ARTICLE pubs.acs.org/EF
Kinetics of Catalytic Pyrolysis of Heavy Gas Oil Derived from Canadian Synthetic Crude Oil Xianghai Meng, Chunming Xu, Li Li, and Jinsen Gao* State Key Laboratory of Heavy Oil Processing, China University of Petroleum, Beijing, 102249, China ABSTRACT: The catalytic pyrolysis of heavy gas oil derived from Canadian synthetic crude oil on zeolite catalyst was conducted in a confined fluidized bed reactor. The optimal reaction temperature, weight hourly space velocity, steam-to-oil weight ratio, and catalyst-to-oil weight ratio were 660 °C, 9.5 h1, 0.6, and 20, respectively. The yield of total light olefins was 36.4 wt % under these optimal conditions. The overall reactants and products were classified into seven species, and a seven-lump kinetic model with 15 rate constants and a catalyst deactivation constant was proposed. Kinetic constants at 620, 640, 660, and 680 °C were estimated by a nonlinear least-squares regression method. Preexponential factors and apparent activation energies were then calculated according to the Arrhenius equation. The feed lump had larger rate constants and smaller apparent activation energies than did the intermediate product lumps. The seven-lump kinetic model showed high calculation precision and the predicted yields agreed well with the experimental values.
1. INTRODUCTION Light olefins (ethene, propene, and butene) are considered as the backbone of the petrochemical industry. Steam cracking (SC) is the most common technology for the production of light olefins,1 and is performed at a high temperature and a short residence time. Fluid catalytic cracking (FCC) has been recently used to produce partly light olefins. Such process involves use of modified ZSM-5 or other novel zeolite-containing catalysts,2,3 recycling FCC naphtha to the main or secondary riser,4,5 operation under severe conditions,6 and modification of the reactor type.6 Catalytic pyrolysis has been recognized as a promising alternative route for the production of light olefins. Compared with SC, catalytic pyrolysis has several advantages such as higher yields of light olefins, easier control of olefin distribution, lower energy consumption, and wider feed scope. The feeds of catalytic pyrolysis include butane,7,8 butene,9,10 C4 hydrocarbons,11 hexene,12,13 C4+ olefins,14 heptane,15 octane,16 natural gasoline,17 naphtha,18,19 coker naphtha,20 FCC naphtha,2022 gas oil,23 heavy oil,24,25 waste tire,26 plastic mixture,27 and bio-oil.28 Catalytic pyrolysis involves a series of complicated reactions,11 and a kinetic study is important to understand the reactions, design and simulate the reactor, predict reaction performance, and optimize operating conditions. The complex mixtures can be described by grouping a large number of chemical compounds into groups of pseudocomponents according to their boiling points and molecular characteristics. Several lumping kinetic models are reported for catalytic pyrolysis of propane,29 C4 hydrocarbons,30 gas oil,23 and even heavy oils.3133 Canadian oilsand bitumen, an unconventional petroleum resource with abundant reserves, has become increasingly important with increasing petroleum consumption. Synthetic crude oil has always been a significant product of oilsand bitumen. Heavy gas oil (HGO) derived from Canadian synthetic crude oil is a potential feed in catalytic pyrolysis. The main objectives of the present study were to investigate the cracking performance of HGO and to describe its reactions through a lumping kinetic model. r 2011 American Chemical Society
2. EXPERIMENTS 2.1. Feed and Catalyst. HGO derived from Canadian synthetic crude oil was used in the present research; its main properties are listed in Table 1. The properties show that HGO is a naphthenic fraction. Compared with the properties of a typical paraffinic fraction, Chinese Daqing vacuum gas oil,24 the density, carbon residue, and aromaticity of HGO were higher, while the molecular weight and H/C atomic ratio were lower. Daqing vacuum gas oil shows good cracking performance and good selectivity of light olefins. The cracking performance of HGO is expectant. A kind of ZSM-5 zeolite catalyst for catalytic pyrolysis of naphthenicbase fractions developed by China University of Petroleum34 was used. Its main properties are listed in Table 2. 