VOLUME 19, NUMBER 5
SEPTEMBER/OCTOBER 2005
© Copyright 2005 American Chemical Society
Articles Maximizing Aromatic Hydrogenation of Bitumen-Derived Light Gas Oil: Statistical Approach and Kinetic Studies Abena Owusu-Boakye, Ajay K. Dalai,* and Deena Ferdous Catalysis and Chemical Reaction Engineering Laboratories, University of Saskatchewan, Saskatoon, Canada
John Adjaye Syncrude Edmonton Research Centre, Edmonton, Canada Received September 2, 2004. Revised Manuscript Received June 14, 2005
Bitumen-derived light gas oil (LGO) was hydrotreated over commercial NiMo/Al2O3 catalysts in a trickle bed reactor. Statistical design of experiments was used to develop response surface models for predicting percentage conversions of aromatics, sulfur, and nitrogen in the LGO feed from Athabasca oil sands. The statistical approach was also used to study the effect of process variables and their interaction on aromatic hydrogenation (AHYD), hydrodesulfurization (HDS), and hydrodenitrogenation (HDN) activities. The two-level interaction between temperature and pressure was determined to affect AHYD significantly, whereas the interaction between temperature and the liquid hourly space velocity (LHSV) was the most important parameter affecting both HDS and HDN activities. Optimal conditions for the conversion of aromatics were observed at a temperature of 379 °C, a pressure of 11.0 MPa, and an LHSV of 0.6 h-1. Under these conditions, a maximum conversion of 63% can be attained. The cetane index of the diesel fraction was affected by changes in the aromatic compounds, as well as by the temperature and pressure of hydrotreating. Product distribution and gasoline yield of the liquid products were also greatly influenced by the reaction temperature, with a slight impact from pressure and LHSV. The kinetics of AHYD was modeled using a singe-site mechanism form of the LangmuirHinshelwood rate of reaction, whereas HDS and HDN were best described by an irreversible pseudo-first-order power-law reaction. Results of the kinetic studies showed significant inhibition of hydrogenation by hydrogen sulfide (H2S) gas produced during the HDS process.
1. Introduction The gradual shift from conventional petroleum crude to synthetic fuels derived from oil sands has intensified the quest for an appropriate technology and catalytic * Author to whom correspondence should be addressed. Telephone: (306) 966-4771. Fax: (306) 966-4777. E-mail address: dalai@ engr.usask.ca.
process for producing high-quality transportation fuels. This is because distillates produced from synthetic crude are characterized by high aromatic contents, which degrade the quality of on-road fuels.1 High concentrations of sulfur and nitrogen species existing in the form of elemental sulfur and fused heterocyclic compounds (1) Yui, S. K.; Sanford, C. S. Proceedings, 5th International Conference on Heavy Crude and Tar Sands; 1988; p 299.
10.1021/ef040080i CCC: $30.25 © 2005 American Chemical Society Published on Web 07/29/2005
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such as benzothiophenes and quinoline are also found in the distillates. Combustion of these fuels leads to the formation of NOx, SOx, and particulate matter in exhaust gases, all of which are detrimental to the environment. The heightened concern to produce cleaner fuels has led to the enforcement of stricter fuel specifications on the quality of diesel produced in both Canada and the United States. For example, in the case of middle distillates, which serve as feedstock for diesel and gasoline fuels, higher-quality standards such as low sulfur contents and reduced polyaromatic concentrations with improved cetane index (CI) and lower densities are required.2 Sulfur and aromatic hydrocarbon contents in diesel are expected to be reduced from the current level of 500 to 15 ppm and 10 vol % by 2006, respectively, with a minimum CI of 42.3 The cetane index measures the auto-ignition properties of diesel. It is also a measure of the hydrogen-to-carbon ratio of the hydrocarbon components and is inversely related to the amount of aromatics and cycloparaffins contained in the diesel fraction.4 For example, high amounts of aromatics (60-70 vol %) observed in catalytically cracked light cycle oil exhibit significantly low cetane values (CI ) 10-20) to be used as a blending component of the diesel fuel.5 Catalytic hydrogenation of aromatic compounds via hydrotreating is the most efficient method for reducing the aromatic contents in middle distillates and consequently increasing the CI value.6 Typical hydrotreating catalysts such as NiMo and NiW on alumina or silica supports are usually used. Noble catalysts such as supported Pd/Pt can also be used but with great caution, because these catalysts have a very low resistance to