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An Improved Kinetics Model for In Situ Combustion of Pre-Steamed Oil Sands Min Yang,*,† Thomas G. Harding,‡ and Zhangxin Chen† †

Department of Chemical & Petroleum Engineering, Schulich School of Engineering, University of Calgary, Calgary, Alberta T2N 1N4, Canada ‡ Nexen Energy ULC, Calgary, Alberta T2P 3P7, Canada ABSTRACT: In situ combustion (ISC) has been recently evaluated as a follow-up process to steam assisted gravity drainage (SAGD) with the expectation to combine the advantages of SAGD and ISC. Before the design of such a hybrid process, it is important to understand the chemical reactions between air (or oxygen) and residual oil within a SAGD chamber in the presence of water and steam in order to simulate the process with a reasonable degree of confidence. In this study, an improved reaction kinetics scheme, in terms of Saturates, Aromatics, Resins, and Asphaltenes (SARA) fractions, is proposed to represent the complex chemical reactions during ramped temperature experiments. From the results of a set of laboratory ramped temperature oxidation (RTO) tests, the oxidation behavior at different temperatures has been carefully analyzed. On the basis of the analysis, a reaction kinetics model consisting of low temperature oxidation, thermal cracking, and high temperature oxidation reactions has been developed. This model has then been incorporated into CMG STARS to simulate RTO experiments. The experimental results of seven RTO tests, including temperature profiles, oxygen consumption, and carbon oxides production, have been successfully matched by tuning kinetic parameters. From the experimental and simulation study, it is found that the coke, which is formed through cracking reactions and traditionally considered to be the main source of fuel in ISC, reacts slowly at high temperatures in the RTO tests. The other source of fuel for combustion in the RTO tests is light hydrocarbons distilled from the original bitumen or cracked from oxidation and cracking reactions. These light hydrocarbons are responsible for the rapid high temperature behavior observed in the RTO tests. This work greatly increases the understanding of fuel sources, and the proposed model is able to predict oxidation/combustion behavior of pre-steamed Athabasca oil sands under a wide range of temperatures.

1. INTRODUCTION

that is based on an understanding of the complex reaction kinetics. Beginning in the 1970s, extensive research has been performed on the oxidation behavior during the ISC process for Athabasca oil sands at different temperatures, including Low Temperature Oxidation (LTO),8−10 Negative Temperature Gradient (NTG) region,11−14 and High Temperature Oxidation (HTO).15,16 On the basis of these valuable learnings on the nature of the chemical reactions, reaction kinetics modeling for numerical simulation prediction had commenced. Belgrave et al.17 originally proposed a comprehensive reaction kinetics model, including LTO, thermal cracking, and HTO. Coke was the only source of fuel in HTO. This model has been widely used to simulate laboratory and field scale ISC processes.18−21 However, bitumen composition in their model was limited to two components, namely, maltenes and asphaltenes. Jia et al.22 split the maltenes into a slow reactive fraction and a more reactive fraction. Since no detailed information on these two fractions was provided, it was difficult to incorporate them into a numerical simulation model. In addition, this model only consisted of LTO and thermal cracking reactions. At the same time, reaction kinetics modeling based on SARA (Saturates, Aromatics, Resins and Asphaltenes) fractions started and Freitag et al.23 presented a comprehensive pyrolysis reaction

Steam Assisted Gravity Drainage (SAGD) is a proven commercial thermal technology for oil sands recovery, although it still has many limitations. Low energy efficiency due to significant heat loss is one of the main concerns in this steam injection process. Also, high capital and operating costs make SAGD vulnerable to low oil prices. In Situ Combustion (ISC) provides an alternative to steam injection in which energy is generated within a reservoir by oxidizing either the original or a modified fraction of the crude oil. During an ISC process, air or oxygen containing gas is injected into the reservoir to oxidize a small fraction of oil, generating heat and forming a high temperature combustion front. Oil is displaced by the propagation of the combustion front as well as the sweep provided by the combustion product gases and steam.1−4 In recent years, ISC has been evaluated as a follow-up process to SAGD with the expectation to combine advantages of steam injection and in situ combustion (ISC).3,5−7 For the ideal situation, steam helps stabilize and preheat a reservoir for combustion, while oxygen provides energy or heat by oxidizing residual bitumen left behind in the steam-swept zone. Before the design of such a hybrid process, it is imperative to understand chemical reactions between air (or oxygen) and residual oil within a SAGD chamber in the presence of water and steam. It is also important for a design of hybrid steam/ combustion processes, development of field tests, and business case analysis to have a reliable numerical simulation capability © 2017 American Chemical Society

Received: October 6, 2016 Revised: February 10, 2017 Published: February 10, 2017 3546

