Article pubs.acs.org/EF
Vapor−Liquid Equilibrium (VLE)-Based Modeling for the Prediction of Operating Regimes in a Heavy Gas Oil Hydrotreater Mugurel Catalin Munteanu and Jinwen Chen* CanmetENERGY, Natural Resources Canada, One Oil Patch Drive, Devon, Alberta T9G 1A8, Canada ABSTRACT: In a pilot-plant trickle-bed hydrotreating reactor, ideal operating regimes (plug flow, full catalyst wetting, and absence of reactor wall effects) are desirable to generate reliable, reproducible, and representative data for kinetics studies and commercial scale-up. In this study, the operating regimes of a heavy gas oil pilot-plant hydrotreating reactor were investigated under operating conditions of commercial interest using real petroleum feedstocks. Vapor−liquid equilibrium experiments and calculations were performed to accurately predict the fluid flows and the physical/transport properties of the vapor and liquid phases. These properties were subsequently used to predict the operating regimes under various operating conditions and to generate a map of operating conditions within which the desired operating regimes can be maintained.
1. INTRODUCTION The performance of a trickle-bed hydroprocessing reactor largely depends upon process operating conditions, which not only affect reaction kinetics and thermodynamics but also determine the fluid hydrodynamics and operating regime inside the reactor. With the current market evolution toward an increased demand for light oil products, such as diesel fuels, and the generally declining quality of crudes, refiners will have to improve their processing units for upgrading heavy oil and residual feedstocks. 1 Hydrotreating (HDT) processes in petroleum refineries are mainly based on trickle-bed reactors (TBRs) operated under trickle flow conditions. Most commercial TBRs run adiabatically at high temperatures and pressures and generally involve exothermic reactions, such as hydrodesulfurization, hydrocracking, hydrodemetallization, and hydrodenitrogenation. Under commercial operating conditions, HDT reactors or hydrotreaters are essentially in a state of vapor− liquid equilibrium (VLE). Hydrogen and petroleum feed/ product are present in both the liquid and vapor phases at the same time. More hydrocarbons are evaporated into the vapor phase if the temperature is increased, the pressure is decreased, or the hydrogen flow rate is augmented. The ratio of the liquidand vapor-phase flow rates in the reactor can alter the hydrodynamics and chemical reactions, which ultimately affect reactor performance. Advanced knowledge of VLE effects on reactor behavior can lead to an improved understanding and control of HDT processes. This knowledge can also be a useful tool in predicting reactor behavior and deviations from the optimal or expected operating regime. It is generally understood that the ideal operating regime for reactors includes the plug flow of the reactants, full catalyst wetting to maximize contact between reactants and catalyst particles, and minimal reactor wall effects. In the literature, HDT is commonly discussed in terms of the following variables: pressure, temperature, liquid hourly space velocity (LHSV), and gas/oil ratio.2−5 VLE effects in threephase reactors are rarely discussed,6−10 and a very limited number of studies even report VLE effects in petroleum hydroprocessing.11,12 The dearth of research on VLE during Published 2011 by the American Chemical Society
petroleum hydroprocessing is due to the difficulty of experimental measurements (or simulations) and poor definition of the problem because of the complexity of the system. Most studies focus on systems that consist of hydrogen and a single hydrocarbon or a replicated hydrocarbon mixture.13−18 A number of studies have also been conducted on systems consisting of hydrogen and authentic feedstocks.19−25 Flow dynamics and operating regimes in a middle-distillate pilotplant hydrotreater were discussed in a previously published paper.25 The present paper discusses a related study of heavy-distillate hydrotreaters. VLE in pilot-plant hydrotreaters was predicted under various operating conditions using a flash calculation program calibrated in-house. The calculated VLE results were used to predict flow dynamics in the hydrotreaters to investigate whether they were operating under the desired operating regime (plug flow, full catalyst wetting, and absence of reactor wall effects). Hydrotreater operating regimes for typical heavy-distillate feed and various operating conditions and parameters are described and defined in detail.
