Effect of Holdup Incomplete Catalyst Wetting and Backmixing during Hydroprocessing in Trickle Bed Reactors J. A. Paraskos and J. A. Frayer Gulf Research & Development Company, Pittsburgh. Pennsyivania 75230
Y . T. Shah* Department of Chemical Engineering. University of Pittsburgh, Pittsburgh. Pennsylvania 75267
Recent literature has shown that hydroprocessing kinetic data obtained in small scale pilot plant trickle bed reactors can be influenced by “liquid holdup” or “incomplete catalyst wetting” effects (caused by maldistribution of the reactant liquid in the catalyst bed) as well as backmixing effects. In this paper, the role of these effects in hydrotreating 53% reduced Kuwait crude and a number of gas oils in a pilot scale trickle bed reactor are evaluated. The effects of variations in the liquid hourly space velocity, average reactor temperature, and catalyst bed length upon the kinetics of sulfur, nitrogen, nickel, and vanadium (where applicable) removal are examined in isothermal pilot plant trickle bed reactors. Kinetic data for desulfurization, denitrogenation, and demetallization are shown to follow apparent first-order kinetics over a wide range of conversion when the holdup model proposed by Henry and Gilbert (1973) and “effective catalyst wetting” model of Mears (1974) are employed. On the other hand, when the holdup and the effective catalyst wetting effects are neglected, the apparent .order of reaction undergoes a change as the conversion level increases.
Introduction Evaluation of kinetic data obtained in single phase, continuous stirred tank reactors is a comparatively simple task. When dealing with a single phase, catalytic flow reactor, backmixing effects can be significant and kinetic data analysis should account for such effects when important. Catalytic reactions which involve gas and liquid phases are often carried out in trickle bed reactors, and kinetic data are usually obtained in laboratory scale pilot plant units prior to commercial reactor design. In the analysis of data derived in a trickle bed, one is faced with an assessment of the relative importance of the various mass transfer resistances (gas-liquid, liquid-solid, backmixing, etc.) and flow behavior (e.g., holdup, incomplete catalyst wetting) effects, as well as the interaction of these effects with the conversion levels observed in the reactor. The trickle bed reactor is extensively employed in the hydroprocessing of petroleum fractions for the removal of impurities such as sulfur and nitrogen, for hydrocracking to lower boiling range materials, and for altering the quality of the feed for subsequent processing. With worldwide demands increasing for “cleaner” fuels, especially those with reduced sulfur levels, new or revised trickle bed processes are being developed to deal with the increased conversion requirements. A t high conversion levels in a pilot plant trickle bed reactor, mass transport and flow effects are relatively more pronounced than a t lower conversions and can begin to have a substantial effect upon the design of a commercial system unless properly accounted for in the analysis of the pilot plant data. Analysis of constant pressure, isothermal kinetic hydrotreating data for various petroleum derived feedstocks obtained in standard small scale pilot plant trickle bed units indicates that the following three effects may be of importance; internal diffusion within the catalyst, reactor backmixing, liquid holdup or incomplete catalyst wetting (Shah and Paraskos, 1974; Mears, 1971; Henry and Gilbert, 1973; Mears, 1974). Internal diffusion within the catalyst is characterized by the effectiveness factor, usually obtained experimentally by measuring kinetic rates using