2.2. Apparatus. The experiments were conducted in a confined fluidized bed reactor, the apparatus diagram of which is shown in Figure 1. The apparatus consisted of five sections, namely, oil and steam input mechanisms, a reaction zone, a temperature control system, and a product separation and collection system. Experiments were carried out in batches. In each experiment, 50 g of the catalyst was loaded into the reactor that had an effective volume of about 580 cm3. Distilled water was pumped into a furnace to generate steam, which was used to fluidize the catalyst and then mixed with the pumped feedstock. The mixture was heated to approximately 500 °C in a preheater and then entered the reactor. Reactions took place as the feed contacted the fluidized catalyst. After the reaction, the oil gas was cooled and separated into liquid and gas samples by the product separation and collection system. Afterward, the spent catalyst was drawn out of the reactor by a vacuum pump. 2.3. Analytical Methods. The gas sample was analyzed using an Agilent refinery gas analyzer (by ASTM D1945, D1946, and UOP 539 standard methods) to determine the volume percentage of the components. Data were converted to mass percentages using the state equation of ideal gases. The liquid sample was analyzed by simulated distillation gas chromatography (AC SIMDIS HT 750; Analytical Controls, Inc.) to Received: April 9, 2011 Revised: June 20, 2011 Published: June 22, 2011 3400
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obtain the weight percentage of gasoline (C5200 °C), diesel (200350 °C), and heavy oil (>350 °C). The coke content of the spent catalyst or inert carrier was measured with a coke analyzer.
3. CRACKING PERFORMANCE OF HGO The effects of reaction temperature, weight hourly space velocity, steam-to-oil weight ratio, and catalyst-to-oil weight ratio on product yields were investigated. In the present study, feed conversion was defined as the sum of the yields of dry gas, liquefied petroleum gas (LPG), gasoline, and coke. 3.1. Effect of Reaction Temperature. The effect of reaction temperature (600700 °C) on conversion and product yields was investigated, keeping the weight hourly space velocity, steam-to-oil weight ratio, and catalyst-to-oil weight ratio constant at 9.5 h1, 0.6, and 17, respectively. Table 3 lists the feed conversion and yields of the products. As the reaction temperature increased, feed conversion and the yields of dry gas, ethene, and coke increased; those of LPG, propene, and butene showed maxima at 640 °C, and that of total light olefins showed a maximum at 660 °C. The yields of gasoline varied slightly below 640 °C and then decreased above 660 °C, and that of diesel decreased. The cracking extent increased with Table 1. HGO Properties item
value
item
value
density (20 °C), g/cm
0.9294
carbon residue, wt %
0.14
C, wt %
87.65
mean molecular weight,
308
H, wt%
11.90
186
S, μg/g N, μg/g
2700 1800
3
elemental analysis
increasing reaction temperature. Dry gas and coke were the end products of catalytic pyrolysis, so their yields increased with the increase in reaction temperature. However, LPG, gasoline, and diesel were intermediate products, so their yields decreased after reaching the highest yields because of secondary reactions.35,36 The optimal reaction temperature was 660 °C for the production of light olefins. 3.2. Effect of Weight Hourly Space Velocity. The effect of weight hourly space velocity (4.817.5 h1) on conversion and product yields was investigated, keeping reaction temperature, steam-to-oil weight ratio, and catalyst-to-oil weight ratio constant at 660 °C, 0.6, and 17, respectively. Table 4 lists the feed conversion and the yields of products. As weight hourly space velocity increased, feed conversion and the yields of dry gas, ethene, gasoline, and coke decreased; those of LPG, propene, butene, and diesel increased, and that of total light olefins increased at a weight hourly space velocity below 9.5 h1 and then varied slightly. High weight hourly space velocity led to short reaction time and low reaction extent. Therefore, feed conversion and the yields of end products such as dry gas and coke decreased as the weight hourly space velocity increased. The yields of LPG and diesel increased due to low secondary reaction extent. The optimal weight hourly space velocity was 9.5 h1. 3.3. Effect of Steam-to-Oil Weight Ratio. The effect of steam-to-oil weight ratio (0.21.8) on conversion and product yields was investigated, keeping reaction temperature, weight hourly space velocity, and catalyst-to-oil weight ratio constant at 660 °C, 9.5 h1, and 17, respectively. Table 5 lists the feed conversion and the yields of products.