2. BASIS OF CALCULATIONS 2.1. VLE Experiments and Calculations. In a paper published earlier,24 we reported VLE experimental findings obtained at different temperatures and pressures for three heavy gas oils (HGOs) of different chemical compositions. The VLE data were then used to estimate the interaction coefficients between hydrogen and hydrocarbon pseudo-components used in VLE flash calculations. The interaction coefficients were further correlated with the boiling point of the hydrocarbon pseudo-component and the aromatic content in the HGO. With the established correlations, VLE can be calculated with reasonable accuracy for any HGO HDT system, as long as the simulated distillation (SimDis), density, and molecular-weight distributions by boiling point are known. VLE calculation for petroleum hydroprocessing systems is normally based on equations of state (EOS). A number of hydrocarbon pseudocomponents (29 in this study) were derived from the SimDis data of the feed being flashed. The densities, molecular-weight distributions by boiling point, and interaction coefficients between hydrogen and Received: October 17, 2011 Revised: November 22, 2011 Published: December 6, 2011 1230
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hydrocarbon pseudo-components were estimated using established correlations. A more detailed description of the flash calculation has been provided elsewhere.25 2.2. System Description. In this study, VLE calculations and determination of operating regimes in trickle-bed HDT reactors were carried out using an in-house pilot-plant hydrotreater. Such units are commonly used in HDT catalyst development, evaluation, and activity stability studies to generate reliable data for commercial scale-up. The specifications of the pilot-plant hydrotreater used in this study are given in Table 1. The properties of the HGO feed are given in Table 2.
be directly scaled up because commercial TBRs normally operate in plug flow mode because of high liquid throughput. Plug flow is more difficult to achieve at high conversions. A simple empirical relation between the reactor bed length and effective particle size is frequently used to verify conditions for plug flow27 (the symbols used here are explained in the Nomenclature).
LB > 100 d pe
Mears proposed a plug flow criterion based on an axial dispersion model with a number of mixers in series.
LB 20n 1 > ln d pe Pe L (1 − α)
Table 1. Pilot-Plant Hydrotreater Specification and Operating Conditions
(2)
In deriving this criterion, the authors assume that the minimum bed length to eliminate axial dispersion should be within 5% of the bed length required for plug flow. In addition, the use of this criterion is limited to reactors with conversions of 90% or less. This criterion was considered to be too conservative and, in most cases, was difficult to meet. Gierman29 modified the criterion by relaxing the deviation in bed length required by plug flow from 5 to 10%.
Specifications of the Pilot Plant reactor ID (cm) 2.54 catalyst volume (mL) 100−200 catalyst particle size (average) 1.5 × 5.0 (mm) diluting particle size (mm) 0.2−0.4 catalyst/diluent volumetric ratio 1:1 Operating Conditions Used Were Typical of Commercial Heavy-Distillate HDT temperature (°C) 350, 360, 370, 380, 390, 400, 410, and 420 pressure (bars) 50, 60, 70, 80, and 90 LHSV (h−1) 1.5 gas/oil ratio (NL/kg) 500, 800, and 1000
LB 8n 1 > ln d pe Pe L (1 − α)
(3)
An even more relaxed criterion with 15% deviation from plug flow was used in our previous and current computations.25
LB 20 n 1 > ln d pe Pe L (1 − α)
Table 2. Physical Properties and Chemical Compositions of the HGO Feed density at 15.6 °C (g/cm3) (ASTM D4052) paraffins (wt %) cycloparaffins (wt %) aromatics (wt %) boiling points (°C) IBP 10 wt % 30 wt % 50 wt % 70 wt % 90 wt % FBP
(1)
28
(4)
The Peclet number PeL in the above equations is based on the catalyst particle diameter. In the present study, the Hochman and Effron correlation30 was used to calculate PeL from the particle-based Reynolds number
0.9609 30.94 7.86 61.2
Pe L = 0.034Re L 0.53
(5)
where 201.8 292.6 350.2 386.4 445.4 529.6 702.6
Re L =
d peuLρ L μL
(6) 30
Note that, in the original Hochman and Efron correlation, there is a term related to the gas flow Reynolds number. In the present work, the term was calculated to have a value close to 1 for any gas rate used in this study and, therefore, was neglected. In this paper, the effective particle size (dpe) was calculated on the basis of the volumetric average of the catalyst particle (cylindrical shape) and the diluent particle (spherical shape) in the reactor with the equation below