two or more catalyst particle sizes, under conditions where reactor holdup, incomplete catalyst wetting, and backmixing effects are negligible. Both backmixing and liquid holdup or incomplete catalyst wetting reduce the efficiency of a trickle bed reactor. Backmixing is solely a mass transfer phenomenon, whereas liquid holdup or incomplete catalyst wetting is caused by maldistribution and bypassing of liquid over the porous catalyst pellets. The maldistribution of the liquid within the catalyst bed causes an ineffective use of active catalyst sites. Although different in their basic nature, both backmixing (effectively characterized by the Peclet number) and liquid holdup or incomplete catalyst wetting have been correlated with the same system variables (Sater and Levenspiel, 1966; Hochman and Effron, 1969; Puranik and Vogelpohl, 1974; Onda et al., 1967). Therefore, it is difficult to isolate the effect of backmixing from that of liquid holdup or incomplete catalyst wetting in the analysis of kinetic data derived from a trickle bed reactor. Recently, Mears (1971) presented criteria for determining the importance of backmixing effects in a trickle bed reactor. He also outlined a backmixing model which was used to correlate experimental data on denitrogenation of West Coast straight-run gas oil in trickle bed reactors of varying catalyst bed lengths. He suggested that backmixing was the most likely cause for the poor performance of the shorter catalyst beds. Henry and Gilbert (1973), on the other hand, presented a model which was used to evaluate the same kinetic data assuming that holdup effects were controlling with negligible backmixing. Parts of Mears’ (1971) data were satisfactorily correlated by the holdup model of Henry and Gilbert. However, Mears (1974) later on showed that his data (1971) with diluted bed cannot be evaluated with the holdup model of Henry and Gilbert (1973). In general, both backmixing and liquid holdup or incomplete catalyst wetting effects would be important in small scale pilot plant reactors. In this paper, we first briefly analyze the roles of these effects on the kinetics occurring in a trickle bed reactor. Subsequently, we analyze the data derived from isothermal, isobaric hydrotreating of 53% reduced Kuwait crude and various types of gas oils in pilot plant trickle Ind. Eng. Chem., Process Des. Dev., Vol. 14, No. 3, 1975
315
Table I. 53% Kuwait Reduced Crude Residual Hydrodesulfurization Experimental Results
-
Run no. Charge stock
-
1
2
Yield, wt % of residue API Sulfur, wt % Metals, ppm Nickel Vanadium Nitrogen, wt % Distillation, D 1 1 6 0 5% 10% 20% 3 0‘) 4 0% 50% 607 7 OLG 80%
4
5
6
7
8
-Cs N-CS i -C, N-C, 2
c, CZ C1
H2S
15.4 3.89
93.42 21.1 1.03
91.44 22.2 0.61
90.23 24.8 0.47
87.20 24.5 0.29
84.13 24.8 0.19
88.87 23.6 0.46
88.87 22.6 0.71
90.54 21.1 1.16
15 54 0.22
7.3 17 0.16
4.9 12 0.14
2.8 7.8 0.11
1.3 4.5 0.08
1.1 3.5 0.09
3.9 8.5 0.13
5.6 12.0 0.15
8.1 19.0 0.16
633 700 752 844 901 967 1024
Operating conditions Days on stream Average reactor temperature, ” F LHSV, vol/hr/vol Normalized hydrogen partial pressure Gas circulation rate. scf/bbl Approx fresh catalyst charge Density, g/cm3 Bed size, cm3 of catalyst length, cni Unit yields, wt (? Residual fuel Lt oil
e . .
...
640 672 740 802 850 910 978
... ...
1.4 740 0.89 0.99 4892
93.42 5.18 0.03 0.08 0.03 0.13 0.16 0.10 0.03 1.89
637 668 729 789 844 903 967 1029
...
3.4 74 1 0.51 0.99 4901
91.44 6.51 0.05 0.10 0.06 0.24 0.24 0.16 0.02 2.54
bed reactors on the basis of “holdup” and “inconiplete catalyst wetting” models. Theoretical Although the following analysis applied to a first-order reaction, extension to other orders of reaction is straightforward. As shown by previous workers (Sater and Levenspiel, 1966; Hochman and Effron, 1969; Mears, 1971; Henry and Gilbert, 1973; Mears, 1974) both Peclet number (sometimes referred to as Bodenstein number) and liquid holdup in a trickle bed reactor are functions of liquid hourly space velocity (LHSV), catalyst size, catalyst bed length ( L ) ,liquid mass velocity, and the fluid properties, such as viscosity, density, and surface tension. For a reactor packed with a certain size catalyst and operating a t constant temperature and pressure conditions, experimental data are usually obtained a t various liquid hourly space velocities. The Peclet number is related to LHSV, 1, and the feedstock viscosity by an empirical relation of the following type (Hochman and Effron, 1969; Sater and 316
3
_______
Ind. Eng. Chem., Process Des. Dev., Vol. 14, No. 3, 1975
610 660 716 774 824 875 922 989 1065
621 653 718 776 826 881 93 8 1022 1028
6.2 739
8.5 74 1
0.29 1.0
0.21 0.99 4798
4958
90.23 8.22 0.05 0.13 0.03 0.17 0.21 0.07 4
.