g/mol distillation range initial boiling point
Table 2. Catalyst Properties item
(IBP), °C final boiling point, °C
583
H/C atomic ratio
IBP200 °C, wt %
0.8
structural group analysis, wt %
200350 °C, wt %
25.6
aromatic carbon
350500 °C, wt %
65.1
>500 °C, wt %
8.5
value
item
value
microactivity index
72
particle size distribution, wt %
surface area, m2/g
130
020 μm
4.8
24.7
pore volume, cm3/g
0.21
2040 μm
17.2
naphthenic carbon
30.8
packing density, g/cm3
0.88
4080 μm
47.8
paraffinic carbon
44.5
>80 μm
30.2
1.63
Figure 1. Diagram of the experimental setup (1, constant temperature box; 2, steam furnace; 3, feedstock; 4, electronic balance; 5, oil pump; 6, water tank; 7, water pump; 8, preheater; 9, reactor furnace; 10, thermocouple; 11, reactor; 12, inlet and outlet of catalysts; 13, filter; 14, condenser; 15, collecting bottle for liquid products; 16, gas collection vessel; 17, beaker; 18, gas sample bag). 3401
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Table 3. Feed Conversion and Product Yield as a Function of Reaction Temperature
Table 5. Feed Conversion and Product Yield as a Function of Steam-to-Oil Weight Ratio
for given reaction temperature
for given steam-to-oil weight ratio
600 °C 620 °C 640 °C 660 °C 680 °C 700 °C conversion, wt %
61.5
67.2
72.0
76.9
78.6
81.1
conversion, wt %
product yield, wt %
0.2
0.3
0.6
0.8
1.3
1.8
76.4
75.9
76.9
74.3
74.8
75.3
product yield, wt %
dry gas LPG
14.8 20.3
14.9 24.1
18.2 26.2
23.4 24.2
27.1 21.2
33.2 17.7
dry gas LPG
22.3 22.7
22.3 23.8
23.4 24.2
22.5 24.6
22.6 24.9
22.9 25.5
gasoline
19.3
19.4
19.0
18.4
17.4
15.9
gasoline
19.3
18.4
18.4
18.1
17.5
17.8
diesel
24.6
21.2
19.8
15.9
15.3
13.1
diesel
16.7
17.0
15.9
17.8
17.3
17.2
heavy oil
13.9
12.6
8.2
7.2
6.1
5.8
heavy oil
6.9
7.1
7.2
7.9
7.9
7.5
coke
7.1
7.8
8.6
10.9
12.9
14.3
coke
12.1
11.4
10.9
9.1
9.8
9.1
ethene
7.8
7.6
9.3
11.2
11.7
14.2
ethene
9.6
10.0
11.2
10.9
11.3
11.7
propene
13.4
14.3
16.1
15.6
13.9
12.7
propene
13.9
14.9
15.6
15.9
16.2
16.7
4.7 25.9
7.7 29.6
8.2 33.6
7.0 33.8
5.8 31.4
3.9 30.9
6.6 30.1
7.1 32.0
7.0 33.8
7.1 33.9
7.3 34.8
7.6 36.0
butene total light olefins
Table 4. Feed Conversion and Product Yield as a Function of Weight Hourly Space Velocity for given weight hourly space velocity 4.8 h1 7.2 h1 9.5 h1 11.6 h1 14.2 h1 17.5 h1 conversion, wt %
83.8
77.4
76.9
74.4
73.9
73.7
24.7 21.8
23.7 22.3
23.4 24.2
21.6 25.0
21.0 25.6
21.2 26.5
product yield, wt % dry gas LPG gasoline
25.2
20.3
18.4
17.5
17.8
16.7
diesel
11.3
15.7
15.9
17.1
18.1
18.1
heavy oil
4.9
6.9
7.2
8.5
8.0
8.2
coke
12.1
11.1
10.9
10.3
9.5
9.3
ethene
11.6
11.5
11.2
10.3
10.3
10.1
propene
14.3
14.8
15.6
15.7
16.1
16.5
5.8 31.7
6.2 32.4
7.0 33.8
7.5 33.5
7.7 34.1
7.9 34.5
butene total light olefins
As steam-to-oil weight ratio increased, feed conversion and the yields of dry gas and diesel varied slightly; those of LPG, ethene, propene, butene, and total light olefins increased, and those of gasoline and coke decreased. Increasing steam-to-oil weight ratio could reduce the partial pressure of oil gas, which favored the cracking of hydrocarbons to low molecular weight components. Furthermore, steam could restrain coking reactions on the catalyst surface and favor the retention of catalyst activity. Large steam-to-oil weight ratio was beneficial to the formation of light olefins, but it resulted in low reactor efficiency and high energy consumption. The optimal steam-to-oil weight ratio was 0.6 for high yields of light olefins and industrial operability in the future. 