2.3. Operating Regime Criteria. Deviations from plug flow, incomplete catalyst wetting, and reactor wall effects are frequently encountered in bench-scale and pilot-plant reactor operations of threephase catalyst reaction systems. Such conditions can lead to inconsistent and nonreproducible experimental data, resulting in unreliable/biased data interpretation and measurements of catalyst activity, reaction kinetics, and scale-up. Therefore, deviations from the ideal operating regime during experiments should be avoided by controlling reactor operating conditions, such as temperature, pressure, gas/oil ratio, LHSV, catalyst particle/diluting particle size, and catalyst bed length and diameter.25 Many other papers have been published that discuss these important issues. A review paper summarizing the criteria that must be met to ensure ideal behaviors in TBRs was published recently.26 The criteria presented in our previous study based on a middle-distillate hydrotreater were also used in the present study. These criteria are briefly discussed below. 2.3.1. Plug Flow Criterion. The plug flow regime (PFR) of the liquid phase in a pilot-plant TBR is always desirable because any axial dispersion or back-mixing in the reactor can complicate the interpretation of kinetics data. In addition, the generated data cannot
⎛ 3L d 2 + 2d 3 ⎞1/3 d ⎟ d pe = ⎜⎜ c c ⎟ 4 ⎝ ⎠
(7)
where Lc and dc are the cylindrical catalyst particle length and diameter, respectively, and dd is the diluent particle diameter. Mary and Chaouki31 published a detailed review considering the homogeneity of the catalyst bed, catalyst wetting, axial dispersion, channeling, isothermicity, and mass-transfer effects on TBRs. Different criteria published in the literature were discussed to minimize the influence of these factors. 2.3.2. Full Catalyst Wetting Criterion. Full catalyst surface wetting in TBRs is critical to maximize catalyst use and performance. Under full catalyst wetting, all of the catalyst surfaces (internal and external) are covered by the flowing liquid. Any partial catalyst wetting could result in bypassing of the liquid feed, thus impairing catalyst performance. In a pilot-plant TBR, the liquid superficial velocity is normally very low because of limitations of reactor configuration and 1231
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operating conditions. Partial vaporization of liquid hydrocarbons under high temperature and hydrogen flow rate in hydrotreaters further reduces the liquid superficial velocity. Therefore, incomplete catalyst wetting is another major concern in operating pilot-plant trickle-bed HDT reactors. In this work, the criterion for wetting efficiency proposed by Gierman29 was used
μLuL ρ Ld pe2g
significant difference in SimDis between the original feed and the blend, while the corrected SimDis is almost identical to the SimDis of the blend, indicating that the developed procedure provides reasonable corrections of SimDis to account for thermocracking at temperatures above 390 °C. 3.2. Plug Flow Operating Regime. The operating regimes that meet the plug flow criterion were calculated using flash calculation programs developed in-house, as well as the operating conditions, reactor configurations, and eqs 4−6. Figure 2 shows the temperature and pressure required to meet
> 5.0 × 10−6 (8)
2.3.3. Absence of Reactor Wall Effects Criterion. In commercial TBRs, the reactor diameter is much larger than the catalyst particle size and reactor wall effects are, therefore, negligible. However, in a pilot-plant or bench-scale reactor, if the catalyst particles are not small enough and/or if no inert diluting particles are used in the catalyst bed, reactor wall effects may be present and sometimes may be significant. To minimize reactor wall effects, the following criterion has to be met:32
DR > 25 d pe
(9)
In the present study, DR/dpe = 49.7 ≫ 25. Therefore, the reactor wall effect is negligible, and no further discussion is given.
3. RESULTS AND DISCUSSION 3.1. Accounting for Thermocracking. During VLE flash experiments, it was observed that, at temperatures above 390 °C, the HGO feed underwent thermocracking, demonstrated by monitoring of the outlet gases from the VLE cell. It was found that the concentrations of light hydrocarbons (C1−C6) in the outlet gases increase with an increasing temperature. Another test performed to verify this observation was to compare the SimDis data of the original feed to those of the proportional blend of the top and bottom fractions from the VLE cell. It was found that the blend became lighter with increasing the temperature because of the thermocracking that occurred in the process. To account for this change, a corrective procedure was developed and implemented prior to the VLE flash calculation. At temperatures above 390 °C, the SimDis data of the feed were corrected to account for thermocracking. The corrections were different for different temperatures because of variations in thermocracking conversion. The corrected SimDis data were then used to derive the 29 hydrocarbon pseudo-components. Figure 1 presents the SimDis data of the original feed, the blend
Figure 2. Plug flow operating regime mapping, temperature−pressure− gas/oil ratio.