.
2.70
87.20 8.95 0.07 0.22 0.14 0.25 0.27 0.35 0.10 3.86
610 643 693 747 793 843 888 956 1022
615 642 707 766 819 868 92 5 990 1073
10.9 781
622 653 714 773 820 875 927 1099
...
12.3 780
0.41 0.99 4970
0.66 0.98 5214
88.87 9.34
...
... ...
0.02 0.28 0.46 0.42
0.01 0.05 0.22 0.20
.
I
.
0.88 0.97 5164
84.13 13.72 0.05
3.48
17.4 781
...
3.62
88.87 8.01 0.13 0.27 0.08 0.39 0.45 0.10 0.38 2.52
620 655 728 774 829 886 944 1031
...
20.2 780 1.62 0.98 5050
90.54 6.40 0.07 0.12 0.03 0.18 0.31 0.28 0.15 3.14
Levenspiel, 1966)
Ped = u(LHSV)O(L)”
( 1) Similarly, the liquid holdup (volume of liquid per unit volume of the reactor) is correlated to LHSV and L by an expression of the type (Satterfield et al., 1969; Otake and Okada, 1953; Hochman and Effron, 1969; Mohunta and Laddha. 1965)
H = b(LHSV)B(L)B ( 2) Equations 1 and 2 are derived for the undiluted catalyst beds. Finally, the effective catalyst wetting is also correlated as (Mears, 1974)
.aff
= c(L)’(LHSV)’ (3) The constants a, b, and c in eq 1 and 2 are dependent upon the catalyst dimensions and other fluid properties. Equations 1 and 2 were developed assuming isothermal and isobaric conditions. In a completely ideal situation where backmixing and holdup or incomplete catalyst wetting effects are negligi-
ble, the governing equation for the reactor performance may be expressed as In--C A i = - k CAO LHSV
(4)
where Cai and Cao are the concentrations of a reactant at the reactor inlet and the outlet, respectively, and k is the apparent kinetic constant which includes the catalyst effectiveness factor and the void fraction of undiluted catalyst (Mears, 1971). Equation 4 indicates that, in the absence of backmixing and holdup or incomplete catalyst wetting effects, a log-log plot of In ( C A ~ / C Avs. ~ ) 1/LHSV should be a straight line with a slope of unity. Furthermore, a t constant LHSV, the conversion should be independent of catalyst bed length. If the holdup effect is important but the backmixing is insignificant, then the holdup model of Henry and Gilbert (1973) combined with eq 2 and 4 gives In-
c.4, c: A0
-
k'( L)* ( LHSV) I d
where k' = k b . Using the holdup correlation of Satterfield et al. (1969), Henry and Gilbert (1973) employed a value for the coefficient p = l/3. However, different values of 3 have also been reported (Hochman and Effron, 1969; Otake and Okada, 1953; Mohunta and Laddha, 1965; Ross, 1965; Mears, 1974). Equation 4 implies that a plot of In (C.L,~/CA~)vs. l/LHSV on a log-log paper should be a straight line with a slope 1 - p . Since p is most probably greater than zero and less than 1, the slope of the above line should be somewhere between 0 and 1. If the conversion data are taken a t constant LHSV but varying cata!yst bed length L, a log-log plot of In ( C A l / C A o ) vs. L should also be a straight line with the slope of the line equal to /3. The same effects of the catalyst bed length and liquid hourly space velocity on the conversion in the absence of backmixing are also predicted on the basis of catalyst wetting model of Mears (1974). According to this model
Mears (1974) takes y = 0.32. In the absence of concrete experimental data it is hard to justify the ranges of validity of eq 5 and 6. Henry and Gilbert (1973) indicate that eq 5 should be valid within the range of Reynolds number 10 to 600. This range, however, must depend upon the precautions taken for the proper distribution of liquid within the bed. The similar range for applicability of eq 6 is even harder to define. Mears (1974) reports that incomplete catalyst wetting has been observed a t a Reynolds number of 5 5 . However, a t large Reynolds number (of the order of 100) incomplete catalyst wetting does not appear to be feasible. At very low Reynolds number (of the order of l), eq 6 may be more valid than eq 5. If backmixing is important, then, in the absence of holdup or catalyst wetting effects, a relation between Ca0 and C A i would be given by (Wehner and Wilhelm, 1959).