sulfur poisoning.7 The hydrogenation reaction is reversible and exothermic and, under typical hydrotreating processes, complete conversion is not feasible, because of thermodynamic equilibrium limitations that occur at elevated reaction temperatures.8 For aromatic hydrocarbons that contain fused heteroatoms, such as pyridine, removal of the heteroatom occurs sequentially, via either hydrogenation followed by hydrogenolysis or vice versa. Hydrogenation involves removal of the ring containing the heteraoatom (e.g., nitrogen), whereas hydrogenolysis involves the direct extraction of the heteroatom by breaking the C-N and C-S bonds.6,9 However, because hydrogenation is exothermic and hydrogenolysis is endothermic, petroleum refinery industries are faced with the challenge of striking a balance among temperature, pressure, liquid hourly space velcoity (LHSV), and H2/oil ratio input between the two reaction mechanisms during hydrotreating. Schematic representation (2) McVicker, G. B.; Daage, M.; Touvelle, M. S.; Hudson, C. W.; Klein, D. P.; Baird, W. C., Jr.; Cook, B. R.; Chen, J. G.; Hantzer, S.; Vaughan, D. E. W.; Ellis, E. S.; Feeley, O. C. J. Catal. 2002, 210, 137. (3) Diesel Fuel Quality, Advance Notice of Proposed Rulemaking, EPA420-F-99-01, United States Environmental Protection Agency (USEPA), Office of Mobile Sources, 1999. (4) Aribas, M. A.; Martinez, A. Appl. Catal., A 2002, 230, 203. (5) Gary, J. H.; Handwerk, G. E. Petroleum Refining: Technology and Economics; Marcel Dekker: New York, 2001; p 465. (6) Stanislaus, A.; Cooper, B. H. Catal. Rev.-Sci. Eng. 1994, 36, 75. (7) Girgis, M. G.; Gates, B. C. Ind. Eng. Chem. Res. 1991, 30, 2021. (8) Wilson, M. F.; Fisher, I. P.; Kriz, J. F. J. Catal. 1985, 95, 155. (9) Fujikawa, T.; Idei, K.; Ebihara, T.; Mizuguchi, H.; Usui, K. Appl. Catal., A. 2000, 192, 253.
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of this mechanism with nitrogen removal from quinoline is shown in eq 1.
Another challenge for the hydrogenation of aromatic compounds in industrial feed is inhibition by organic sulfur and nitrogen compounds present in the feedstock, as well as hydrogen sulfide (H2S) and ammonia (NH3) produced by hydrodesulfurization (HDS) and hydrodenitrogenation (HDN) processes, respectively.6,9 In the inhibition process, NH3 and H2S are strongly adsorbed on the hydrogenation centers of the hydrotreating catalysts, compared to the other catalytic centers responsible for the hydrogenolysis reaction. This condition presents a competitive environment for the adsorption of nitrogen, sulfur, and aromatics compounds in the feed toward the hydrogenation sites. To overcome these challenges, it is necessary to predict the kinetic parameters that control aromatic hydrogenation (AHYD) at any point in time during hydrotreatment. Knowledge of the kinetics of AHYD will serve as the basis for selecting the best process variables and catalysts for maximum hydrogenation of aromatics.6 However, information on the kinetics of AHYD in synthetic middle distillates is scarce, because of the complexity of interpreting the experimental data. Almost all available results existing in the literature10-12 have been obtained using model compounds, which makes direct application of these results in the prediction of the kinetic parameters for industrial processes quite difficult. The main objective of this study is to use the statistical design of experiments to study the interaction effects of temperature, pressure, and LHSV on hydrotreatment and also determine the best combination of process variables for maximum conversion of aromatics in an LGO feed derived from Athabasca bitumen. Liquid product distribution and gasoline yield of the hydrotreated feed subjected to the different hydrotreating severities will also be investigated. Kinetics of AHYD, HDS, and HDN will also be studied by varying temperature at the optimum pressure and LHSV conditions obtained from the statistical study. 2. Experimental Section 2.1. Statistical Design of Experiments. The search for an appropriate predictive model to estimate the percent conversions of aromatics, sulfur, and nitrogen and to obtain the optimum operating conditions for maximum saturation of aromatics was investigated using a response surface methodology (RSM) via the central composite inscribed (CCI) design.13 RSM consists of a group of statistical and mathematical techniques for empirical model building and exploitation that (10) Spare, A. V.; Gates, B. C. Ind. Eng. Chem., Process Des. Dev. 1981 20, 68. (11) Rautanen, P. A.; Aittamaa, J. R.; Krause, A. O. I. Chem. Eng. Sci. 2001, 56, 1247. (12) Huang, T.-C.; Kang, B. C. Ind. Eng. Chem. Res. 1995, 34, 1140. (13) Box, G. E. P.; Draper, N. R. Empirical Model-Building and Response Surfaces; Wiley: New York, 1987; p 515.
