained bygravity override of the displacing gas and the areal sweep reduced bypreferential flow to one well of the pattern. Injected oxygen,overriding the combustion zone, created problems at the productionend and the overriding hot steam and combustion gases did little toheat the formation ahead of the burn zone. The displacementgeometry of the process requires that the mobilized oil be displacedahead of the combustion front into the colder immobile oil,increasing oil saturation and further reducing mobility , with thelimited producibility of the vertical production wells unable toalleviate the situation.

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2 EXPERIMENTAL AND NUMERICAL SIMULATIONS OFA NOVEL TOP DOWN IN-SITU COMBUSTION PROCESS SPE 3029Incorporating horizontal wells into an in-situ combustion operationcreates new prospects. Horizontal wells have higher producibilitythan vertical wells and should not restrict the displacement of the oilahead of the front, providing improved sweep efficiency. Thispremise was convincinglydemonstrated in the laboratory by Greaveset al3,4 and field tests5,6 with horizontalwells placed in existing in-situcombustion operations, have shown promising results.Some of the most successful applications of in-situ combustion havebeen in the Suplacau de Barcau7,8 field ofRomania, in reservoirs thatcontain a slanted orientation. The combustion front is initiatedthrough gas injection at the top of the slant and the front advancesdown slope. Production wells, placed at the lower end of the incline,collect the mobilized oil. Adaptation of this burning down concept tooil sand and heavy oil deposits is the premise behind the top downprocess assessed in this paper.The conceptual strategy of the top down in-situ combustion processinvolves the stable propagation of a high temperature combustionfront from the top to the bottom of a heavy oil or oil sand reservoir.Combustion is initialized and maintained by injection of an oxygencontaining gas at the top of the reservoir, with mobilized oil drainingto a lower horizontal producer well. Most of the injected oxygen isconsumed in the high temperature combustion reactions at thecombustion front. Oxygen that passes unreacted through the front,reacts in lower temperature reactions to produce a layer of cokewhich is subsequentlyburned as the combustion front moves through.Hot combustion gases and thermally cracked light ends mix with theoil ahead of the high temperature front, heating, upgrading anddriving the oil by a top down gas drive. Gravitational forces helpdrain the oil to the horizontal producer.Although the top down process holds great promise, there arepotential problems which must be addressed before it can beconsidered for the field. The high bitumen saturation and viscosity ofvirgin heavy oil reservoirs must be overcome to obtain initialinjectivity. Methods to obtain injectivity need to be assessed as doesthe ability to successfully apply the process to reservoirs that havealready been partially depleted by a previous recovery operation. Thestable advancement of the combustion front through the reservoir andthe efficient draining of the mobilized oil to the producing wells bothneed to be proven.

3-D EXPERIMENTSThe top down in-situ combustion process is currently underinvestigation at the Alberta Research Council using a combination of3D physical simulation in combination with numerical modelling.Experience gained from previous 1D combustion experiments andnumericalmodel predictions were used to design the 3D experimentsResults from the seven 3D experiments are presented in this paper.

Experimental SetupSchematics of the experimental set up and cylindrical test packs ashown in Figures 1 and 2. The cylindrical oil sand pack, 29 cmdiameter and 40 cm in height, was prepared by manually tampinAthabasca oil sand into a stainless steel sheath. Each end of the pawas sealed by welding a stainless steel end cap to the sheath. All tesexcept Test 2, incorporated a water soaked central frac sancommunications path extending from the top to the bottom of tpack. Table 1presents details of each test pack.Thermal measurements within the packwere obtainedwith an arrof thermocouple rods inserted throughout the pack. A serieselectrical guard heaters, cemented in place around the stainless stesleeve, reduced heat losses during the experiments. The electricheaters were also used in Tests 1 to 4 to pre-heat the pacMeasurements from the thermocouple array were displayed in retime on a computer screen as a colour animation, enabling diredetermination of the progress of the experiment.Injection fluids (water, steam and gases), at a controlled rate atemperature, were injected through a slot (0.6cm long x 0.3 cwide) representing a vertical injection well, in the middle of the tend cap. An electrical igniter was located next to the gas injectipoint.Production was from the bottom through a 22 cm long, 0.8 cm wiscreened slot, representing the horizontal well. The produced fluflowed through a back pressure regulator maintained at 1.9 MPbefore entering a low pressure (0.3 MPA) separator. Gas from tseparator was cooled to remove any condensable components athen flowed through an on-line gas chromatograph for analysLiquid from the separator was collected in jars and analyzed f%bitumen, '%water, and %solids. Properties such as asphaltene asaturate concentrations, and bitumen density and viscosity wedetermined for selected samples. The pH of produced water samplwas also measured.Beginning in Test 7 a controlled flow of cooling water was circulatthrough the production well to maintain the produced fluids belo150 cC. This was to keep the produced fluids below the lotemperature oxidation range, a condition which had previously causpremature plugging of the production well.The sealed oil sand pack was enclosed within a steel pressure vessthat was filled with an insulating material. Nitrogen gas pressurapplied to the annular region to simulate overburden pressure, wmaintained at a constant 1.0 MPa above the injection pressure byregUlating/relief system. The pressure vessel was heatedsurrounding steam coils and additional thermocouples attached to tinside and outside of the pressure vessel aided heat loss estimatio