the plug flow criterion when other operating and reactor configuration parameters (gas/oil ratio, LHSV, bed length, particle size, etc.) are fixed. In the figure, each individual curve divides the two-dimensional plane into two zones. The zone above the curve, which represents relatively low operating temperatures and high pressures, meets the plug flow criterion, while the zone below the curve, which represents relatively high temperatures and low pressures, does not. As seen in Figure 2, an increased reactor temperature requires increased pressure to ensure plug flow of the liquid phase. This is understandable because elevated temperature tends to volatilize more oil into the vapor phase to reduce the liquid Peclet number, while at the same time, increased pressure tends to push the vaporized oil back into the liquid phase to maintain a constant Peclet number. At a given temperature or pressure, an increased gas/ oil ratio requires increased pressure or decreased temperature to maintain liquid plug flow, because a higher gas/oil ratio tends to increase the vaporization of oil. During pilot-plant studies, the operating conditions, catalyst particle size, and diluting particle size are determined according to the commercial process being simulated or the operating limitations of the pilot-plant unit. In this case, the catalyst volume or the catalyst bed length becomes important to maintain plug flow of the liquid inside the reactor. A minimum catalyst bed length is required under a particular set of operating conditions. Figure 3 shows the required catalyst bed length at different reactor temperatures and pressures for a fixed LHSV of 1.5/h and gas/oil ratio of 800 NL/kg. Again, each curve in the figure represents the boundary of the temperature and catalyst bed length that determines whether the plug flow criterion is met, while other operating conditions are fixed. The zone above the curve, which represents relatively
Figure 1. Experimental versus estimated SimDis at T = 420 °C.
of the top and bottom fractions from the CLE cell, and the corrected SimDis data at 420 °C. As seen in the figure, there is 1232
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boundary between two zones of temperature and diluting particle size. The zone under the curve represents the temperature and diluting particle size that meet the plug flow criterion, while the zone above the curve represents those that do not meet the plug flow criterion. As seen in this figure, an increasing temperature requires smaller diluting particles to ensure plug flow of the liquid, as readily explained by eqs 3 and 4. Similarly, for a given temperature, reduced pressure also requires smaller diluting particles to maintain liquid plug flow. As mentioned earlier, the plug flow criterion used in this study is dependent upon reaction conversion (in this study, we consider the hydrodesulfurization conversion), as shown in eq 4. Higher conversion means increased difficulty in meeting the plug flow criterion. From the standpoint of pilot-plant operation, this requires a high temperature and LHSV, low pressure and gas/oil ratio, longer catalyst bed, and smaller diluting particles. Figure 5 presents the required temperature
Figure 3. Plug flow operating regime mapping, temperature−pressure− catalyst bed length.
low temperatures and longer catalyst beds, meets the plug flow criterion, while the zone below the curve, which represents relatively high temperatures and shorter catalyst beds, does not. As seen in this figure, an increased temperature requires increased catalyst bed length to ensure plug flow of the liquid because an increased temperature reduces the liquid Peclet number, resulting in higher required LB/dpe, as shown in eqs 3 and 4. At a higher temperature and lower pressure, the required length of the catalyst bed increases sharply. In most pilot-plant studies, the operating conditions and the catalyst particle size are fixed according to test requirements. Therefore, inert diluting particles (such as glass beads and silicon carbide particles) are used to maintain plug flow of the liquid inside the reactor, improve liquid distribution, and achieve better heat transfer for isothermal operations. Mathematically, the effective particle diameters (dpe) in eqs 3 and 4 are calculated on the basis of the catalyst particle size and the diluting particle size. The smaller the diluting particle, the smaller the calculated effective particle size for a given catalyst particle size. Therefore, a certain diluting particle size is required under a particular set of operating conditions. Figure 4
Figure 5. Plug flow operating regime mapping, temperature− conversion−pressure.