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Table 1. Actual and Coded Levels of the Design Parameters
Table 2. Composition and Physical Properties of Light Gas Oil (LGO) Feed
temp (°C)
pressure (MPa)
liquid hourly space velocity, LHSV (h-1)
code
340 350 365 380 390
6.9 8.2 9.6 11.0 12.4
0.5 0.8 1.25 1.7 2.0
-* -1 0 +1 +*
relate an output or a response to a number of predictors or the input variable that affect it. The input variables investigated in this study are reaction temperature, pressure, and liquid hourly space velocity (LHSV) at a constant H2/oil ratio. The central composite design consists of an embedded factorial and fractional factorial design, characterized by a central point (denoted by “o”), axial points (denoted as -1, +1), and star points (denoted with /). The star points represent the extreme values of each process variable. Table 1 shows the coded and actual levels of the process variables used to design the experimental program for this study. The total number of runs (N) required was 20, calculated from x
N ) 2 + 2x + 6 ) 20 trials
(2)
where 6 is the number of replicates at the center levels and x is the number of design factors or input variables under investigation. The Design Expert software (Version 6) was used to design the experiments and process the data. 2.1.1. Test for Significance of Models and Estimated Coefficients. Analysis of Variance (ANOVA) approach was used to test for the adequacy of the regression models for explaining the experimental data at a 95% confidence interval. A model was considered desirable when its probability (p-value) from ANOVA was 0.1 are desired for the lack of fit test. 2.2. Materials and Method. Hydrotreatment was conducted under isothermal conditions in a stainless-steel, fixedbed reactor 10 mm in diameter and 285 mm long. Extrudate commercial NiMo/Al2O3 catalysts (∼1.2-2 mm in size) were loaded in the reactor. The Brunauer-Emmett-Teller (BET) surface area of the catalyst, as measured by an automated gas (N2) adsorption analyzer (Micrometrics model ASAP 2000) was 165 m2/g with a pore volume of 0.45 cm3/g.16 The catalyst bed was maintained at a height of 12 cm and diluted with a 90mesh size silicon carbide (SiC), to provide complete catalyst wetting and reduce radial dispersion. Below and above the catalyst bed were successive layers of 60, 46 and 16 mesh SiC, as well as a layer of 3-mm-diameter glass beads to prevent heat loss. The large pellet sizes of the catalyst, as well as the compact nature of catalyst bed packing provided by the 90 mesh SiC, reduced the bed porosity, thus minimizing any diffusion effects and providing plug flow conditions for reaction. Plug flow pattern was also ensured by the different layers of SiC and inert glass beads. Schematic representation of the experimental setup is shown in Figure 1. The catalyst bed was calibrated with a thermocouple that was placed in direct contact with the middle part of the bed. Calibration was performed under the same conditions as the experiments, except that, during the experiments, the ther(14) Design Expert Software, Version 6, and User’s Guide, Stat-Ease, Inc., 2003. (15) Rigas, F.; Panteleos, P.; Laoudis, C. Global NEST: Int. J. 2000, 2, 245. (16) Bej, S. K.; Dalai, A. K.; Adjaye, J. Energy Fuels 2000, 15, 377.
property
value
density @ 15 °C (g/cm3) carbon content (wt %) hydrogen content (wt %) total sulfur content (wppm)a total nitrogen content (wppm)a basic nitrogen content (wppm) aromatics content (wt %)b cetane index, CIc simulated distillation (°C)d initial boiling point, IBP 10 wt % 30 wt % 50 wt % 80 wt % 90 wt % final boiling point, FBP
0.9007 85.5 12.06 17420 461 247 17.1 35.7 169.6 183.0 230.2 279.6 349.1 375.5 439.2
a From ASTM D-4629. b According to 13C NMR analysis. c From ASTM D-976. d From ASTM D-2887 HT.