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SPE 30295 R. COATES, S. LORIMER, J. IVORY 3Experimental ProcedureAll the tests had a period of pre-heating prior to commencing airinjection. Tests 1 through 4 were heated by the external electricheaters and Tests 5 through 7 were pre-heated by injection of steaminto the pack. During the pre-heating phase of Tests 1 to 4, heliumgas was passed through the pack to maintain open communication.Once air injection was started, the electrical igniter was turned on atapproximately 180watts and turned off when a temperature of 400Cwas observed in the pack. This was considered as the start of ignition.Tests 1 through 5 had normal air (21% 02) injection and Tests 6 and7 had oxygen enriched (50%00 air injection. The injection flow ratewas continually adjusted depending on state of the combustion frontand the composition of produced gases. Monitoring of the computeranimation of temperatures indicated if combustion was decreasingand more oxygen was necessary. Experiments were continued untileither the pressure drop became prohibitively high (plugging) oroxygen composition ofproduced gases became high (gas channellingand combustion front extinguishing).A summary of the ultimate operating strategy for each test ispresented in Table 2.

EXPERIMENTAL RESULTSTables 3, 4 and 5 summarize the experimental results obtianed fromthe tests. Details of the results are discussed in the following text.The degree offormation pre-heating, required to obtain rapid ignitionwas investigated in Tests 1 to 4 by varying the amount of heatsupplied by the external guard heaters. Figure 3 displays the generaltrend obtained by the investigation, with the degree of pre-heatingrepresentedby the average temperature of the top 10 cm of the pack.The time required to obtain ignition increased dramatically as thetemperature decreased, with an average temperature near 200 cCrequired to produce rapid ignition. Note that the data point displayedfor Test 3 is an estimation. In the actual test, when ignition was notattained after several hours, air injection was stopped and pre-heatingwas continued. Several attempts were made to recommence airinjection. However ignition was not obtained until an averagetemperature of about 200 cC was obtained.The air injection flux is an important operational parameter in the topdown process. The flux must be sufficient to supply the oxygenrequired to sustain combustion. However too high a flux mayjeopardize the stabilizing effect of gravity which is desired in the topdown process. Test 1 used an air (normal, 21%oxygen) injection fluxof 5 std. m3/m2/h. At this injection flux, ignition was achieved, butafter approximately4 hours the combustion front began to extinguishas shown in temperature contours of Test 1 at 1hours, 4 hours and 8hours after commencing air injection (Figure 4). Air is injected intothe formation through a slot and as the combustion front expandsoutward radially, the injection flux needs to be continuously increased

to supply sufficient oxygen to sustain the combustion. For theexperimental configuration the outward progress of the combustionfront is limited by the radius of the pack and it was determined insubsequent tests using normal air, that a flux of approximately 10 std.m3/m2/h was required. In the field, the continual increase of therequired injection flux will be more pronounced as the combustionzone expands radially. The outward limit of the combustion zone isa subject of interest for the numerical model and will ultimatelydetermine the vertical interwell distance.The importance of supplying a communications path to assist ininitial injectivity was assessed, in Test 2, by air injection into a preheated pack that had no communications path. Air injection fluxstarted at 10 std. m3/m 2/h and although ignition was obtained, thepressure drop across the pack increased to an extremely high valueof2.4MPa.. The high pressure drop, caused by the creation of an oilbank ahead offue combustion front, ultimately caused fue combustionfront to become unstable and channel through to the production well,(Figure 5). The presence of the channel enabled injected oxygen topass unreacted through to the production well. High oxygenconcentrations combined with temperatures near 200C, caused lowtemperature oxidation of the oil and resulted in plugging theproduction well.Figure 6 presents the final recoveries obtained from the externalelectrically preheated tests. Test 1 had the lowest recovery becauseof insufficient oxygen supply. The absence of a communication pafuin Test 2 resulted in significant oil (11 %) being produced during thepre-heat which contributes to a high final recovery. Test 3 exhibitsa very slow production rate because of the low pack temperaturethroughout the test. The significant increase in production rate uponco-injection of water in Test 3, at approximately 17 hours, shows thatthis mode of operation may have important potential to improverecovery efficiencies. Although having initial slow oil productionrates, because of a low initial temperature, Test 4 produced the bestfinal oil recovery during the in-situ combustion period. Thecombination of suitable injection flux and communication pathcreated conditions giving a fairly stable advancement of thecombustion front and drainage of the mobilized oil. Co-injection ofwater in Test 4 did not show the large jump in production rate as inTest 3. This mode of operation may only show an advantage whencommenced in a formation that has not been significantly depleted.Figure 7 presents the viscosities measured for selected productionsamples from fue tests. It is apparent that the experimental top downcombustion process produces an oil with a viscosity reduced by abouta factor of four from the native oil. Note that steam injection duringfue pre-heatperiod of Tests 5, 6 and 7 created an initial produced oilof much higher viscosity than the native oil. However this alsoeventually decreased as the combustion period progressed. Analysisof the produced fluids indicated that the reduced viscosities wererelated to decreased asphaltene and increased saturate content.The purpose of Tests 5 to 7 was to examine combustion achieved in