where q = dl + 4h/Pe(LHSV) and Pe = PedL/d,. Here d, is the diameter of the catalyst. For a large Pe or a small deviation from plug flow, eq 7 can be approximated by
-AS C = C'Ai
exp
[- LHSV k + ___
(LHSV)* k2
h] Ped
(8)
Combining (1)and (81, one gets
where ( k ' ) 2 = k 2 d , / a It is clear from the above relationship that a log-log vs. 1/LHSV obtained under a set of plot of In (C,,/C4,,) reaction conditions would not be a straight line. Similarly, vs. L a t constant LHSV and a log-log plot of In (C,,/C,,,) other reaction conditions would also be expected to show curvature. The effect of backmixing in the presence of liquid holdup and incomplete catalyst wetting has recently been discussed by Mears (1974). This result once again indicates that even when backmixing is present in the presence of holdup or incomplete catalyst wetting, log-log plots of In ( C A , / C A ~ vs. ) l/LHSV (at constant L ) and In ( C A ~ / ( ' ~ ~ ) vs. L (at constant LHSV), would not be straight lines for first -order reactions. Experimental Results The above analysis was verified with two sets of hydrotreating data. The first set of data was obtained for hydrotreating of 53% reduced Kuwait crude in a reactor of given length. Conversions of sulfur, vanadium, nickel, and nitrogen were obtained as functions of LHSV. The second set of data was obtained for hydrotreating of various gas oils. These data were obtained to examine the effects of catalyst bed length on the conversion of sulfur and nitrogen in the gas oils. These two sets of data are evaluated separately below.
A. Hydrotreating of 53% Reduced Kuwait Crude. I . Experimental The experimental data were obtained in an isothermal. automated bench scale trickle bed reactor. The reactor was 1 in. in diameter and 44l/4 in. long. The catalyst bed was preceded by a 5-in. deep pre-mix zone packed by quartz chips. A li4-in. 0.d. thermowell was inserted a t the center of the reactor to measure the axial temperature distribution in the reactor. In the present study the catalyst bed length to the catalyst diameter ratio was varied from approximately 111 to 1554. The reactor was preceded by a 5 ft long preheater where oil and hydrogen are brought in intimate contact. The reactor was charged with a partially spent catalyst extrudates. The feed to the reactor was 5370 reduced Kuwait crude having the inspections shown in Table I. Two sets of data. at average reactor temperatures of 740 and 780°F. were obtained. For each temperature, four different space velocity runs, ranging from 0.21 to 1.62 h r - l , were made resulting in desulfurization levels from about 65 to 9570. The inlet gas, containing about 9070 hydrogen, was passed over the catalyst at a rate corresponding to about 5000 scf/bbl and a normalized hydrogen partial pressure of 1.0. Off gases were scrubbed with DEA, mixed with fresh makeup hydrogen, and recycled to meet the fresh feed at the entrance to the preheater. The fresh feed-hydrogen mixture was heated to reaction temperature and processed over the partially spent catalyst under isothermal conditions. The partially spent catalyst was removed from a commercial residual oil hydrodesulfurization process. This catalyst had acquired substantial levels of coke, nickel, vanadium, and other deposits but was still sufficiently active and stable to provide lined out data for the kinetic analysis of residual oil hydrotreatment. Ind. Eng. Chem., Process Des. Dev.. Vol. 14, No. 3, 1975
317
2,o
,
9
f i r s t order reaction
second order reactlop
1
I
I
I I I I I
I
I
I
I
I
I I I I I
s l o p e = 0 747
1
-
0
0z 0z
F
f e r 0 = 780
0 5
41
0
I
o
a
1
I 16
12
1
l
20
24
0 2
1 0
0 5
01
Figure 1 . Catalyst activity decline during residue HDS kinetic
study.