mocouple was removed, to avoid disturbance in the packing and improve the flow pattern. The catalyst bed was pretreated (sulfided) with 2.5 wt % butanethiol at temperatures of 193 and 343 °C for 2 days at a constant pressure and space velocity of 9.0 MPa and 1.0 h-1, respectively. Following sulfidation, the catalyst bed was then stabilized with a heavy gas oil feed for 7 days at a temperature of 375 °C, pressure of 9.0 MPa, and LHSV of 1.0 h-1. Experiments were performed using light gas oil (LGO) feed from Athabasca oil sands, the properties of which are provided in Table 2. For the first part of the experirments involving the statistical study, experiments were conducted by varying the reactor temperature (340-390 °C), pressure (6.9-12.4 MPa), and LHSV (0.5-2.0 h-1). The H2/feed ratio was maintained constant (550 mL/mL) throughout the study. These process variables were selected based on typical hydrotreating conditions that are used in petroleum refining industries and also by various researchers.1-7 The same catalyst loading was used for the entire experiments. However, for the kinetic study, fresh NiMo catalyst was loaded in the reactor and the same pre-experimental procedures were used. Experiments for the kinetic studies were performed by varying the temperature from 340 °C to 390 °C at constant pressure (11.0 MPa), LHSV (0.6 h-1), and H2/oil ratio (550 mL/mL). Each experimental condition was run for a period of 3 days and products were collected at 24-h interval. To ensure reproducibility of the results, deactivation studies were performed at a temperature of 375 °C, pressure of 9.0 MPa, and LHSV of 1.0 h-1 after every 20 days. Results from the deactivation studies gave variations of 4%, 3%, and 2.6% in the aromatics, sulfur, and nitrogen conversions, respectively. The product samples were analyzed for total aromaticity using 13C NMR spectroscopy (Bruker model Avance 500). Sulfur and nitrogen concentrations were measured using a sulfur and nitrogen analyzer (ANTEK 9000). The boiling point distribution of petroleum fractions in the products was determined by gas chromatography (GC)-simulated distillation (Model CP 3800), according to the ASTM D2887 HT procedure.
3. Results and Discussion 3.1. ANOVA. The ANOVA procedure was used to develop regression models for AHYD, HDS, and HDN activities to predict the conversions of aromatics, sulfur, and nitrogen, respectively, during hydrotreatment. The significant main and interaction effects in the predictive models were selected based on their probability values. Individual and interaction terms with p-values of >0.05 were removed from the final expressions of the models,
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Figure 1. Schematic diagram of the experimental setup. Table 3. Analysis of Variance (ANOVA) for the Overall and Reduced Quadratic Model of AHYD Overall Model source model T P LHSV T2 P2 LHSV2 T×P T × LHSV LHSV × P lack of fit R2
sum of squares
p-value
3929.66 0.0001 692.8 0.0001 943.7 0.0001 188.6 0.0014 696.1 0.0014 1254.7 0.0001 37.89 0.0657 434.0 0.0001 23.36 0.1328 6.60 0.3997 66.51 0.1232 0.9833
Reduced Model sum of squares
p-value
3870 746.57 969.75 187.52 684.20 1236.65
0.0001 0.0001 0.0001 0.002 0.0001 0.0001
399.54
0.0001
126.17
0.1082 0.9684
using a backward elimination method (Tables 3 and 4). The predictive response models were based on conversion, which is defined as
Conversion (%) )
[feed] - [products] × 100 (3) [feed]
where [feed] and [products] are the concentrations of the species in the feed and product samples, respectively. 3.1.1. Effect of Process Variables on Aromatic Hydrogenation. The reduced quadratic model for AHYD showing only the significant factors is represented in eq 4. These factors were selected based on the ANOVA parameters in Table 3. Factors with p-values of 425
3.3. Hydrotreatment and Mild Hydrocracking (MHC). During hydrotreatment, some of the heavy materials in the feed are converted to lighter fractions under the conditions of mild hydrocracking (MHC).19 MHC is an imperative catalytic process for producing high-quality fuel from heavy feedstock. MHC occurs (19) Koseoglu, R. O.; Phillips, C. R. Fuel 1988, 67, 1201.