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4 EXPERIMENTAL AND NUMERICAL SIMULATIONS OFA NOVEL TOP DOWN IN-SITU COMBUSTION PROCESS SPE 302a reservoir that has been previously depleted and pre-heated by asteam process. Steam injection was continued into the packs of Tests5 to 7 until the average temperature of the top IOcm reachedapproximately2QO C. At that time air injection was commenced andthe igniter turned on. Time to attain ignition is longer with steam thanwith the external electricalpre-heating(Figure 3). This is likely dueto the increased water saturation caused by the steam injection.Despite significant depletion from steam injection (oil recoveriesprior to air injection were 6%, 25% and 13% for Tests 5, 6, and 7respectively), enough fuel remained to sustain combustion. Athabascaoil sand contains an initial solid hydrocarbon component (up to 1.8%of total oil sand weight in the oil sand used, given in Table I as initialcoke) which may remain behind after a steam process. This initialsolid material can account for as much as 50% of the fuel reported forcombustion of Athabasca oil sand. Lighter oils may not contain thiscomponent and may not exhibit the same behaviour.To minimize the pressure drop during the steam pre-heat it wasdecided to increase the permeability of the communication path forthese tests. The permeability of the communications path used in Test5 , 135 11m2, proved to be too high. Heating during the steaminjection was prirnarilyin the region closely surrounding the path andonce combustion was achieved, the front moved rapidly through thecommunications path to the production well. Figure 8 displays atemperature profile of the Test 5 pack, showing the high temperaturezone reaching the production well within 2 hours after commencingair injection.Lower permeability communication paths, 19.5 11m2, used in Tests6 and 7, produced improved areal heating and did not result in thefront rapidly advancing through to the production well. However,asthe experiments progressed, the combustion front in each test,ultimately preferentially advanced down one side of the pack (Figure9). Pre-heating of the pack by steam injection appeared to promotethe channelling of the combustion front. This is evidenced in thesteam tests by the higher percentages of injected oxygen passingunreacted through the pack (Table 4).Despite channelling and extensive oil depletion during the steam preheating period, additional recoveries, shown in the insert of Figure10, were obtained during the combustion period of the Tests 6 and 7,that are comparable to Test 4, the best of the externally pre-heatedtests. Additionally the ratio of cummulative injected oxygen tobitumen produced during combustion also approach similar values,shown in Figure 11 for Tests 4 and 7. Starting ratios differ becausethe oil is already being produced at the commencement of airinjection in the steamcase whereas little oil is being produced in theexternal electric case.Tests 6 and 7 also utilized enriched (50% O2) air to obtain a loweroverall injection flux (4 std. m3/m2/h ) but at an equivalent oxygeninjection flux used in previous tests. The use of enriched air had nodetectable effect on the oil production.However, use of enriched air

showed evidence of greater oxygen consumption in low temperatoxidation (LTO) reactions. The apparent hydrogen/carbon (Hratios, which are an indication of the degree of LTO and shouldapproximately 1.5 for the burning of Athabasca bitumen, wslightly higher during combustion for Tests 6 and 7 than other teValues of 1.1-1.7 (Table 4), were obtained for normal air tests avalues of2.4-2.7 were obtained for enriched air tests. High valuesapparent HlC ratios were also observed for all tests prior to ignitand when the unreacted oxygen had channelled through toproduction well.For all the tests reported in this paper the viability of the top doprocess is threatened by the eventual channelling of the combustfront through to the production well. Channelling can be clearly sin the temperature profiles as well as the post run pack profile, shoin Figure 12 for Test 7. The degree of channelling was affectedoverall injection gas flux, the permeability of communication pand by depletion by previous steam injection. Channelling of the frthrough to the production well results in large amounts of the it*coxygen passing unreacted through the pack, leading to uneconomO/bitumen ratios and potential problems at the production end.If the mobilized bitumen does not efficiently drain away fromcombustion zone, stable advancement of the combustion front wnot occur, negating the conceptual advantage of the proceEvidence that the oil is not being displaced from the burn z o ~ 1encountered in the excessively high (compared to published valuefuel loads obtained in the experiments (Table 1). Literature valquote the fuel loads in the range of 24-48 kg fueVm 3 reservhowever experimental values were greater than 100 kg/Additionally, amaterial balance of the oil and coke, shown in Ta6 for Test 7, indicates that a sizable amount of initial oil is beconverted to coke and being burned, in this case 19% of the origibitumen was converted to coke of which 75% was burned. Simvalues were observed in other tests. The large amount of oil beconsumed in the combustion process further leads to uneconom02/bitumen ratios. Because the experimental packs were compoof actual Athabasca oil sand and the height of the pack was onlycm, capillary forces within the pack were comparable togravitational forces. It is hypothesised that this equivalencyresponsible for the mobilized oil not efficiently draining away frthe combustion zone, ultimately leading to the channelling. Infield, where gravitational forces are more dominant, it is anticipathat with proper operating strategy, channelling will not be caufrom this situation. To test the hypothesis, future experimesimulations are planned where the gravitational forces will be scato be dominate over the capillary forces.