3.4. Effect of Catalyst-to-Oil Weight Ratio. The effect of catalyst-to-oil weight ratio (825) on conversion and product yields was investigated, keeping reaction temperature, weight hourly space velocity, and steam-to-oil weight ratio constant at 660 °C, 9.5 h1, and 0.6, respectively. Table 6 lists the feed conversion and the yields of products. As catalyst-to-oil weight ratio increased, feed conversion and the yields of dry gas, LPG, coke, ethene, propene, and
butene total light olefins
light olefins increased. However, gasoline yield and diesel yield showed a decreasing trend. Butene yield varied slightly at catalyst-to-oil weight ratio below 17 and then increased. High catalyst-to-oil weight ratio led to high average activity of catalyst and extent of cracking, which were favorable for the formation of light gas components and coke. However, a high ratio would also increase catalyst attrition, energy consumption, and operating difficulty. The optimal catalyst-to-oil weight ratio was 20 for high yields of light olefins and industrial operability in future. For catalytic pyrolysis of HGO, the optimal reaction temperature, weight hourly space velocity, steam-to-oil weight ratio, and catalyst-to-oil weight ratio were 660 °C, 9.5 h1, 0.6, and 20, respectively.
4. DEVELOPMENT OF THE SEVEN-LUMP KINETIC MODEL 4.1. Establishment of the Physical Model. The first step of the lumping kinetic study was to determine the lumps. The feed, HGO, could be considered as one lump (lump 1). Gasoline was a byproduct of catalytic pyrolysis and could be considered as another lump (lump 2). The target products of catalytic pyrolysis were light olefins, so gas components could first be split into two lumps, namely, light olefins and light alkanes. The formation mechanism of light gas components are different from that of heavy gas components;37 thus, gas components were further divided into four lumps, namely, propene plus butene (lump 3), propane plus butane (lump 4), ethene (lump 5), and hydrogen plus methane plus ethane (lump 6). Coke is a significant byproduct since coke yield is an important parameter for the design of the regenerator, and its formation mechanism is different from those of the other products. Therefore, coke could be considered as another lump (lump 7). The hydrogen-transfer and dehydrogenation reactions between gaseous components could be neglected. No reaction was considered between lumps 3 and 4, and between lumps 5 and 6. A seven-lump kinetic model with 15 reactions was established on the basis of the above analysis. The reaction network is shown in Figure 2. An advantage of this model is that it could predict the yields of the desired products (ethene and propene plus butene) directly. However, one limitation of this model is that the kinetic 3402
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Table 6. Feed Conversion and Product Yield as a Function of Catalyst-to-Oil Weight Ratio for given catalyst-to-oil weight ratio
conversion, wt %
8
12
17
20
25
73.1
75.3
76.9
78.5
79.3
22.4 23.9
23.3 24.5
23.4 24.2
24.5 26.1
25.4 26.4
due to catalyst deactivation. 0
ri ¼ ki ðFCi Þ
product yield, wt % dry gas LPG gasoline
18.8
17.8
18.4
16.5
15.9
diesel
18.1
16.9
15.9
15.0
14.1
heavy oil
8.8
7.8
7.2
6.5
6.6
coke
8.0
9.7
10.9
11.4
11.6
ethene
10.3
10.7
11.2
12.2
12.6
propene
15.1
15.4
15.6
16.8
17.1
7.1 32.4