during high severities of hydrotreatment at temperatures of 390-420 °C. Unlike the conventional hydrocracking process, MHC improves hydrogen consumption economy and minimizes the formation of undesirable lighter products.20 Studies show that MHC also helps to improve the CI value of diesel fuels by converting naphthalene and branched alkanes to lighter, gasolinerange products.2,21 The feed used in this study is predominantly LGO with a large boiling range of 190-435 °C. Analysis of the feed shows that it contains some tail ends, such as the heavy gas oil (HGO) and vacuum gas oil (VGO); in addition, because hydrotreatment is performed at high severities, it is expected that the feed will undergo MHC. This section of the paper studies the effect of process variables such as temperature, pressure, and LHSV and MHC on product distribution and gasoline yield in the feedstock. The feed and sample products have been grouped into five main fractions in Table 6, based on their boiling ranges. Data obtained from the GC-simulated distillation are based on liquid products only. Gasoline yield is defined as
Yield (%) )
specific product amount (wt%) × 100 feed amount (wt% ) (8)
3.3.1. Effect of Temperature on Product Distribution and Gasoline Yield. Figure 5a illustrates the effect of temperature on liquid product distribution. It is ob(20) Yui, S. M. AOSTRA J. Res. 1989, 5, 211. (21) Botchwey, C.; Dalai, A. J.; Adjaye, J. Energy Fuels 2003, 17, 1372.
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Figure 6. Impact of pressure on (a) liquid product distribution and (b) gasoline yield at a temperature of 365 °C, LHSV of 1.25 h-1, and a H2/oil ratio of 550 mL/mL.
served that, with increasing temperature, more gasoline fractions are produced with a corresponding decline in the heavy cuts such as HGO and VGO. This may be as a result of the ease at with which species with large molecular weights (such as sulfur and nitrogen) in the feed are removed at high reaction temperatures. In addition, increasing the temperature enhances hydrocracking, which leads to the formation of lighter fractions with lower molecular weights and boiling points. It is interesting to note that kerosene and the LGO fractions remain relatively the same throughout the reaction. This indicates that the majority of the lighter fractions are produced from the HGO and VGO fractions. Figure 5b shows the effect of temperature on gasoline yield. It is observed that the yield increases from 0 wt % to 31 wt % as the temperature was varied in the range of 340-390 °C. Thus, more gasoline is produced at higher temperatures of hydrotreatment. 3.3.2. Effect of Pressure on the Distribution of Products and Gasoline Yield. The effect of pressure on product distribution and gasoline yield was investigated by varying the reaction pressure from 6.9 to 12.4 MPa at a constant temperature and LHSV (365 °C and 1.25 h-1 respectively). It is observed from Figure 6a that pressure does not have a significant impact on the overall distribution of petroleum fractions. This is in agreement with the studies by Kim et al.,22 who reported an insignificant contribution of pressure to product distribution when PR-spring bitumen oil was hydrotreated over a HDM catalyst. PR-spring bitumen oil is one of (22) Kim, J. W.; Hanson, F. V. Energy 1992, 23, 3.
Owusu-Boakye et al.
Figure 7. Effect of LHSV on (a) liquid product distribution and (b) gasoline yield at a temperature of 365 °C and a pressure of 9.6 MPa.