NUMERICAL SIMUIJATIQNNumerical simulations of the experimental tests were performusing the CGM model STARS by incorporating a reaction model

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SPE 30295 R. COATES, S. LORIMER,' 1. IVORY 5in-situ combustion of Athabasca bitumen compiled by Belgrave etal. lO The reaction model is a consolidation of individual kineticstudies on thermal cracking l !, low temperature oxidation (LTO)12 andhigh temperature oxidation (HTO) of coke13 and uses a pseudocomponent approach to characterize the oil. In the model thefollowing reactions are proposed:

Thermal Cracking :(1) Maltenes .... 0.327 Asphaltenes(2) Asphaltenes 83.223 Coke(3) Asphaltenes 37.683 Gas

LTD:(4) Maltenes + 3.439 2.... 0.473 Asphaltenes(5) Asphaltenes + 7.5882.... 101 .723 Coke

HTO:(6) 0.811 Coke + 02.... 0.811 Gas + 0.46H20

Stoichiometric coefficients for the above reactions are determinedfrom estimates of the molecular weights of the pseudo components(Table 7).The rate of material conversion in the above reactions can be writtenas:

where: Ct =Concentration of reactant IC2= Concentration of reactant 2 ( if it exists)k = Rate constantn,m =reaction ordersWhen oxygen is a reactant, its concentration and reaction order aredefined in terms of the oxygen partial pressure. The rate constant, k,which is a function of temperature, follows Arrhenius Law:

where: A, =Arrhenius constant (reaction frequency factor)E =Activation energyR = Universal gas constantT =TemperatureValues for the parameters n, m, A, and E, were derived in theindividual experimental investigations and consolidated by Belgrave.The values quoted are reproduced in Table 7, converted to unitsappropriate for input into STARS. Also given in Table 7 are heats of

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reaction for each reaction, as well as gas/liquid k values, componenheat capacities and component densities in units required for inpuinto STARS.The reaction model suggested by Belgrave recognized that chemicachanges in the oil would results in changes to its density andviscosity. Changes in the density were obtained by inputting differendensities for each pseudo oil component and changes in the oiviscosity were derived within STARS based on the mixing rule:

Individual pseudo component viscosities were input into STARS aa temperature-viscosity table (Table 6), and STARS interpolatedwithin the table according to the exponential correlation:!!j= a exp bf f

Additionally it was recognized that the formation of the solid cokecomponent in the pore spaces would produce a restriction to fluidflow. This plugging mechanism was accommodated in STARS by thcalculation:

R= 1+ rrstx(C coke- so!idmin )where: rrst =plugging factor =.00374

Ccoke =Coke concentration (mole/m3pore volume)solidmin =Minimum coke concentration for plugging toccur = 5350 mole/m3 pore volume

The mobility of each flowing fluid is reduced by the factor R. Valuefor rrst and solidmin were obtained from history matching earlier airsteam experiments (unpublished).For the numerical simulation, the cylindrical packwas approximatedby aCartesian geometry shown in Figure 13. This geometry producedamodel having a similar external surface area to the physical modela requirement that was found necessary to closely match heat lossesA constant pressure producer runs through a bottom central row oblocks and aconstant volume injector is located in a top centre blockThe igniter is modeled by constant addition of heat to the injectionblock and the communications path is characterised by adjusting thpermeability of the central column of grid blocks.Results of the history match of Test 5 are presented in this paperHistory matches of other tests have been previously reportedl4 Tes5 had a period ofsteam injection prior to commencing air which ialso included in the numerical simulation. Initial conditions for thsimulation were:

Absolute permeability=0.5 11m2

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6 EXPERIMENTAL ANDNUMERICAL SIMULATIONSOFA NOVELTOP DOWN IN-SITU COMBUSTION PROCESS SPE 3029Water saturation=0.15Oil saturation = 0.85Initial temperature = 33 DCInitial coke concentration=2200gmole/m3 pore vol.Pressure=2100 kPa 'Communication Path Permeability:Test 5 = 135 J.Ull2