and pressure at various hydrodesulfurization conversions ranging between 80 and 95%. It should be pointed out that the conversion was fixed at 90% for all of the calculations discussed in this paper, except for those presented in Figure 5. The zone above each individual curve in Figure 5 represents the temperatures and pressures under which the plug flow criterion is met, while the zone below the curve represents those under which the plug flow criterion is not met. As the figure shows, higher conversion requires either higher pressure or lower temperature (or both) to maintain liquid plug flow in the reactor. The pressure requirement does not vary much as the conversion increases from 80 to 90%. However, when the conversion increases from 90 to 95%, the required pressure increases sharply. Another set of calculations were performed to evaluate the minimum LHSV required to ensure liquid plug flow. Figure 6 depicts the required LHSV and temperatures for liquid plug flow at different pressures. Again, for each individual curve, the zone above the curve represents the operating LHSV and temperatures that can meet the plug flow criterion, while the zone below the curve represents those that cannot. As seen in Figure 6, at low temperatures (350−380 °C), there are no significant differences between the LHSVs required at different pressures. In this temperature range, the required LHSV is almost constant for a given pressure. This is no longer the case when the temperature is above 400 °C. In this case,
Figure 4. Plug flow operating regime mapping, temperature−diluting particle size−pressure.
shows the required diluting particle size at different reactor temperatures and pressures for a fixed LHSV of 1.5/h and gas/ oil ratio of 800 NL/kg. Each curve in the figure represents the 1233
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cannot. As seen in the figure, at fixed gas/oil ratios, an increased reactor temperature requires increased pressure to ensure full catalyst wetting. The explanation is similar to that for the plug flow, as discussed earlier. At a given temperature, an increased gas/oil ratio also requires increased pressure to maintain full catalyst wetting. It is interesting to note by comparing Figures 2 and 7 that, at a given temperature and gas/oil ratio, the pressure required to achieve complete catalyst wetting is lower that that required for plug flow. It seems that, for the present system, the full catalyst wetting criterion is easier to meet or more relaxed than the plug flow criterion. A similar observation has been made for middle-distillate hydrotreaters.25 For the present system, as long as the plug flow criterion is met, full catalyst wetting is achieved automatically. However, this conclusion does not necessarily hold when the liquid properties, reactor diameter, and/or catalyst/diluting particle sizes change. For example, in a commercial middle-distillate hydrotreater, the plug flow criterion is easier to meet than the full catalyst wetting criterion.33 The diluting particle size also plays an important role in catalyst wetting, as seen in eq 8. The smaller the diluting particles, the easier it is to achieve full catalyst wetting. Therefore, under certain conditions, the diluting particle size must remain below a certain maximum to avoid partial catalyst wetting. Figure 8
Figure 6. Plug flow operating regime mapping, temperature−LHSV− pressure.
the required LHSV increases dramatically when the pressure decreases. In addition, the required LHSV also increases significantly when the temperature increases. At a given temperature, a lower pressure requires higher LHSV to meet the plug flow criterion. 3.3. Full Catalyst Wetting Regime. The liquid-phase flow rate in a pilot-plant hydrotreater can be significantly affected by partial volatilization of the oil into the vapor phase. A reduction in the liquid flow rate could result in incomplete catalyst wetting and the bypassing of the liquid feed, leading to an impaired reactor performance. In addition, partial catalyst wetting could also promote the formation of a hot spot inside the catalyst bed, resulting in catalyst damage. It is therefore essential to ensure a fully wetted catalyst bed in a pilot-plant hydrotreater throughout the entire testing process to generate reliable and reproducible data. As seen in eq 8, the full catalyst wetting criterion is independent of the catalyst bed length, unlike the plug flow criterion (eqs 3 and 4). Only the operating conditions and the catalyst and diluent particle sizes determine whether or not full catalyst wetting is achieved. Figure 7 shows the mapping of the required temperatures and pressures to ensure full catalyst wetting at different gas/oil
Figure 8. Full catalyst wetting operating regime mapping, temperature− diluent particle size−pressure.
shows the maximum diluting particle size at different reactor temperatures and pressures with fixed LHSV of 1.5/h and gas/ oil ratio of 800 NL/kg. For each curve in the figure, the zone under the curve represents the diluting particle size and temperature that meet the full catalyst wetting criterion, while the zone above the curve represents those that do not. An increased temperature and/or a decreased pressure require smaller diluting particles to ensure full catalyst wetting, similar to the plug flow requirement discussed earlier. As seen in eq 8, the value of the left-hand side is more sensitive to the effective particle size (dpe) than to other parameters (velocity, density, viscosity, etc.) because of the power of 2. This might also explain why the full catalyst wetting criterion is more relaxed than the plug flow criterion in this study, because the Peclet number is proportional to the 0.53 power of effective particle size (eqs 3 and 4). In this study, the diluting particle size (0.3 mm) is so small that full catalyst wetting would not become a problem. However, if a larger diluting particle size had been chosen for the calculation,
Figure 7. Full catalyst wetting operating regime mapping, temperature− pressure−gas/oil ratio.