DOI: 10.1021/acs.energyfuels.6b02582 Energy Fuels 2017, 31, 3546−3556

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Figure 1. Schematic of (a) RTO reactor26 and (b) simulation grid system.

model based on SARA fractions. Oxidation characteristics of each SARA fraction have also been studied.24,25 Considering reactivity of SARA fractions to oxygen addition, Sequera et al.26 improved Jia et al.’s22 model and presented a SARA based LTO model. In their improved model, aromatics and resins exhibited similar behavior and could be oxidized in the presence of an intermediate product at low temperatures. Such a model can be readily incorporated into a commercial simulator. Chen et al.4 further extended the aforementioned LTO model to a complete reaction model with consideration of NTG and HTO reactions. Both coke combustion and aromatic combustion reactions were included in HTO, while the roles of each fuel source in the ISC process were not clear. Recently, the principal reaction mechanisms have been screened to identify the generally accepted reaction regimes. In addition, key oxidation reactions have been summarized for an ISC process.27 For sources of fuel in HTO, it is commonly believed15,17,28 that a solid material (i.e., coke) is the main fuel for combustion. In recent years, attempts have been made to add other sources of fuel for HTO, such as the extension of the Belgrave et al.17 model by Yang and Gates30 to include methane and gas combustion reactions, in which the gas component is produced from thermal cracking of asphaltenes. Recent work investigating the application of air injection in light oil recovery (i.e., high pressure air injection)31,32 has shown that the combustion of solid residue (coke) is characterized by its slow reaction rate and high carbon dioxide concentration in the produced gas. It was concluded that the rate of oxygen consumption associated with burning of coke was much lower than expected based on the previous literature. Therefore, it was concluded that vapor phase combustion of hydrocarbons was responsible for the distinct reaction wave observed in their experiments and was included in their proposed model. No attempts, however, were made to examine the contribution of coke combustion to heat generation in an ISC process for oil sands. In addition, the

presence of connate water is seldom considered in previous studies. It is of practical and fundamental importance to analyze the sources of fuel and provide an insightful reaction kinetics model to predict oxidation/combustion behavior with the presence of water and steam. In this work, a comprehensive reaction kinetics scheme based on SARA fractions has been proposed to represent the complex chemical reactions with ramped temperature oxidation experiments using pre-steamed Athabasca oil sands. Bitumen samples for the tests were obtained from the Long Lake project operated by Nexen Energy ULC. On the basis of the analysis of a set of Ramped Temperature Oxidation (RTO) tests, some key modifications have been made and an improved reaction kinetics model has been presented consisting of LTO, NTG, and HTO reactions. This model has then been incorporated into CMG STARS to simulate RTO experiments. Kinetic parameters, including pre-exponential factor and activation energy, have been tuned to history match temperature profiles and produced gas compositions from RTO measurements.

2. RTO EXPERIMENTS AND NUMERICAL MODEL Ramped temperature oxidation (RTO) testing is one of the experimental methods used for oxidation kinetics study. During a RTO test, a one-dimensional oil-saturated core is heated with a flowing stream of air (or oxygen containing gas) under controlled conditions. Results from RTO tests allow for a comprehensive understanding of the global oxidation behavior under a wide range of temperatures. Reaction kinetics obtained from RTO tests can be directly incorporated into numerical simulation of in situ combustion processes.4,28 In the RTO apparatus, two identical reactors are mounted in an aluminum heating block. One reactor, which acts as the active reactor, is packed with core containing reservoir oil. The other reactor, which is the reference reactor, is packed only with a clean, dry core matrix. Figure 1a depicts the schematic of an RTO reactor. Temperature within each reactor is measured by five thermocouples located along the reactor. Air is injected from the bottom, and mass flow meters are 3547


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Energy & Fuels Table 1. Operating Conditions of Each RTO Test tests Test Test Test Test Test Test Test

#1 #2 #3 #4 #5 #6 #7

porosity (%)

water saturation (%)

oil saturation (%)

Injection gas rate (L(st)/h)

O2 concentration (mol %)

set point temperature (°C)