Figure 4. First-order kinetic plots based on the holdup model for the nitrogen removal reaction. I
60
,
l o
1
1
1
1
I
I
1 I ! .
0 5
01
.-
I
1
I
l/LHSV
1
I
04
l
I
I
I I I I
1 1
1
1
01
1
1
1
l
1
I
I
t o
0 5
1 ! 1 l 1 1 ! 10 0
5 0
/ L H S V Ihrl
'
I
I
1
c
s l o o e = ~619
Figure 5. First-order kinetic plots based on the holdup model for the nickel removal reaction.
I
t r
5 0
l
1
the vanadium removal reaction. 1
l
10 0
(hrl
Figure 2. First-order kinetic plots based on the holdup model for
10 0
'
l ! l / l
! 50
1 0
l
c
i
n/
i
1
t
1
sloDe = 0 646
10 0
5 0
lhll
1/LSHV
DAYS ON STREAM
i
/
L
1
1 0
0 1
I 05
1
#
!
1
,
1 0 l/LHSV
!
1
1
l
l
5 0
,
l 10 0
lhrl
oil
>
I
1
e
10'2
t 5C
Figure 3. First-order kinetic plots based on the holdup model for
the sulfur removal reaction.
11. Results and Discussion Residual Oil Hydrotreatment. A summary of the experimental results is given in Table I. An analysis of the kinetics of residual oil hydrodesulfurization, demetallization, and denitrogenation using the method outlined above follows. In the analysis of the data, corrections in the product concentrations of sulfur, nickel, vanadium, and nitrogen were made to account for variations in product yields. As shown in Figure 1, the catalyst activity decline was assumed to be a linear function of time. The product concentrations were reevaluated so that they are based on constant product yield and the apparent rate constants were adjusted to unit catalyst activity prior to kinetic analysis. The relative catalyst activity decline observed for the desulfurization reaction was assumed to hold for the removal of metals and nitrogen also. For the desulfurization and the demetallization reactions, standard first- and second-order kinetic plots using a plug-flow model were first investigated. These plots appeared to indicate t h a t sulfur removal follows second-order kinetics a t low conversions and first-order kinetics at higher conversion a t both temperatures investigated. The 318
Ind. Eng. Chem., Process Des. Dev., Vol. 14, No. 3, 1975
01 01
I
I
I
1 I , , # l 35 10
I/LHSV
I
I
I 1 I Illl3 5c
103
PC
Figure 6. Second-order kinetic plots based on the holdup model for the sulfur and nitrogen removal reactions.
apparent reaction order changed a t a conversion level of approximately 80%. Similarly, the apparent reaction order vanadium) reactions for demetallization (nickel changed from second order (at low conversion) to the first order (at higher conversion) at about 75% metals removal. The experimental data were subsequently analyzed using the analysis presented earlier which includes the effects of holdup, incomplete catalyst wetting, and backmixing. Assuming first- and second-order kinetics, log-log plots of In ( C . A . ~ / C ~and ~) [ ( ~ / C A ~ )( ~ / C A , )vs. ] l/LHSV for sulfur, nitrogen, nickel, and vanadium removal were obtained. These plots are described in Figures 2 to 7 . The results indicate that first-order kinetics correlate both desulfurization and demetallization (both nickel and vanadium) reactions quite well over the entire range of conversions.