the seven special tar sands deposits in the Uintah Basin in the United States. However, it is interesting to observe that gasoline yield (Figure 6b) increased from ∼8 wt % to 17 wt % as the pressure was increased. The explanation for this observation is not quite clear, but the pressure effect on gasoline yield may be due to the presence of compounds with low molecular weights produced during the hydrotreatment and just before the commencement of hydrocracking reactions. 3.3.3. Effect of LHSV on Petroleum Fractions and Gasoline Yield. The impact of the reaction time on product fractional distribution and gasoline yield was studied in terms of the liquid hourly space velocity by varying the LHSV over a range of 0.5-2.0 h-1 at a constant temperature and pressure (365 °C and 9.6 MPa, respectively). From Figures 7a and 7b, it is observed that both the liquid products and gasoline yield remained invariably the same throughout the range of space velocities. It can be inferred that prolonging the reaction time of hydrotreatment has minimal impact on the distribution of products and gasoline yield. Results from this study show that, to optimize gasoline production, the most important process variable that should be considered is the reaction temperature. 3.4. Kinetics Studies. Following the statistical study to determine the optimum conditions for maximum aromatic saturation, a separate set of experiments was performed to determine the apparent kinetic parameters controlling AHYD, HDS, and HDS. These experiments were performed by varying the reaction temperature over a range of 340-390 °C at the optimum LHSV and pressures conditions of 0.6 h-1 and 11.0 MPa, respectively. Because of the complexity of identifying all the
Aromatic Hydrogenation of Bitumen-Derived LGO
Energy & Fuels, Vol. 19, No. 5, 2005 1771 Table 7. Kinetic Parameters for Aromatic Hydrogenation over Sulfided NiMo/γ-Al2O3 Catalyst at a Pressure (P) of 11.0 MPa and a Liquid Hourly Space Velocity (LHSV) of 0.6 h-1 kinetic parametera
E [kJ/mol]
ln(k0 or K0)
R2
ln k ln KA ln KH2S ln KH
80.7 ( 4.0 26.7 ( 1.3 52.2 ( 2.6 8.4 ( 0.42
15.4 ( 0.77 -2.7 ( 0.14 -7.0 ( 0.35 -1.8 ( 0.09
0.9916 0.9837 0.9900 0.9044
a ln k ) ln k - E/(RT) and ln K ) ln K - E/(RT), where i is 0 i i0 either A, H2, or H2S.
Figure 8. Effect of temperature on aromatics, sulfur, and nitrogen conversions at the optimum pressure and LHSV conditions of 11.0 MPa and 0.6 h-1, respectively.
various forms and types of aromatics, sulfur, and nitrogen species occurring in the feedstock, total concentrations of the species were used. The concentration patterns for AHYD, HDS, and HDN activities in Figure 8 show that, for the temperature range selected for this study, AHYD covers both the kinetic and equilibrium domains. This is because, at low temperatures (340-380 °C), conversion of aromatics (AHYD) is kinetically controlled while at high temperatures (390 °C) aromatic saturation is dominated by thermodynamic equilibrium as the reverse reaction becomes increasingly more rapid. However, for HDS and HDN activities, the reactions (conversion) are kinetically controlled throughout the entire range. To provide a common basis for comparison for the kinetic studies, the hydrotreatment data for all the reactions were treated in the low-temperature regime (i.e., 340-380 °C), where the equilibrium effect is negligible. 3.4.1. Kinetics of Aromatic Hydrogenation. Using the total aromaticity data from the 13C NMR spectrometer, AHYD was kinetically modeled using a single-site mechanism form of the Langmuir-Hinshelwood (LH) rate of reaction equation. The LH model was used to determine any inhibition effects during hydrogenation. The selected rate equation was based on the following assumptions: the surface reaction was rate-limiting; equilibrium effects were negligible, hence, hydrogenation was irreversible and could be represented by a pseudo-first-order reaction; H2S produced from the HDS process inhibits the hydrogenation reaction; and aromatic species and H2S gas are both adsorbed onto the catalyst surface. The single-site rate equation was defined as
kKAKHPHCA dCA ) -rA ) dt 1 + KACA + KH2SPH2S
(9)
where -rA and k are the rate and apparent rate constant of AHYD respectively; KH, KA, and KH2S are the adsorption constants of hydrogen, aromatics, and H2S, accordingly, whereas CA is the concentration of the total residual aromatics, PH2S, and PH are the partial pressures of H2S and H2 gas, respectively, and t is the residence time. The partial pressure of H2S was included in eq 9 to account for any inhibition effect by the gas during saturation of aromatics.
Figure 9. Arrhenius and Van’t Hoff plots for aromatic hydrogenation. Conditions were as follows: a temperature of 340-380 °C, a pressure of 11.0 MPa, an LHSV of 0.6 h-1, and a H2/oil ratio of 550 mL/mL.