ratios. The zone above the curve represents the operating temperatures and pressures that meet the full catalyst wetting criterion, and the zone below the curve represents those that 1234
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the operating regimes in a pilot-plant HGO hydrotreater to determine under what operating conditions and with what reactor configurations the reactor could be operated in the desired operating regimes (plug flow, full catalyst wetting, and absence of reactor wall effects). The calculation results show that liquid plug flow and full catalyst wetting inside the HGO HDT reactor are favored by a higher pressure and LHSV, lower gas/oil ratio and temperature, smaller diluting particles, and longer catalyst beds. A higher desulfurization conversion also requires more stringent controls to meet the desired operating criteria. It should be pointed out that the predictions cannot be verified by performing experiments because an accurate measurement of actual gas and liquid flow rates in a hydrotreater operated under high temperature and pressure is extremely difficult (if not impossible). However, the VLE model used in this study was indeed calibrated and validated with experimental data obtained by conducting flash experiments with HGO and hydrogen under similar operating conditions. The flash experiments are considered to represent the VLE behavior in a TBR by assuming that VLE is achieved anywhere inside the reactor. In addition, the criteria used for determining the plug flow and full catalyst wetting have been well-established, used, and recommended by researchers in this field. Therefore, the methodology established in this work can be a useful tool in predicting the actual gas and liquid flow rates in a TBR and the deviation from the desired ideal operating regimes.
the full catalyst wetting criterion might have superseded the plug flow criterion. Figure 9 shows the mapping of full catalyst wetting at different LHSVs. Each curve in the figure represents the
Figure 9. Full catalyst wetting operating regime mapping, temperature− LHSV−pressure.
boundary of two zones of temperature and pressure for a certain LHSV. The zone above the curve represents the operating temperatures and pressures that meet the full catalyst wetting criterion, while the zone below the curve represents those that do not. Under fixed LHSV, an increased reactor temperature requires increased pressure to ensure full catalyst wetting. Alternatively, at a given temperature, a decreased LHSV requires increased pressure to maintain full catalyst wetting. Calculations were also performed to evaluate the minimum LHSV required to meet the full catalyst wetting criterion. Figure 10 depicts
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AUTHOR INFORMATION Corresponding Author *Telephone: 780-987-8763. Fax: 780-987-5349. E-mail:
[email protected].
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ACKNOWLEDGMENTS The authors are grateful to Dennis Carson and other CanmetENERGY pilot-plant staff for conducting the VLE experiments, the analytical lab staff for sample analysis, and Conrad Gietz for editing this paper. Partial funding for this study was provided by the Canadian Interdepartmental Program of Energy Research and Development (PERD 1.1.3).
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Figure 10. Full catalyst wetting operating regime mapping, temperature− LHSV−pressure.
the calculation results. At the lower temperatures (350−380 °C), there is no significant difference between the LHSVs at which the full wetting criterion is met. However, when the temperature increases to 400 °C and beyond, the minimum required LHSV for full catalyst wetting increases substantially.
4. SUMMARY In this study, VLE calculations were performed with a HGO HDT system. The calculated results were then used to predict 1235
NOMENCLATURE BP = boiling point (°C) Caromatics % = aromatic content of feed (wt %) dC = catalyst particle diameter (cm) dd = diluting particle diameter (cm) dpe = effective particle size (cm) DR = diameter of the reactor bed (cm) EOS = equation of state g = gravitational acceleration (981 cm/s2) HDT = hydrotreating HGO = heavy gas oil LB = reactor bed length (cm) LC = catalyst particle length (cm) LHSV = liquid hourly space velocity (h−1) n = reaction order P = pressure (bars) PeL = particle Peclet number PFR = plug flow regime ReL = particle Reynolds number T = temperature (°C) TBR = trickle-bed reactor uL = liquid superficial velocity (cm/s) dx.doi.org/10.1021/ef201616f | Energy Fuels 2012, 26, 1230−1236
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(31) Mary, G.; Chaouki, J.; Luck, F. Int. J. Chem. React. Eng. 2009, 7, 1−70. (32) Chu, C. F.; Ng, K. M. AIChE J. 1989, 35, 148−158. (33) Chen, J.; Mulgundmath, V.; Wang, N. Ind. Eng. Chem. Res. 2011, 50, 1571−1579.