Figures 14 to 17 summarized results of the history match. Recoveryand pressure drop predictions are very comparable. Predictions of thetemperature profile are not as good. Although the simulation displaysthe rapid channelling down the communications path observed in theexperiment it does not duplicate the progression ofa hightemperature zone through the pack. Fine tuning of thecharacterization of the communications path in the numerical model(relative permeabilities and capilary pressure data) could provide abetter match. The outcome of the numerical simulations proved to bequite sensitive to these parameters.Formation of coke' ahead of the combustion front, which is animportant aspect of the reaction model, can be observed in Figure 17.Areas ofhigh contour density in the figure indicate regions of highcoke density. The maximum coke density is on the order of 25000gmole/m3 of pore volume. These regions of high coke density occurahead front of the advancing high temperature front, a behaviourwhich is also evidenced in post run pack appearance, Figure 12,where coke bands appear along the edge of the burnt channel.Discussion of Field Scale Simulation

The purpose of history matching experimental tests is so that anacceptable reaction scheme and other operational parameters can beobtained for the field scale simulations which will ultimately predictprocess performance in the field. However the results obtained fromhistory matching experimental simulations cannot be used directly forfield scale simulation. The size of grid blocks used in theexperimental history matching are comparable to the actual size ofthe combustion front, however field scale grid blocks can beseveralmagnitudes larger. In the numerical model, temperature is the averagetemperature over the entire grid block. It may adequately representthe temperature in the reaction zone in the experimental case but isinappropriate to represent the peak combustion zone temperature inthe field case. Because the kinetic reaction model is temperature andtime dependant, application of it in situations with large grid blockswill not produce reliable results.The challenge for the researcher is to apply a technique that canassess the controlling mechanisms of in-situ combustion withoutbeing impaired by fine grid requirements and numerical dispersion.Several solutions have been offered that consist of either, simplifyingthe reaction model 15 , modifying the kinetic parameters16 or alteringthe grid block temperatures I? The ideal combustion simulator wouldincorporate amethod ofdynamic localized grid refinement that would

track the advancing combustion front. Although this technique habeen introduced l819 it has yet to be developed to a stagewhere it cabe used in a commercial simulator.

CONCLUSIONS1. Employing a suitable air injection flux'is essential to obtaininggravity stable combustion front. The injection flux must bsufficient to supply enough oxygen to initiate and sustacombustion but not too high as to lead to channelling. As thcombustion zone grows radially, the injection flux must bprogressively increased until outward advancement has reacheits limit.2. The time to reach ignition after commencing air injection is highdependent on the degree of pre-heating. I f pre-heatingaccomplished by steam injection, ignition time will take longbecause of the increased water saturation in the formation.

3. After a steam process in an Athabasca formation, there is enougfuel remaining to initiate and sustain an in-situ combustioprocess. Additional recoveries achievable by in-situ combustioin a formation partially depleted by a steam process, wecomparable to recovery by in-situ combustion process in aundepleted formation, however the partial depletion of the pacduring the steam pre-heat promoted channelling of thcombustion front through to the production well.4. Wet combustion shows the potential to increase production rate

it is commenced before pack has become depleted.5. The combustion front can be maintained at a lower overall gflux with the use of oxygen (50%) enriched air with no effect obitumen production. However the use of enriched air showevidence of increased oxygen consumption in low temperatuoxidation reactions.6. The top down in-situ combustion process produces bitumen wia lower viscosity than the original native bitumen.7. Provision of a vertical communication path aids in establishininitial injectivitry, however selecting the proper path permeabili

is of utmost importance, as a combustion front can rapidadvance along a high permeabilitypath.8. Use of an actual oil sand in the experimental pack, resultedcapillary forces that were equivalent to the gravitational forcand as a consequence, mobilized oil did not drain away from thcombustion front. This condition produced high fuel loads,excessive amount of fuel being burned and eventually unstabadvancement of the front.

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SPE 30295ACKNOWLEDGEMENTS

R. COATES, S. LORIMER, 1. IVORYJournal of Petroleum Technology, pp 1154-1164, Oct., 1962.

7

The authors wish to thank members of the ADOE (formallyAOSTRA)-ARC-CANMET-Industry Research program forpermission to present this material. In addition, they wish to thankGurdeep Purewal and Kerry Scott for their assistance in conductingthe experiments.

REFERENCES1. Butler, R. M.: Thermal Recovery of OIL and Bitumen, PrenticeHall inc., pp 416-417, 1991.

2. Farouq Ali, S. M.: "Redeeming Features of In Situ Combustion,"presented at DOEINIPER, Symposium on In Situ CombustionPractices-Past, Present and Future, Tulsa, OK, USA, April 2122,1994.

3. Greaves, M., Tuwill, A A, Field, R. W.: "Horizontal ProducerWells In In Situ Combustion (ISC) Processes," presented atCIMIAOSTRA Technical Conference, Banff, Alta, Can. April21-24,1991.

4. AI-Shamali, 0., Greaves, M.: "In Situ Combustion (ISC)Processes: Enhanced Oil Recovery Using Horizontal Wells,"Chemical Engineering Research and Design (UK), TransIChernE, Vol. 71, Part A, May, 1993.