MAPLE 6.0 was used to evaluate the concentration of aromatics, which is defined as
[ { ]
[
CA ) (1 + KH2SPH2S) LambertW KA exp kKAKHPH t CAOKA + ln(CAo) + ln(CAOKH2SPH2S) kKAKHPH
}/
/(1 + KH2SPH2S)
]
(1 + KH2SPH2S) /KA (10) where
LambertW(x) ) 3 8 125 5 54 6 x - x2 + x3 - x4 + x - x + (o)7 (11) 2 3 4 5 and
{
[
x ) KA exp kKAKHPH t -
]
CAOKA + ln(CAo) + ln(CAOKH2SPH2S) kKAKHPH
}/
/(1 + KH2SPH2S)
(1 + KH2SPH2S) (12) The kinetic parameters in eq 10 were determined using a nonlinear least squares approach by fitting the experimental data to eq 10. High regression coefficients (Table 7) obtained from the Arrhenius and Van’t Hoff plots (Figure 9), together with the plot of the predicted versus actual conversion values of AHYD in Figure 10, show a good fit of the data to the model. The activation energy for hydrogenation, evaluated directly from the slope of the Arrhenius plot in Figure 9, was 80.7 kJ/
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Figure 10. Plot of predicted versus experimental data for the conversion of aromatic compounds. Conditions were as follows: temperatures of 340-380 °C, a pressure of 11.0 MPa, an LHSV of 0.6 h-1, and a H2/oil ratio of 550 mL/mL.
mol. This value is similar to the activation energy of 83 ( 3 kJ/mol obtained by Wilson and Kriz23 when they conducted a study on the hydrogenation of middle distillates over NiMo/γ-Al2O3. Although all the adsorption constants decreased as the temperatures increased, the adsorption constants for hydrogen sulfide (KH2S) were observed to be higher than that of the aromatic species (KA). This is an indication of the higher adsorption strength of H2S, as compared to the aromatics, thus leading to H2S inhibition during hydrogenation of aromatic compounds. This is because higher temperature reactions favor sulfur removal, thus leading to the formation of more H2S gas with higher partial pressures. Adsorption of the H2S gas can modify the catalyst surface by converting the active hydrogenation sites to hydrogenolysis centers and, in effect, reduce the overall rate of conversion of aromatics in the reaction.24 3.4.2. Kinetics of HDS and HDN. Unlike AHYD, the apparent kinetic parameters for HDS and HDN were best described by the irreversible pseudo-first-order power law model. The resulting rate equation for sulfur and nitrogen removal is defined as
-ri ) -
dCi ) kiCni dt
(13)
where Ci is the concentration of the heteroatom species; ri and ki are the rate of reaction and rate constants, respectively, and n is the order of reaction, with respect to either the sulfur or nitrogen species. HDS and HDN activities were assumed to occur in large hydrogen excess at constant partial pressure. The activation energies were also computed directly from the slopes of the regression lines in the Arrhenius plots (Figure 11), which predicted experimental data very well with higher R2 values (>0.99). Energy requirements of 42.7 and 27.5 kJ/mol were obtained for sulfur and nitrogen removal, respectively. Activation energies for HDS and HDN are low, in comparison with those reported in the literature.7,25-27 Most of the (23) Wilson, M. F.; Kriz, J. F. Fuel 1984, 63. (24) Zdrazil, M. Catal. Today 1988, 3, 269. (25) Ancheyta, J.; Angeles, M. J.; Macias, M. J.; Marroquin, G.; Morales, R. Energy Fuels 2002, 16, 189. (26) Kataria, K. L.; Kulkarni, R. P.; Pandit, A. B.; Joshi, J. B.; Kumar, M. Ind. Eng. Chem. Res. 2004, 43, 1373. (27) Kabe, T.; Akamatsu, K.; Atsushi, I.; Otsuki, S.; Godo, M.; Zhang, Q.; Qian, W. Ind. Eng. Chem. Res. 1997, 36, 5146.
Figure 11. Arrhenius plots for HDS and HDN. Conditions were as follows: temperatures of 340-380 °C, a pressure of 11.0 MPa, an LHSV of 0.6 h-1, and a H2/oil ratio of 550 mL/ mL.