7.2 33.3
7.0 33.8
7.4 36.4
7.6 37.4
butene total light olefins
Fb ε
ð3Þ
Equation 4 could be deduced from eqs 2 and 3. ∂FCi ∂Ci F 0 þ GV ¼ ki ðFCi Þ b ∂t x ∂x t ε
ð4Þ
The residence time of oil gas was much shorter than that of the catalyst for a confined fluidized bed reactor; this indicates that the rate of concentration change with respect to time is much lower than that with respect to position. Accordingly, eq 4 may be simplified to eq 5, which can also be expressed as eq 6 with the definition of the nondimensional length of the x cross-section in the bed. dCi F 0 ¼ ki ðFCi Þ b dx ε
ð5Þ
GV dCi F 0 ¼ ki ðFCi Þ b L dX ε
ð6Þ
GV
Equation 7 could be obtained by the definition of the mass velocity in the cross-section of the reactor (GV) and the weight hourly space velocity (SWH). GV ¼
SWH Fb L ε
ð7Þ
Equation 8 could be deduced from eqs 6 and 7. dCi 1 0 ¼ ki ðFCi Þ SWH dX Figure 2. Reaction network of the seven-lump kinetic model.
parameters are dependent on the feed and catalyst properties. Thus, the model could be improved by dividing the feedstock into several lumps to predict the cracking performance of different feeds. 4.2. Development of the Mathematical Model. In the reaction network of the seven-lump model, the reaction of lump I forming lump J could be expressed in the following mathematical model: ki
I sf vij J
ð1Þ
Several assumptions were made to simplify the mathematical model. The confined fluidized bed reactor could be considered as an isothermal plug-flow reactor, and the radial dispersion in the reactor was negligible. In the isothermal, gaseous-phase, and plug-flow reactor, the reaction system could be described by the continuity equation
∂FCi ∂t
x
∂Ci þ GV ¼ ri ∂x
ð8Þ
The actual rate constant ki0 equals the product of the intrinsic rate constant ki and the catalyst deactivation function (ϕ), which is shown in eq 9, where ki is a constant. 0
ki ¼ ki ϕ
ð9Þ
The density of oil gas may be expressed as eq 10, with the assumption that the oil gas follows the state equation of ideal gases. F¼
PM ̅ RT
ð10Þ
Equation 11 could be deduced from eqs 810 and was the model equation of the first-order reaction. dCi 1 PMϕ ̅ ki Ci ¼ SWH RT dX
ð11Þ
The mean molecular weight (M) is not a constant; it varies with X. Equation 12 shows the calculation method, where, i = 0 describes the influence of steam on mean molecular weight. n
ð2Þ
Reaction rate ri is proportional to the molar concentration of lump I (FCi) and the ratio of the catalyst mass density to the gas volume (Fb/ε), as shown in eq 3. Rate constant ki0 is not a constant and it decreases with the residence time of the catalyst
M ̅ ¼
∑ C i Mi i¼0 n
∑ Ci i¼0
¼
1 n
∑ Ci i¼0
ð12Þ
According to the reaction network of the seven-lump model and eq 11, the mathematical equations of the seven lumps may be 3403
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written as follows:
Table 7. Calculated Kinetic Constants
dC1 1 PM ̅ ðk12 þ k13 þ k14 þ k15 þ k16 ¼ SWH RT dX þ k17 ÞC1 ϕ
for given reaction temperature reaction constant
ð13Þ
3 1
k12, [g/cm ]
3h k13, [g/cm3]1 3 h1
dC2 1 PM ̅ ½v12 k12 C1 ðk23 þ k24 þ k25 þ k26 ¼ SWH RT dX þ k27 ÞC2 ϕ
k14, [g/cm3]1 3 h1 k15, [g/cm3]1 3 h1 k16, [g/cm3]1 3 h1
ð14Þ
dC3 1 PM ̅ ¼ ½ðv13 k13 C1 þ v23 k23 C2 Þ ðk35 þ k36 ÞC3 ϕ dX SWH RT
ð15Þ
dC4 1 PM ̅ ¼ ½ðv14 k14 C1 þ v24 k24 C2 Þ ðk45 þ k46 ÞC4 ϕ dX SWH RT
ð16Þ
k17, [g/cm3]1 3 h1 k23, [g/cm3]1 3 h1 k24, [g/cm3]1 3 h1 k25, [g/cm3]1 3 h1 k26, [g/cm3]1 3 h1 k27, [g/cm3]1 3 h1 k35, [g/cm3]1 3 h1 k36, [g/cm3]1 3 h1