41.5 39.0 40.3 39.2 40.1 36.0 42.2

68.2 56.6 68.5 67.3 70.3 2.08 0.57

17.5 19.5 19.5 20.2 18.2 16.3 19.5

18.5 24.7 18.3 23.6 24.6 18.6 25.4

12.36 33.42 12.35 33.40 33.44 12.30 33.34

260 260 350 350 450 450 450

used to control the flow to the reactor. Combustion gases are produced from the top of the reactor. Produced gas composition is analyzed using gas chromatographs (GCs). More detailed information about the RTO apparatus and experimental procedures was presented by Moore et al.34 All RTO experiments associated with this study were conducted by the In Situ Combustion Research Group (ISCRG) at the University of Calgary. The oil samples used in this set of experiments were steamproduced Long Lake bitumen that was packed at around 20% oil saturation to represent approximately the residual oil saturation. The operating pressure used was around 2300 kPa, which is close to the steam injection pressure for Long Lake operations. The operating conditions of each test are listed in Table 1. In these RTO tests, only 6 or 7 g of bitumen was packed in the core and the corresponding initial oil saturation was less than 20%, representing the residual oil saturation left in the steam swept zone. The initial water saturation varied between 2% and 70%. Two levels of oxygen concentration in the injected gas, 12 and 33 mol %, were used. The maximum set point temperature applied in individual tests was one of 260, 350, or 450 °C. Simulation of RTO experiments was performed with a commercial simulator (STARS, Version 2014.10, Computer Modeling Group Ltd.). The active reactor was modeled in a three-dimensional (3D) radial (cylindrical) coordinate system, consisting of 4 grid blocks in the radial direction, 1 grid block in the azimuthal direction, and 27 grid blocks in the vertical direction. Previous studies4,26 associated with RTO experiment simulation adopted one-dimensional Cartesian grid systems without considering heat conduction between reactor and reactor wall, reactor wall and insulation, as well as insulation and heating block. In this study, four grid blocks in the radial direction were designed to represent the reactor, reactor wall, insulation, and heating block, respectively. In this way, heat loss during the test can be accurately simulated. The inner cross-sectional area for the reactor is 3.524 × 10−4 m2. The inlet and outlet of the reactor were packed with coarse sands, so the thickness of the first and last layer can be calculated based on the density and mass of the coarse sands. The thickness of other layers in the vertical direction was assigned as 1 cm, as suggested by previous grid size sensitivity evaluation.26 Five thermocouples were located in layers 24, 19, 14, 9, and 4, respectively. Figure 1 shows the schematic of the RTO reactor and simulation grid system. The input core porosity, fluid saturations, and operating parameters varied for different RTO tests. Fluid properties in terms of SARA fractions were characterized using CMG WinProp (version 2014.10, Computer Modeling Group Ltd.). Molecular weight and specific gravity of SARA fractions for Athabasca bitumen were presented in previous literature.35 SARA analysis was performed using a chromatography method based on ASTM D-2007. Critical properties of SARA fractions were adjusted to match bitumen density and viscosity at different temperatures. Table 2

displays properties of each SARA fraction,35 and Table 3 summarizes bitumen density and viscosity properties at different temperatures.7

Table 3. Bitumen Density and Viscosity at Different Temperature Conditions7

Saturates Aromatics Resins Asphaltenes

average molecular weight (g/mol)

density (kg/m3)

16.14 31.79 30.78 21.29

381 408 947 2005

882.7 994.9 1033.7 1200.2

density (g/cm3)

temperature (°C)

viscosity (cP)

25 40 55

1.0218 1.0130 1.0035

30 40 55

795 000 175 000 26 080

Because asphaltene is the heaviest component in bitumen and it does not vaporize, the K values for Asphaltenes are adjusted to be zero. Relative permeability curves are the same as those presented by Chen et al.4 in which RTO experiments were performed with Athabasca oil sands. Long Lake oil sands used in this work belong to the Athabasca region. Therefore, it is reasonable to use the same relative permeability curves in the Athabasca region. Simulation runs without chemical reactions were made first to ensure that the initial condition of the simulation matched the original mass of oil and water in the experiments. In the RTO tests, a heating rate of 40 °C/h was applied until the maximum set point temperature was achieved. This process can be simulated with a heat transfer model coupled in CMG STARS with the combination of keywords “UHTR” and “TMPSET”. Detailed information about the heating process simulation can be found in previous publications.4,26,32

3. DEVELOPMENT OF REACTION KINETICS MODEL The reaction kinetics model proposed by Chen et al.4 was selected as a starting point in this study to simulate oxidation behavior, as this model provides a comprehensive reaction scheme, including LTO, NTG, and HTO reactions. When incorporating such a model into the RTO simulation, an acceptable match between these simulation and experimental results was not achieved. The peak temperatures and shape of the temperature profiles could not be predicted well. Also, the timing of coke consumption was not accurate. The formation of carbon monoxide (CO) in HTO reactions was not considered. As a result, it was determined that an improved reaction kinetic scheme is needed to properly model the oxidation behavior in the RTO tests. The RTO data used by Chen et al.4 were collected on a wide variety of bitumen samples from the Athabasca region, but all the RTO data for the present study came from the Long Lake project. It is likely that the Long Lake bitumen has properties different enough from average Athabasca bitumen to require specific tuning of the reaction kinetics model. In order to improve the previous model, a large number of cases were created to investigate the effect of different parameters on the temperature profiles and produced gas composition, including activation energy, pre-exponent factor, and heat of reaction. Many different iterative attempts were conducted to match the experimental results. For the sake of brevity, only the key modifications and final reaction kinetics