5. Morgan, R. J.: "Can Horizontal Wells Inject Life Into an OldCombustion Project?," presented at 5th PetroleumConference ofthe South Saskatchewan Section, Regina, Sask., Can., oct. 18-20,1993.

6. Ames, B. G., Grams, R. E., Pebdani, F. N.: "Improved SweepEfficiencyThough the application Of Horizontal well TechnologyIn Mature Combustion Project, Battrum Field, Saskatchewan,Canada," presented at SPE/CIM/CANMET InternationalConference on Recent Advances in Horizontal Well Applications,Calgary, Alta., Can., March 20-23, 1994.

7. Aldea, G., Turta, A, Zamfir, M.: "The In Situ Combustion andIndustrial Exploitation of Suplacu De Barcau Panonian Field inthe Socialist Republic of Romania", 4 th International Conferenceon Heavy Crude and tar Sands, Edmonton, Alta., Can., Aug. 7-12,1988.

8. Carcoana, A, Machedon, V. c., Pantazi, I. G., Petcovici, V. c.,Turta, A. T.: "In Situ Combustion - An Effective Method toEnhance Oil Recovery in Romania," 11 th World PetroleumCongress, London, UK, 1983.

9. Alexander, J. D., Martin, W. L., Dew, J. N.: "Factors AffectingFuel Availability and Composition During In Situ Combustion,"

10. Belgrave, J. D. M., Moore, R. G., Ursenbach, M. G., Bennion, DW.: "A Comprehensive Approach to In-Situ CombustionModelling," Paper SPEIDOE 20250 presented at SPEIDOE 7tSyrnposiumonEnhanced Oil Recovery, Tulsa, OK , USA, Apri22-25, 1990.

11. Hayashitani, M., Bennion, D. W., Donnelly, J. K, Moore, R. G,"Thermal Cracking of Athabasca Bitumen," OIL Sands oCanada-Venezuela, CIMSpecial Volume 17, pp 233-247, 1977

12. Adegbesan, K 0. , Donnelly, J. K., Moore, R. G., Bennion, DW.: 'LowTemperatureOxidation Kinetic Parameters for In-SituCombustionNumerical Simulation," paper SPE 12004 presentedat 58 th Annual Technical Conference and Exhibition, SanFrancisco, CA., USA, Oct 5-8, 1983.

13. Thomas, F. B., Moore, R. G., Bennion, D. W.: "KineticParameters for the High-Temperature Oxidation of In-SituCombustion Coke," J. Can. Pet. Tech., pp 60-67, Nov.lDec.1985.

14. Ivory, J., Lorimer, S., Redford, D., Nzekwu, B.: "Development oa Top-Down Combustion Process," presented at the AOSTRAConference"Oil Sands- Our Petroleum Future", Edmonton, AltaCan., April 4-7, 1993.

15. Kumar, M.: "Simulation of Laboratory In-Situ Data and effect oprocess Variations," paper SPE 16027 presented 9th SPESymposium on Reservoir Simulation, San Antonio, Tex., USAFeb. 1-4, 1987.

16. Marjerrison, D. M., Fassihi M. R.: "A Procedure for ScalinHeavy-Oil Combustion Tube Results to a Field Model," papeSPEIDOE 24175 presented at SPEIDOE 8th Symposium oEnhanced Oil Recovery, Tulsa, OK , USA, April 22-24, 1992.

17. Coats, K H.: In-Situ Combustion Model," Soc. Pet. Eng. J., p533-554, Dec., 1980.

18. Onyekonwu, M. 0. , Pande, K, Ramey, H. J., Brigham, W. G"Experimental and Simulation Studies of laboratory In-SitCombustion Recovery," paper SPE 15090 presented at thCalifornia Regional Meeting, Oakland, CA., USA, April 2-41986.

19. Ewing, R. E., Lazarov, R. D.: "Adaptive local Grid Refinement,paper SPE 17806 presented at SPE Rocky mountain Regionameeting, Casper, WY., USA, 1988.

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8Table 1. Experimental Pack Details

Test Pack Composition! Est. Saturation Communication path# Mass Bit. Wat. Coke2 So Sw Sg Dia. Sand Perm.kQ % % % cm mesh um.1 52.5 12.65 1.45 1.5 0.79 0.09 0.12 0.6 20-40 502 52.7 12.30 1.95 1.8 0.77 0.12 0.11 N.A. N.A. N.A.3 50.8 12.90 1.05 1.3 0.80 0.07 0.13 0.6 20-40 504 53.7 12.60 1.50 1.5 0.79 0.09 0.12 1.0 20-40 805 50.9 15.13 0.90 0.57 0.90 0.05 0.05 1.3 20-40 1356 50.8 15.30 1.49 0.32 0.89 0.09 0.02 1.3 60-70 19.57 48.5 15.48 0.71 0.28 0.89 0.09 0.02 1.3 60-70 19.5I-The packs for Tests 1-4 were prepared from a different batch of oil sand than Tests 5-7.2_A solid material initially present in oil sand that is not soluble in toluene but contributes to fuel load by burningat high temperature.