dC5 1 PM ̅ ½v15 k15 C1 þ v25 k25 C2 þ v35 k35 C3 ¼ SWH RT dX þ v45 k45 C4 ϕ
k45, [g/cm3]1 3 h1
C7 ¼
ð18Þ
1 C 1 M 1 C2 M 2 1 þ RSO
C3 M3 C4 M4 C5 M5 C6 M6 =M7
ð19Þ
Coking, poisoning, and sintering were the three reasons for the catalyst deactivation, with coking as the most important one. The catalyst deactivation function (ϕ) could be described by ϕ ¼ expðRtÞ
ð20Þ
5. SOLUTIONS OF MODEL PARAMETERS Fifteen rate constants and the catalyst deactivation constant of the seven-lump kinetic model were estimated by solving the system of differential equations [eqs 1319]. The system of differential equations was solved by a program compiled in Matlab language using the RungeKutta method based on the Taylor theorem. Equation 21 shows the objective function. The LevenbergMarquardt algorithm was used to accelerate the convergence of the objective function. SðkÞ ¼
∑½yobsd ycalcd ðt, cðt, kÞÞT ½yobsd ycalcd ðt, cðt, kÞÞ
k46, [g/cm3]1 3 h1
ð17Þ
dC6 1 PM ̅ ½v16 k16 C1 þ v26 k26 C2 þ v36 k36 C3 ¼ SWH RT dX þ v46 k46 C4 ϕ
1
ð21Þ
Experiments were conducted at four temperatures, 620, 640, 660, and 680 °C, to obtain enough data for the fitting of kinetic constants. At each temperature, 24 groups of data were obtained with catalyst-to-oil weight ratio ranging from 7 to 26, weight hourly space velocity ranging from 5.5 to 15.5, and steam-to-oil weight ratio ranging from 0.3 to 1.0. The kinetic constants were calculated on the basis of the above program and experimental data. Table 7 lists the kinetic constants at 620, 640, 660, and 680 °C. The rate constants of the
R, h1
620 °C
640 °C
660 °C
680 °C
7352
8822
10859
13075
8804
10541
12594
13896
654
745
900
976
2651
3346
4804
6031
2951
3653
5810
8599
2851
3288
5184
6906
271
379
503
622
116 613
154 832
192 1214
217 1875
714
1015
1521
2661
1168
1622
2252
4483
540
790
1238
1984
326
535
942
1801
253
457
801
1388
213
436
813
1726
30.8
37.5
44.3
51.2
cracking of the feed lump were much larger than those of the secondary cracking of intermediate products, such as gasoline, propene plus butene, and propane plus butane. This suggests that the cracking of the feed played an important role in the catalytic pyrolysis of HGO, and the proportion of secondary cracking of intermediate products was low. The rate constants of conversion of gasoline to propene plus butene and to propane plus butane were low, and that of gasoline to coke was high. These indicate that the cracking performance of gasoline was poor and the condensation performance was good since the intermediate gasoline contained a large amount of aromatic rings with short side chains.36 Table 8 lists the preexponential factors and apparent activation energies calculated according to the Arrhenius eq 22. Many of the apparent activation energies were above 100 kJ 3 mol1. For reactions forming the same product lump, the apparent activation energy of the primary reactions was lower than that of the secondary reactions. The apparent energy of catalytic cracking is usually between 42 and 125 kJ 3 mol1, and that of thermal cracking is usually between 210 and 290 kJ 3 mol1.38 Many of the apparent energies of the catalytic pyrolysis of HGO were higher than that of catalytic cracking and lower than that of thermal cracking; these suggest that catalytic pyrolysis followed the carbonium ion and free radical mechanisms.24 E k ¼ k0 exp ð22Þ RT The apparent activation