Table 2. Composition and Properties of SARA Fractions35 mass percent (%)

temperature (°C)

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fractions, OxidAro, LightOil, and Gas), the solid component (i.e., coke), noncondensable gases (O2, CO2, and CO), and water. OxidAro represents the oxidation of the Aromatics component to hydroperoxides and falls into the Resins fraction.26 Coke is represented as a solid component. It was experimentally shown that a measured hydrogen/carbon ratio was 1.13, which gave a coke molecular weight of 13.13 g/mol.29 In this work, the molecular weight of coke is defined as 13 g/ mol. 3.1. LTO Reactions. The temperature range for heavy oil with LTO is from 150 to 300 °C.34 LTO is also known as oxygen addition reactions. The rate of oxygen addition is controlled by the diffusion rate of oxygen molecules into the crude oil.29 It is well-known that LTO reactions involving oxygen addition increase the molecular weight of components and thereby increase the oil viscosity. Reactions 1−3 are LTO reactions, in which reaction 1 acts as the precursor of reactions 2 and 3. The intermediate product (i.e., OxidAro) participates in the oxidation of aromatics and resins fractions, forming asphaltenes, carbon oxides, and water. A value of 90 kcal/mol O2 was chosen as a mean value for different oxidation reactions.26 The LTO scheme is essentially the same as that presented by Chen et al.4 and Sequera et al.,26 except that stoichiometric coefficients of oxygen, CO2, and water in reactions 2 and 3 have been adjusted. In the original LTO model presented by Sequera et al.,26 an average value of 0.5 g O2/g oil was used to form carbon oxides and water in reactions 2 and 3. Sequera et al.26 also pointed out that this value was too low for their specific tests conducted. In this work, an average of value of 0.65 g O2/g oil was used to form water and carbon oxides. The detailed information to calculate stoichiometric coefficients of O2, water, and CO2 can be found in the literature.41 Activation energy for reactions 2 and 3 in Table 4 was determined from history matching to be around 1.25 × 105 J/ mol, while most activation energy values reported previously in the literature were around 7 × 104 J/mol.25,26 It is possible that low oil saturation resulted in the large activation energy. The initial bitumen mass in this set of RTO tests is around 7 g, while the reported initial bitumen mass by Sequera et al.26 was 24.7 g. Oil saturation associated with this group of tests was less than 20%, and most of the pore space was saturated with either water or gas. It is believed that low temperature oxidation reactions are limited by the rate of diffusion of oxygen into the reactive hydrocarbon phase. Moreover, the global kinetics are mass transfer or diffusion controlled.36 As for this group of RTO tests, the molecules of oxygen were not able to easily come into a direct contact with reactive hydrocarbon during the heating ramp process, so the mass transfer rate was reduced. Mathematically, this phenomenon was modeled by increasing the activation energy in order to decrease the reaction rates. 3.2. NTG Reactions. The temperature range for NTG is between 280 and 330 °C, in which the oxidation rate decreases as temperature increases.34 NTG is the transition zone from the LTO to HTO.34 Thermal cracking reactions or pyrolysis reactions are dominant in this temperature range. It is believed that thermal cracking is responsible for the deposition of fuel for the subsequent combustion. In the proposed reaction scheme, reaction 5 is the Asphaltenes cracking reaction, which is modified from the Asphaltenes thermal decomposition reaction in previous literature.4,23 A laboratory pyrolysis study37 of the Long Lake bitumen indicates that pyrolysis products consist of coke, light hydrocarbon, and gases. The

model have been explained and listed as follows. The stoichiometric coefficients in each reaction represent conversions in mole percent. The tuned kinetic parameters for each reaction are presented in Table 4. Aromatics + 0.260O2 → 0.100OxidAro + 0.850Aromatics (1)

Aromatics + 0.100OxidAro + 12.159O2 → 0.300Asphaltenes + 4.018CO2 + 4.911H 2O

(2)

Resins + 0.100OxidAro + 29.447O2 → 0.670Asphaltenes + 9.327CO2 + 11.399H 2O

(3)

Saturates → 3.000LightOil

(4)

Asphaltenes → 0.780Saturates + 116.140Coke + 0.101CO2 + 6.480Gas

(5)

LightOil + 12.910O2 → 10.000H 2O + 6.750CO2 + 2.250CO

(6)