Table 2. Experimental Test Operating Strategiesest 1 2 3 4 5 6 7Total Run Time (11) 32.8 44.2 30.5 46.2 11.2 28.4 32ypeof Pre.heat Electric Electric Electric Electric Steam Steam Steam

f'!'e.heat Period (11) 5.6 22.6 9 29.1 6.6 8.6 10op Temperature at Start of 208 212 129 164 178 198 207f\ir Iniection (OC)lr'imeToObtain Ignition (11) 0.08 0.17 8 3.2 1 1.45 2bxveen in Injection Gas (%) 21 21 21 21 21 2 1 ~ 5 0 501\verageGas Injection Rate 0.4 0.46 0.63 0.68 0.32 0.42 0.39kg/h) (Target) (0.4) (0.8) (0.8 (0.8) (0.35) (0.35) (0.35)purationOf Gas Only

27.2 21.6 11.1 12 4.6 10 12lliection (h )Puration Of Gas.Steam . . 10.4 5.1 . 9.8 9.3njection (11)

Table 3. Oil ProductionIrest 1 2 3 4 5 6 7Final Recovery (%) 34.8 48.2 59.3 51.6 11.0 67.0 57.~ e c o v e r y at End 5.2 11.0 2.8 0.0 5.7 25.0 13.bf Pre-heat (%)~ a x . Bit. Recovery 2 4 13 5 2 5 6Rate (% OOIPIh)Max. Bit. Prod. 0.13 0.26 0.85 0.34 0.15 0.39 0.4Rate (ke!h)otal Bitumen 2.3 3.13 3.88 3.49 0.85 5.21 4.3Production (kg)Pinal Cum. OiBi!.! 1.10 0.74 0.88 0.78 0.44 0.69 0.9Pinal Cum. 0 ~ B i t . 2 1.36 0.96 0.92 0;81 1.23 1.11 1.21 _ Includes bitnmen produced during preheat - Based on cumulative 02injected2_Does not include bitumen produced during pre-heat Based oncumulativeO2 injected

Table 4. Produced GasIrest 1 2 3 4 5 6 7Max. CO2(mole %) 16;6 17 20.7 16.2 11.4 34.8 33Max. CO (mole %) 2.5 2.4 3 5.3 6.1 7.5 6.9M a x . C O ~ C O 7.2 9.6 >10 9.4 6.3 7.4 6.51\vg. CO2CO 5.7 7 8.1 6 2 4.7 5I\vg HlC Ratios' 1.65 1.76 1.10 1.09 4.54 2.68 2.40njected O2unreacted 15.1 11.5 6.07 14.4 30.8 18.0 19.8%O.ini,)njected O2forming CO2 57.5 67.8 86.6 64.9 35.5 50.6 41.3nd CO(%O2inj.)Reacted O2forming CO2 67.6 83.3 92.1 75.8 51.3 60.0 63.0nd CO(%O2reacted)'-Average dnring the distinct high temperaturecombustion period

Table 5. Combustion ParametersTest 1 2 3 4 5 6 7BurntVolume 0.0053 0.0079 0.0121 0.0073 0.013 0.009m3) -BumtVolume 20.0 29.9 45.8 27.7 49.0 37.0(%ofTotal) .!Fuel COllSUmed 0;68 0.68 1.20 0.77 0.07 0.81 0.98ke)lFuel Content 128 86 99 105 62 102(kg/m3) .O2Requirements! 2.8 2.5 2.1 2.6 3.7 3.3 3.0std. m3/ke fue))O2Requiremene 2.3 2.2 1.9 2.2 2.5 2.8 2.4(std. m3/kg fuel), - Based on injected oxygen2_Based on reacted oxygen

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Table 6. Bitumen and Coke Material Balance Test 79

Bitumen CokeInitial (kg) 7.51 0.14

Produced (kg) 4.601 0.982Final (kg) 1.51 0,33Difference (kg) 1.403 -1.1741 _ Includes an estimated 0.27 kg from produced gas condensate (20% oil)2 _ Fuel consumed3 _ Mass of bitumen that was converted to coke4 _ Mass of coke formed