energies of formation of ethene and of hydrogen plus methane plus ethane were higher than those of formation of propene plus butene and of propane plus butane. These differences are due to the predominant carbonium ion mechanism in the formation of LPG and the predominant free radical mechanism in the formation of dry gas. The concentration of aromatic hydrocarbons in the gasoline product is above 80%,39 and the apparent activation energies of the gasoline lump to gas lumps were high. These indicate that the catalytic cracking performance of the gasoline lump was poor. Similarly, the 3404
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Table 8. Preexponential Factors and Apparent Activation Energies preexponential factor reaction
value
apparent activation energy, kJ/mol
6.93 107
68
1.22 107
54
3.75 105 1.25 109
47 97
7.45 1010
127
3h k23, [g/cm3]1 3 h1
3.78 109
105
1.43 108
98
2.35 106
73
k25, [g/cm3]1 3 h1
3.30 1010
133
9.07 1011
156
k27, [g/cm3]1 3 h1 k35, [g/cm3]1 3 h1
2.47 1012 5.30 1011
160 154
k36, [g/cm3]1 3 h1
2.13 1014
202
k45, [g/cm3]1 3 h1
1.41 1014
201
k, [g/cm3]1 3 h1
5.95 1016
247
R, h1
9.79 104
60
k12, [g/cm3]1 3 h1 3 1
k13, [g/cm ]
1
3h k14, [g/cm3]1 3 h1 k15, [g/cm3]1 3 h1 k16, [g/cm3]1 3 h1 3 1
k17, [g/cm ]
1
k24, [g/cm3]1 3 h1 k26, [g/cm3]1 3 h1
Figure 3. Comparison of the experimental yields (points) and the model-predicted yields (line) for catalytic pyrolysis of HGO at 660 °C.
Table 9. Average Relative Error of Conversion and Lump Yields at Four Temperatures, % for given reaction temperature 620 °C
640 °C
660 °C
680 °C
conversion
3.8
4.1
3.6
4.2
gasoline
6.1
7.4
6.7
6.3
propene + butene
4.6
4.5
4.2
4.7
propane + butane
9.8
10.5
12.5
12.8
ethene
5.5
5.3
6.0
6.6
hydrogen + methane + ethane
5.1
4.5
4.0
5.3
coke
5.6
4.9
4.6
5.2
Figure 4. Predicted and experimental yields of light olefins as a function of reaction temperature (catalyst-to-oil weight ratio of 17, weight hourly space velocity of 9.5 h1, and steam-to-oil weight ratio of 0.6).
apparent activation energies of propene plus butene and propane plus butane lumps were above 150 kJ 3 mol1; these indicate that the catalytic cracking performance was poor and it was mainly thermal cracking reactions that took place at high temperature.
6. MODEL TEST AND PREDICTION 6.1. Calculation Error. The calculated average relative errors of the conversion and the yields of product lumps are listed in Table 9. The calculated average relative error of conversion, gasoline, propene plus butene, propane plus butane, ethene, hydrogen plus methane plus ethane, and coke were 3.9, 6.6, 4.5, 11.4, 5.9, 4.7, and 5.1%, respectively. The calculated average relative error was around 6.0%, which indicated that the calculated conversion and lump yields were close to the experimental values. The comparison of the experimental yields (points) and the model-predicted yields (line) at 660 °C is illustrated in Figure 3. The predicted yields were close to the experimental values. All of these show that the seven-lump kinetic model had good calculation accuracy and that the calculated results were reliable. 6.2. Variation of Product Yields with Reaction Temperature. Rate constants at different temperatures were calculated on
Figure 5. Predicted and experimental yields of product lumps as a function of reaction temperature (catalyst-to-oil weight ratio of 17, weight hourly space velocity of 9.5 h1, and steam-to-oil weight ratio of 0.6).