Gas + 3.250O2 → 3.000H 2O + 1.500CO2 + 0.500CO (7)

Coke + 1.125O2 → 0.750CO2 + 0.250CO + 0.500H 2O (8)

Table 4. Reaction Kinetic Parameters for Proposed Reaction Kinetics Scheme reactions

pre-exponent factor A (day−1)

R1 R2 R3 R4 R5 R6 R7 R8

7.2 1.44 1.44 1.44 1.04 9.91 9.91 2.74

× × × × × × × ×

104 1011 1011 1012 1014 1011 1011 1011

activation energy Ea (J/mol)

ΔH (J/mol)

oxygen reaction order

× × × × × × × ×

1.97 × 104 4.3 × 106 10.8 × 106 0 0 4.83 × 106 2.43 × 106 1.5 × 105

0.283 1.114 1.114 1 0.732 1 1 1

4.02 1.25 1.25 1.04 1.34 7.76 5.76 1.38

104 105 105 105 105 104 104 105

For each chemical reaction, the reaction rate can be calculated as displayed below29 Ri =

dCi = −kiPOa2Cib dt

(9)

where Ri is reaction rate. t is the time. Ci, ki, and PO2 are instantaneous concentration, rate constant, and oxygen partial pressure, respectively. a and b are reaction orders with respect to oxygen partial pressure and concentration, respectively. The reaction constant is governed by the Arrhenius expression ki = Ai

⎛ Ea , i ⎞ ⎟ ⎜− e ⎝ RT ⎠

(10)

where Ai is the pre-exponential factor and Ea is activation energy. R is ideal gas constant, and T is absolute temperature. In the proposed reaction scheme, a total of eight reactions are presented. The bitumen was characterized as Saturates, Aromatics, Resins, and Asphaltenes. Twelve components are involved, including the oil phase components (i.e., 4 SARA 3549


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Figure 2. Comparison between experimental and simulation results for Test #1.

produced gases contain more hydrocarbon gases than CO2. Therefore, a Gas component has been added in reaction 5, which represents lumped properties of all hydrocarbon gases with a molecular weight of 30 g/mol. The stoichiometric coefficients for CO2 and Gas are calculated based on the mole fraction of CO2 and all hydrocarbon gases. In addition, pyrolysis results indicate that an amount of the light hydrocarbon with a boiling point temperature less than 220 °C was produced.37 These light hydrocarbons range from C4 to C12 and fall in the Saturates fraction. These light hydrocarbons are lumped as the LightOil component in the reaction model with a molecular weight of 127 g/mol. At certain operating conditions, LightOil will be distilled from liquid phase and become part of the vapor phase. Therefore, reaction 4 was added in this work as a simplified representation for LightOil component vaporization. It should be noted that LightOil in reaction 4 represents light hydrocarbons distilled from the original bitumen as well as cracked from Asphaltenes. It should be noted that reaction 4 is not a true cracking reaction. 3.3. HTO Reactions. HTO is also known as the bond scission reaction, which involves the destructive oxidation of the hydrocarbon to produce water and carbon oxides. HTO is believed to ensure the success of an ISC process because its dominant temperature is higher than 350 °C4 and it has a high heat of reaction. Carbon dioxide (CO2), carbon monoxide (CO), and water are principal products of HTO. Coke was traditionally believed to be the main source of fuel, and therefore, the coke combustion reaction was included in HTO as reaction 8. In the previous literature, it was widely accepted that coke combustion generated significant thermal energy and sustained continuity of combustion. The recent laboratory test results indicate that coke combustion was characterized by its slow reaction rate.32,37,38 The experimental obtained activation energy for coke combustion was in the range of (1.2−1.5) × 105 J/mol.38 This result has been incorporated in the proposed reaction kinetics (see Table 4). As a result, coke burning was not able to support rapid and significant combustion performance. The Gas component produced through cracking reaction contains hydrocarbon gases, and therefore, it is proposed that Gas combustion contributes to HTO reactions as reaction 7. The LightOil component represents vaporized light hydrocarbons and acts as another source of fuel in HTO reactions as reaction 6. In summary, a total of three combustion reactions