existing data on the kinetics of HDS and HDN were derived using model compounds and HGO feeds. In this study, LGO that was derived from Athabasca bitumen was used. Thus, differences in the activation energies could be due to the different composition of the feedstocks used in present and past studies. It may be noted that heavy fractions contain less reactive (high-molecular-weight) sulfur and nitrogen species, whereas LGO fractions have more reactive (low-molecular-weight) heteroatom species, thus leading to a decrease in the slope of the Arrhenius curve, generating low activation energy. 4. Implication of Results Product quality is a very significant issue with hydrotreated products from distillates derived from bitumen such as the Athabasca oil sands. Results from this study suggest that reaction temperature and pressure are the key process parameters that affect the hydrogenation of aromatic compounds in the LGO feed, whereas HDS and HDN activities are greatly influenced by the interaction between temperature and LHSV. Hence, to adequately reduce the aromatic contents in middle distillates, high-pressure hydrotreatment at moderate reaction temperatures is recommended to essentially drive equilibrium toward the formation of saturated compounds (products). Because the CI of diesel fuel is affected by changes in the aromatic content and process conditions, the CI value of the fuel is also significantly affected by temperature and pressure. Under the conditions of hydrotreatment and MHC, the liquid product distribution and gasoline yield are highly influenced by temperature. Varying the reaction pressure and LHSV has minimal impact on the distribution of the liquid products. Although H2S enhances the hydrogenation activity by modifying the catalyst surface during hydrotreatment, higher concentrations of H2S produced from the HDS process may inhibit the hydrogenation of aromatic compounds in middle distillate fractions from Athabasca oil sands. Analysis of the HDS and HDN kinetics also revealed that, under the same hydrotreating conditions, sulfur species are more difficult to remove, in comparison to the nonrefractory nitrogen species. Hence, higher severities of hydrotreatment are required to produce distillates that have low
Aromatic Hydrogenation of Bitumen-Derived LGO
sulfur contents. The highest activation energy obtained for AHYD confirms that, under typical hydrotreatment conditions, AHYD is the most difficult, compared to the HDS and HDN activities. This may be due to the occurrence of a high concentration of refractory aromatic species (such as monoaromatics) in the feed. Because these groups of aromatics have stable resonance structures, more energy is usually required to hydrogenate the aromatic rings to their corresponding saturates.
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(7) The activation energies of 80.7, 42.7, and 27.5 kJ/ mol were obtained for AHYD, HDS, and HDN, respectively. AHYD has the highest activation energy, and this may be due to the occurrence of a high concentration of refractory aromatic species (such as monoaromatics) in the feed. Nomenclature Variables
5. Conclusions (1) Interaction between temperature and pressure is the most significant factor affecting aromatic hydrogenation (AHYD), whereas hydrodesulfurization (HDS) and hydrodenitrogenization (HDN) are greatly influenced by the interaction between temperature and space velocity. (2) The optimal conditions of AHYD for the maximum conversion of aromatic compounds was found at a temperature, pressure, and LHSV combination of 379 °C, 11.0 MPa, and 0.6 h-1, respectively. Under these conditions, the highest conversion of 63% could be attained. (3) Just as for AHYD, the cetane index (CI) also passes through a maximum, becuase of thermodynamic effects at high temperature severities of hydrotreatment. However, with increasing reaction pressure, the CI increases linearly. The highest CI value calculated from the ASTM D976 correlation was 47. This is only a 12% increase from the minimum specification of 42. (4) Among the operating variables considered in this study, temperature is the only parameter that has a significant impact on product distribution and gasoline yield. (5) The kinetic model developed for AHYD, using the single site mechanism of the Langmuir-Hinshelwood (LH) approach, predicted the experimental data well. (6) The higher adsorption constant of hydrogen sulfide (H2S), compared to that of the aromatics, shows that the hydrogenation of aromatics in the feed is inhibited by H2S.
CA ) concentration of aromatics (%) CAO ) concentration of aromatics in the feed (%) Ki ) adsorption constant with respect to species i (either aromatics, hydrogen sulfide, or hydrogen) ki ) rate constant of species i (h-1) LHSV ) liquid hourly space velocity (h-1) n ) order of reaction P ) pressure (MPa) PH2 ) partial pressure of hydrogen PH2S ) partial pressure of hydrogen sulfide R2 ) regression coefficient T ) temperature (°C) Abbreviations AHYD ) aromatic hydrogenation A ) aromatic compound Ar-H ) saturated aromatic compound CI ) cetane index CN ) cetane number H2S ) hydrogen sulfide gas HDN ) hydrodenitrogenation HDS ) hydrodesulfurization HGO ) heavy gas oil HT ) hydrotreated sample LGO ) light gas oil NH3 ) ammonia gas PG ) pressure gauge TC ) temperature controller VGO ) vacuum gas oil EF040080I