VLE = vapor−liquid equilibrium Suffixes L = liquid phase Greek Symbols α = conversion of the component μL = liquid-phase viscosity (cP) ρL = liquid-phase density (g/cm3)
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REFERENCES
(1) Trambouze, P. Engineering of hydrotreating processes. In Chemical Reactor Technology for Environmentally Safe Reactors and Products; Plenum Press: New York, 1993; NATO Advanced Study Institute Series E, pp 425−450. (2) Speight, J. G. The Desulfurization of Heavy Oils and Residua; Marcel Dekker, Inc.: New York, 1981. (3) Bej, K.; Ajay, D.; John, A. Energy Fuels 2001, 15, 377−383. (4) Botchwey, C.; Ajay, D.; John, A. Energy Fuels 2003, 17, 1372− 1381. (5) Mapiour, M.; Sundaramurthy, V.; Dalai, A. K.; Adjaye, J. Fuel 2010, 89, 2536−2543. (6) Tesser, R.; Di Serio, M.; Santacesaria, E. Catal. Today 2003, 79 and 80, 323−331. (7) LaVopa, V.; Satterfield, C. N. Chem. Eng. Sci. 1988, 43, 2175− 2180. (8) Akgerman, A.; Collins, G. M.; Hook, B. D. Ind. Eng. Chem. Fundam. 1985, 24, 398−401. (9) de Jong, K. P. Ind. Eng. Chem. Res. 1994, 33, 3141−3145. (10) Smith, C. M.; Satterfield, C. N. Chem. Eng. Sci. 1986, 41, 839− 843. (11) Bellos, G. D.; Papayannakos, N. G. Catal. Today 2003, 79 and 80, 349−355. (12) Kocis, G. R.; Ho, T. C. Chem. Eng. Res. Des. 1986, 64, 288−291. (13) Wiegand, K.; Strobel, B.; Hofmann, H. Chem. Eng. Technol. 1989, 12, 280−288. (14) Laugier, S.; Richon, D.; Renon, H. Proceedings of the Thermodynamics Sessions of the 2nd World Congress on Chemical Engineering; Montreal, Quebec, Canada, 1983. (15) Simnick, J.; Sebastian, H.; Li, H.-M.; Chao, K.-C. J. Chem. Eng. Data 1978, 23, 339−340. (16) Chen, J.; Yang, H.; Ring, Z. Proceedings of the 2005 AIChE National Spring Meeting, Topical A: 8th Topical Conference on Refinery Processing; Atlanta, GA, April 10−14, 2005. (17) Connolly, J.; Kandalic, G. J. Chem. Eng. Data 1986, 31, 396− 406. (18) Ramanujam, S.; Leipziger, S.; Well, S. Ind. Eng. Chem. Process Des. Dev. 1985, 24, 364−368. (19) Chen, J.; Te, M.; Yang, H.; Ring, Z. Pet. Sci. Technol. 2003, 21 (5and 6), 911−935. (20) Chen, J.; Ring, Z. Fuel 2004, 83, 305−313. (21) Chen, J.; Yang, H.; Ring, Z. Catal. Today 2004, 98, 227−233. (22) Peng, D. Y.; Robinson, D. B. Ind. Eng. Chem. Fundam. 1976, 15, 59−64. (23) Chen, J.; Mclean, N.; Briker, Y.; Hager, D.; Ring, Z. Proceedings of the 2007 AIChE National Spring Meeting, Topical A: 10th Topical Conference on Refinery Processing; Houston, TX, April 22−27, 2007; Paper 3d. (24) Munteanu, M. C.; Chen, J.; Farooqi, H. Hydrocarbon Eng. 2010, 15, 39−47. (25) Chen, J.; Wang, N.; Mederos, F.; Ancheyta, J. Ind. Eng. Chem. Res. 2009, 48, 1096−1106. (26) Mederos, F.; Ancheyta, J.; Chen, J. Appl. Catal., A 2009, 355, 1−19. (27) Carberry, J. J.; Wendel, M. M. AIChE J. 1963, 9, 129−133. (28) Mears, D. E. Chem. Eng. Sci. 1971, 26, 1361−1366. (29) Gierman, H. Appl. Catal. 1988, 43, 277−286. (30) Hochman, J.; Effron, E. Ind. Eng. Chem. Fundam. 1969, 8, 63−71. 1236
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