(i.e., reactions 6, 7, and 8) are included to model the HTO. As mentioned in previous sections, the average molecular weights for Gas, LightOil, and Coke are 30, 127, and 13 g/mol, respectively. The stoichiometric coefficients are calculated based on mass balance and element balance. All of the HTO reactions are assumed to be the first-order reactions. The heat of combustion for each reaction is calculated based on complete combustion heat (CO2) of 4.39 × 105 J/mol O2 and partial combustion heat (CO) of 3.77 × 105 J/mol.26 Both LightOil and Gas will burn in the vapor phase, while coke will burn in the solid phase, albeit at a slower rate than the gas phase components. In order for the vapor phase to participate in the combustion, hydrocarbon vapor must fall in the flammable range.32 If the hydrocarbon vapor concentration is less than the lower limit of the flammable range, the hydrocarbon vapor is too lean to burn. On the other hand, if the hydrocarbon vapor concentration is higher than the upper limit of the flammable range, the hydrocarbon vapor is too rich too burn. Currently, CMG STARS does not have a specific option to model the flammable range. The key word RXEQFOR was adapted in the current simulation to mimic this behavior. The detailed information about flammable range simulation can be found in earlier publications.32,39

4. RESULTS AND DISCUSSION After performing sensitivity analysis with the modified reaction model, history matching of RTO tests was performed. For each test, temperature responses measured by five thermocouples and three major produced gas composition curves were predicted with the modified reaction model. Comparisons between experimental and simulation results are displayed in Figures 2−8. Both temperature and produced gas composition profiles were well matched with the measured results, indicating that the key oxidation behaviors are captured by the improved reaction model. It should be noted that the concentration of helium was measured during the tests because of the safety consideration. When the oxygen concentration returned to 0.5%, helium was injected until the produced helium concentration reached 99.5%. At this moment, the system was depressurized and allowed to cool to ambient temperature. In all tests, helium was injected after 20 h, while all kinds of oxidation activities occurred before 11 h. Therefore, the presence of helium did not affect the experimental and simulation results. The effects of three operating parameters on oxidation behavior were examined, including the oxygen 3550


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concentration, set point temperature, and initial water saturation. 4.1. Effect of Oxygen Concentration. The injected oxygen concentration was kept at 12.36% in Test #1 and Test #3, while the injected oxygen concentration was increased to 33.42% in Tests #2, #4, and #5. The oxidation behaviors of Tests #1 and #2 were compared because their other operating conditions were kept the same. In Test #1, limited exothermic activities in temperature were observed from 6 to 7 h in the

test. Meanwhile, the oxygen concentration reached zero, indicating that lack of adequate oxygen restricted the combustion performance. In Test #2 (Figure 3), an excellent match between simulated and measured profiles was achieved; especially, the thermal spikes in Zone 4 and Zone 5 were well matched. LTO reactions were dominant in Zones 1, 2, and 3 because simultaneous thermal bumps were observed below 300 °C. HTO reactions were dominant in Zone 4 and Zone 5 with peak temperatures over 350 °C. With elevated oxygen 3551


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concentration in Test #2, the produced oxygen concentration did not reach zero, indicating sufficient oxygen was present to support the thermal spikes in Zone 4 and Zone 5. A similar phenomenon can also be observed in Test #4 (Figure 5) and Test #5 (Figure 6). There were thermal bumps in Test #3 (Figure 4) at the temperature of 350 °C, accompanied by oxygen consumption and CO2 production, indicating coke combustion. It seems that these thermal bumps were not as obvious as those in Test #4 (Figure 5). This is because the injection oxygen concentration in Test #3 was lower than that in Test #4, resulting in the oxygen concentration from 8 to 10 h approaching zero. The lack of sufficient oxygen restricted coke consumption, leading to lower temperature in thermal bumps. 4.2. Effect of Temperature. The set point temperatures for Tests #2, #4, and #5 were 260, 350, and 450 °C, respectively. A heating ramp was applied until the set point temperature was reached. Experimental results in Test #2 showed that over 50% of the initial oil remained in the core as coke after the test. Similarly, a large amount of coke remained in the core after the tests in both experimental and simulation results in Test #1, in which the set point temperature was also 260 °C. This finding leads to the conclusion that coke is not the main fuel responsible for the thermal spikes in Zone 4 and Zone 5 in Test #2. Simulation results show that around 50% mass of the initial oil remained in the reactor as coke, which is matched with residual coke in the experiment. Each reaction has been studied separately, and it was found that significant thermal spikes in Zone 4 and Zone 5 are the result of LightOil and Gas combustion, while coke combustion contributes little to the temperature increase because of its slow reaction rate. Though this finding questions the traditional assumption of coke as the main fuel for combustion reactions, similar results have been documented in the recent literature from both laboratory test results33,42 and field pilot testing.40 The pilot test results indicated that the main fuel for the combustion appeared to be a lighter oil fraction which was a product of oxidation and cracking reactions. The operating conditions for Test #4 were the same as those in Test #2, except that the heating ramp was extended to 350 °C. The temperature profiles for Tests #2 and #4 (Figures 3 and 5) exhibited the exact same behavior before 6 h of the tests because operational conditions for those two tests were the same during that time. The secondary thermal bumps were observed when the heating ramp was extended to 350 °C in Test #4. These bumps were accompanied by oxygen uptake