Table 7. Component Properties For Input Into STARSComponent Water Asphaltene Maltene No Gas (COy) 0 0 CokeMolecularWt. 0.Q18 1.0928 0.406 0.028 0.043 0.032 0.01313(kg/gmo1e)Liquid Density [1] 1052 2410 - - - -(gmole/m3)Solid Phase Den. - - - - - - 1.05x10'(gmole/m3)LiQ. Compres. (kPa-') [1] 9.47xlO- 9.53x10- - - - -Coef. of Thermal [1] 4.5x10-4 5.85x10'4 - - - -Expansion (OCI )Critical Pres. (kPa) 22107 792 1478 3394 7176 5046 -Ctitical Temp. (0C) 374 904 619 -146.95 22 -119 -Vapor Heat Capacity = epg1 + cpg2 x T+ cpg3 x T' +c Ig4 x T" where T = temperature in OK (JIgmoleOC)cpg1 [1] - 9.92x102 [1] 19.8 28.1 -cgp2 [1] - 0 [1] 7.34xlO'" -3.68x10-o -CPjl,3 [1] - 0 [1] -5.6x10-' 1.74x10-' -cpg4 [11 - 0 [1] 1.71x10- 1.065xlO-" -Vaporization Enthalpv = hvr x (Tcrit- no. 3M (JIgmo1e) ,hv!' [1] - 1.03x104 - - I - I -Liquid Viscosity - Interpotaion of the following table based in corelation: U=aexpDI urn")30 DC r1l 1x10' 960 - - - -300 [11 6.5 0.22 - - - -500 1] 1.0 0.1 - - - -GaslLiQuid k-values = (kv11P + kv2 x P + kv3) x exp(kY 1(1- kY'))kv1 [1] - 1.888x10 - - - -kv2 [11 - 0.0 - - - -kv3 1] - 0,0 - - - -kv4 11 - -6562.3 - - - -kv5 11 - -79.98 - - - -

[1]- Values used from STARS mternal properties table

Table 8 Reaction Model ParametersReaction # Reactant A,. E (J/gmole) m n Enthalpy (J/gmole)

1 Maltenes 7.86x10 d- 2.347x10' 1.0 - 0.02 Asphaltenes 3.51xlO' d- 1.772x10' 1.0 - 0.03 Asphaltenes 1.18x10 J4 d' 1.763xl0' 1.0 - 0.04 Maltenes 11.1xlO" (d- kPa- L) 8.673xl04 1.0 0.425 (O?) 1.296xl0"5 Asphaltenes 3.58xl0" (d- kPa") 1.856x10' 1.0 4.57 (O?) 2.857xl0"6 OXVl!en 150.2 (d- kPa- l ) 3.476xlO 1.0 (0, ) 1.0 3.5xl0'

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VISOOSITY OF PRODUOED FLUIDS Teat 5 0.5 h. Teat 5 1 h. Teat 5 - 1.5 h.

Figure 8. Temperature Profiles - Test 52050

l.. ..........

START OFJI' AIR INJEOTION01: ' ----............----......-------. . . . .-----'o 5

200. . . . . . . . . . . . : - ~ - - l r . _ - + - - - - - - - - - - - - _ I

600r--. . . . .- - - . . - - - ~ - - - - - ~ .................. - ~,'./ " __ EXTERNAL ELEOTRIO PREHEAT'I . . -. STEAM INJEOTION PREHEAT' , , ' _ ORIGINAL}, ," . . . . L...- ....I, \ , , , \, \ " A ", --,, ,, ,, ,

RUN TIME (h)Figure 7. Viscosity of Produced Samples Recovery

3000

" ' ~ 7 - - Till 4-Electric Pnt-heat-- Till S-Staam Pre-h.at40/ ' ~ ' " - - - Teet ll-Steam Pre-heat...... - - Till 7-Steam Pnt-h.at"I20 ", OS~

I?OS ...... -00 10 ", ... ..9720 ... 30 ___04

. / ... -tart of ,. ... / : ~ ......Air In).... '" / "I"I ... ",'" " /

... / /" /I,.- OS L~ /' "...oo

40

80

20

80

100

Teat 712 h..at 7 8 h..at7 - 2 h.

Figure 9. Temperature Profiles - Test 7 Run Time (h)Figure 10. Recoveries of Steam Pre-heat Tests

Oummulative Oxygen/Bitumen Ratio1.6.............- - . , . . . . - -....... . . , . . . - - - - . . - - -. . . . .~ - - . . . . ,

LEVEL517.5 em)

BOTTOM(40om)

LEVEL 4(1205 em)

LEVEL 0(37.5 em)

LEVEL 3(1.5 em)

LEVELS(3205 em)

LI:VEL 2(2.5 em)

LEVEL 7(27.5 om)

VISUAL APPEARANCE OFTOCO? POST RUN PACK

LEVELS(22.5 em)

LI:VEL 1(0.0 em from top)

000255 20

'Ooe8 not Include bitumenproduced during pre-heat.

to

- EXTERNAL ELEOTRIO PREHEAT- - STEAM INJEOTION PREHEAT

0 ....-...4- - ' -"--- '- ......o

0,4

Run Time (h)Figure 11. Injected 0iProduced Bitumen Ratio Comparison of External Electric and Steam Pre-heat

BLAOKANo DARK ANDCONSOLIDATl!O BOFTllANO OILllAND

Figure 12. Post Run Pack - Test 7

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Figure 13. Numerical Simulation Grid Configuration

20 r--.--...,.....---.----,,...--.....---....,.....--.--.....,-

1

Tes t 5

-234

15

e108

5

Run Time (h)Figure 14. Bitumen Recovery - Comparison ofNumerand Experimental Results - Test 5

vvvV .clemVVVL-V

HorixoIItolProductionWellBIo