the basis of the Arrhenius equation. The product yield as a function of reaction temperature within 600700 °C was predicted for the catalytic pyrolysis of HGO, keeping the catalyst-tooil weight ratio, steam-to-oil weight ratio, and weight hourly space velocity constant at 17, 0.6, and 9.5 h1, respectively. 3405
dx.doi.org/10.1021/ef200545m |Energy Fuels 2011, 25, 3400–3407
Energy & Fuels The model-predicted product yields and the experimental values are illustrated in Figures 4 and 5. The predicted yields were close to the experimental ones. As the reaction temperature increased, the yields of ethene, hydrogen plus methane plus ethane, and coke increased monotonically, and those of propene plus butene and gasoline showed maxima at about 640 °C; that of total light olefins showed a maximum at about 660 °C, and that of propane plus butane varied slightly. The predicted variation of product yields with reaction temperature was similar to the experimental variation.37
7. CONCLUSIONS (1) The optimal reaction temperature, weight hourly space velocity, steam-to-oil weight ratio, and catalyst-to-oil weight ratio, for catalytic pyrolysis of HGO, were 660 °C, 9.5 h1, 0.6, and 20, respectively. Under these reaction conditions, the yields of ethene, propene, and total light olefins were 12.2, 16.8, and 36.4 wt %, respectively. (2) The reactants and products of catalytic pyrolysis of HGO were classified into seven lumps, namely, HGO, gasoline, propene plus butene, propane plus butane, ethene, hydrogen plus methane plus ethane, and coke. A seven-lump kinetic model was established to describe the reactions. The kinetic model had 15 rate constants and a catalyst deactivation constant. (3) The kinetic parameters were calculated. The rate constants of the feed lump were larger than those of the intermediate product lumps, and the apparent activation energies of the feed lump were lower than those of the intermediate product lumps. (4) The model-predicted yields agree well with the experimental values. The predicted variation of product yields with the reaction temperature was similar to that of the experimental data. ’ AUTHOR INFORMATION Corresponding Author
*Tel.: 8610-8973-3993. Fax: 8610-6972-4721. E-mail: jsgao@ cup.edu.cn.
’ ACKNOWLEDGMENT Financial support was provided by the National Science Fund for Distinguished Young Scholars of China (Grant Nos. 20525621 and 20725620), and the Major Research Plan of the Ministry of Education of China (Grant 307008). ’ NOMENCLATURE R = catalyst deactivation constant F = gas density, g 3 cm3 Fb = density of catalyst in the bed, g 3 cm3 ε = void fraction of the bed ϕ = catalyst deactivation function vij = chemical measurement coefficient for the reaction of lump I to lump J c = concentration vector calcd = calculated data Ci = concentration of lump I in the gas phase, mol 3 g1 E = apparent activation energy, kJ 3 mol1 GV = mass velocity in the cross section of the reactor, g 3 cm2 3 h1 I, J = lump species I and J k0 = preexponential factors, g1 3 cm3 3 h1 ki = rate constant of lump I, g1 3 cm3 3 h1
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L = total length of the bed, cm M = average molecular weight of gaseous components, g 3 mol1 Mi = average molecular weight of lump I in the gas phase, g 3 mol1 obsd = experimental data P = reaction pressure, Pa ri = reaction rate of lump I, mol 3 cm3 3 h1 R = gas constant, J 3 mol1 3 K1 RSO = steam-to-oil weight ratio S = objective function SWH = total weight hourly space velocity, h1 t = running time or residence time of catalyst, h T = reaction temperature, K T0 = transpose of vector matrix x = distance from the reactor entrance to the x cross-section in the bed, cm X = nondimensional length of the x cross-section in the bed y = observed variable
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