and CO2 production. Simulation results indicate that such simultaneous bumps are the result of coke combustion. However, it seems that both oxygen consumption and CO2 production from 8 to 9 h are higher in the simulation than in measured results. It is possible that coke production from the Asphaltenes cracking reaction (i.e., reaction 5) is overestimated. In Test #5, the heating ramp was further extended to 450 °C, while the other operational conditions were kept the same as those in Test #4. Figure 6 displays experimental and simulated results for Test #5. Both temperature and produced gas composition profiles exhibited the same behavior as those obtained in Test #4. The extension heating ramp did not extend the second thermal bumps. Both experimental and simulation results indicated that all hydrocarbons have been either produced or consumed under this operating condition. All RTO tests in which the maximum temperature ramps were either 350 or 450 °C displayed the second exothermic activity. In addition, the post core analysis for all of these tests indicated that limited amounts of coke remained. When experimentally obtained coke combustion kinetic parameters were incorporated into the simulation model, the simulated temperature as well as produced gas composition indicated a good match with the RTO experimental results. Such agreements confirm previous assumptions that coke combustion is responsible for the second region of simultaneous bumps and significant coke combustion occurs at a high temperature of around 350 °C. One limitation of this model is that coke production is overestimated using the modified model. 4.3. Effect of Initial Water Saturation. The initial water saturation in Tests #1−5 was high, varying between 56.6% and 70.3%, while the low initial water saturation was packed in Tests #6 and #7 as 2.08% and 0.57%, respectively. In Test #1 (Figure 2), a series of thermal dips were observed in temperature profiles between 3 and 5 h in the test as shown in Figure 2b, which is consistent with the measured results (see Figure 2a). These thermal dips are the result of the propagation of water evaporation fronts. The initial water saturation in this test was as high as 68.2%. During the heating ramp process, vaporization of water from the liquid phase to the vapor phase absorbs part of the heat and results in the endothermic activity in the temperature profile. Once the vapor phase of water had passed by, the temperature in each zone began to increase. Similarly, a series of thermal dips were observed in the temperature profiles of Tests #2−5 (Figures 3−6) because of the high initial water saturation. In Tests #6 and #7 (Figures 7 3552


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(Figure 7) did not match well. The experimentally obtained oxygen concentration in the produced gas approached zero during LTO (see Figure 7a), while the simulated oxygen concentration did not reach zero during 5−7 h of the test (see Figure 7b), indicating that oxygen consumption in LTO is underestimated by the proposed model. The simulated oxygen concentrations in Tests #1 and #2 approached zero due to the fact that a portion of hydrocarbon is combusted with oxygen. This is also the reason that thermal fronts in Zones 4 and 5 are observed to exceed 300 °C in both Tests #1 and #3 (see Figures 2b and 4b). The underestimation of oxygen consumption was also reported in the previous literature.4,26 The authors26 who originally developed the LTO model suggested that there might be a different reaction, involving the oxidation of another fraction of bitumen (such as saturates) that had not been considered. However, the key oxidation behavior has been captured. The comparison of the experimental and simulated results is presented in Figures 2−8. In order to make the comparison quantitative, statistical analysis results in terms of the mean absolute percentage error (MAPE) and root-mean-square deviation (RMSD) are presented in Table 5. More specifically, a total of 45 data points was selected from each curve and 35 of the points were selected between 4 and 11 h of the test, during which time the oxidation behavior was observed. The mean absolute percentage error (MAPE) of the temperature profiles for seven tests ranges from 1.94% to 2.89%, indicating a good

and 8), thermal dips in temperature ramps from previous tests were not observed in this test because of extremely low initial water saturation. This finding is the same as that documented in the literature.33 In Test #7, the operating conditions were the same as those in Test #5 except that the initial connate water saturation was packed as 0.57%. Figure 8 presents the comparisons between measured and simulated results for Test #7. The temperature responses are matched well. The first simultaneous thermal bumps in all five zones from 4 to 6 h of the test were dominated by LTO reactions. One interesting observation was that rapid bond scission reactions in previous tests associated with a high level of oxygen concentration (i.e., Tests #2, #4, and #5) were not present in this test. This suggested that hydrocarbon vapor concentration in this test was too low to support significant combustion performance. Without initial connate water, the hydrocarbon vapor could be freely distilled into the air stream so that the concentration was below the lower limit of the flammable range. On the basis of the above discussion, the presence of sufficient oxygen helps support thermal spikes in the temperature profiles. In addition, light hydrocarbon behaves as the major source of fuel and is responsible for these thermal spikes. Significant coke consumption occurs when the temperature is higher than 350 °C. With the proposed model, there is a good match between the measured and simulation temperature profiles. However, the produced gas composition in Test #6 3553


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Energy & Fuels

This is mainly because some of the measured O2 and CO2 mole fractions were less than 0.05. Therefore, even a small deviation between measured and simulated values resulted in a large MAPE. Moreover, it is difficult to estimate the errors of the experimental measurements due to their complexity. The rootmean-square deviation (RMSD) was also calculated. The RMSD for gas compositions was within 0.05 for each test. These statistical analysis results support the predictive capability of the proposed reaction kinetics model.

Table 5. Comparison of Measured and Simulated Results test Test #1

curve type temperature

gas composition

Test #2

temperature

gas composition

Test #3

temperature

gas composition

Test #4

temperature

gas composition

Test #5

temperature

gas composition

Test #6

temperature

gas composition

Test #7

temperature

gas composition

Zone Zone Zone Zone Zone N2 O2 CO2 Zone Zone Zone Zone Zone N2 O2 CO2 Zone Zone Zone Zone Zone N2 O2 CO2 Zone Zone Zone Zone Zone N2 O2 CO2 Zone Zone Zone Zone Zone N2 O2 CO2 Zone Zone Zone Zone Zone N2 O2 CO2 Zone Zone Zone Zone Zone N2 O2 CO2

1 2 3 4 5

1 2 3 4 5

1 2 3 4 5

1 2 3 4 5

1 2 3 4 5

1 2 3 4 5

1 2 3 4 5

MAPE (%)

RMSD

1.36 1.35 1.75 2.68 2.63 0.96 57.35 27.92 1.48 2.75 2.86 3.04 2.09 0.71 6.97 14.97 1.72 2.19 2.43 3.28 3.42 1.54 90.81 83.41 1.90 1.95 2.81 2.42 2.84 1.70 18.13 83.41 1.37 2.01 2.71 3.04 2.78 1.73 17.25 34.93 1.61 1.59 2.11 1.54 2.85 1.60 77.31 87.58 2.59 2.85 2.63 2.73 4.10 2.08 24.97 56.25

4.71 4.52 5.87 12.68 13.29 0.02 0.03 0.01 5.79 10.82 10.54 12.99 7.52 0.02 0.03 0.02 5.28 6.56 7.77 12.85 14.32 0.02 0.03 0.02 6.87 8.97 10.64 9.23 30.84 0.02 0.05 0.03 4.85 9.63 11.93 10.44 19.91 0.02 0.05 0.03 5.38 5.04 6.27 4.92 11.64 0.02 0.02 0.02 8.01 12.26 10.89 10.91 15.50 0.02 0.08 0.05

5. CONCLUSIONS An improved reaction kinetics scheme based on SARA fractions has been proposed to represent the main oxidation reactions for pre-steamed bitumen during RTO experiments. The proposed reaction scheme is able to reasonably predict temperature profiles and produced gas compositions in RTO experiments. Application of the reaction scheme into numerical simulation has indicated that coke formed through pyrolysis reactions is not the only source of fuel and that distilled or cracked light hydrocarbons participate significantly in the reactions. Moreover, light hydrocarbon reacts with oxygen quickly and contributes to the rapid combustion behavior. Coke reacts with oxygen slowly, and significant coke consumption occurs only at high temperature conditions (higher than 350 °C). The low initial oil saturation and high water saturation in the core samples of the RTO tests reduces the diffusion rates between oxygen and hydrocarbon and, therefore, results in the large activation energy in the LTO reactions. The proposed reaction kinetics scheme in this study is ready to be applied in numerical simulation of an ISC process.

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AUTHOR INFORMATION

Corresponding Author

*E-mail: [email protected]. ORCID

Min Yang: 0000-0001-7963-1267 Thomas G. Harding: 0000-0002-5506-9457 Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors would like to thank Nexen Energy ULC for financing the laboratory studies and for permission to publish this paper. Also, the In Situ Combustion Research Group at the University of Calgary is acknowledged for collecting the laboratory data. This research has been made possible by contributions from the NSERC/AIEES/Foundation CMG Industrial Research Chair in Reservoir Simulation and the AITF (iCore) Chair in Reservoir Modelling, and the Frank and Sarah Meyer Foundation CMG Collaboration Centre.

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REFERENCES

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Energy & Fuels Calorimetry for Air Injection Enhanced Oil Recovery Processes. SPE Reservoir Eval. Eng. [Prepr.] 2016; SPE-180451-PA.http://dx.doi.org/ 10.2118/180451-PA.

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