+
Table 11. Summary of Experimental Data for Furnace Oil Run no. Charge stock furnace oil
1
2
3
4
5 ~-
API Sulfur, wt R Nitrogen, wt ?$ Distillation, D86, "F Over point End point 1OR
34.3 1.10
0.0184 405 624 455 473 489 504 518 534 552 571 593 609
20% 3 0% 40% 50% 60% 70% 80% 90% 95LZ Operating conditions Average reactor temperature, "F LHSV. vol/hr/vol Normalized hydrogen partial pressure" Gas circulation rate, scf/bbl Catalyst charge density, g/cm3 Bed Size: cm3 of catalyst length, cm Fkferred to same reference value employed in Table I
36.7 0.11 0.014 392 62 8 450 474 488 502 516 532 550 572 592 607
36.8 0.11
...
387 624 446 467 482 497 512 528 547 567 592 609
36.9 0.08 0.011
374 627 453 4 74 489 503 518 53 1 54 7 567 591 610
36.7 0.08
36.7 0.07
0.010
0.010
399 627 449 472 486 500 515 532 550 568 593 61 1
407 630 452 472 486 501 516 532 550 569 593 609
660 4.95 0.24 1461
660 5.02 0.24 1540
64 5 3.02 0.24 1523
64 5 3.03 0.24 1464
644 2.99 0.24 1476
87 17.71
175 35.63
87 17.71
175 35.63
262 53.34
< - - - - - - - -. - - - - - - _ - _ _ - _ - - - -0.788---------------------------
Note that if the plug flow prevailed, the slope of the lines in Figures 2 through 5 would be equal to 1.0. The results shown in Figures 6 and 7 indicate that second-order kinetics do not correlate the experimental data for desulfurization and demetallization reactions over the entire range of conversion. The slopes of the In ( C , ~ , / C Avs. ~ ) 1/LHSV plots for the desulfurization and the demetallization reactions illustrated in Figures 2, 3, and 5 vary from 0.532 to 0.922. This range includes the value 2/3 used by Henry and Gilbert (1973) in their data analysis for the hydrodesulfurization of gas oils and 0.68 suggested by Mears (1974). A slight temperature dependence of the slope of the Figures 2, 3, and 5 and the nature of the reactions implies that in a reacting system, the power-law coefficient in the holdup or the effective catalyst wetting-LHSV relationship may be dependent upon the reaction conditions (e.g., temperature) as well as the nature of the reaction. The denitrogenation of gas oils is commonly believed to be a first-order reaction (Mears, 1971; Henry and Gilbert, 1973). Experimental data for denitrogenation were also first evaluated assuming plug flow. Once again, neither first- nor second-order kinetic plots rendered a straight line. Subsequently, first- and second-order kinetic plots based on the holdup or catalyst wetting models were investigated. These results are described in Figures 4 and 6. These results indicate that just like desulfurization and demetallization, denitrogenation of residual oil also follows first-order kinetics. The slopes of a log-log plot of In ( C N , / C N ~vs. ) 1/LHSV were found to be approximately 0.766 at 740°F and 0.747 at 780°F. The results of Figures 2 through 7 indicate that the kinetics of desulfurization, demetallization, and denitrogenation reactions are well correlated with first-order kinetics when either liquid holdup or catalyst wetting effects are taken into account. The straight line relation-
>
ships of Figures 2 through 5 also imply that the backmixing effect in these experiments was negligible (see eq 9). Since the maximum ReL in these experiments was of the order of one and the catalyst bed length was approximately 53 cm, the study of Mears (1974) indicates that the backmixing effect in these experiments should indeed be negligible.
B. Hydrotreating of Various Gas Oils. I. Experimental These experiments were also performed in an isothermal, bench scale trickle bed reactor. All reactor temperatures were controlled within *5"F of the average reactor temperature. The reactor was charged with standard 1Ae-in. catalyst extrudates which were presulfided for 12 hr. The compacted density of the catalyst bed was 0.788 g/cm3. The reported product sulfur values are arithmetic averages of the results from analyzing five samples with an X-ray absorption analyzer. For product from a given period, the maximum difference among five individual analyses was 0.04 wt 70. 11. Results F u r n a c e Oil. A 50-50 blend (by volume) of Kuwait and California furnace oils was desulfurized at a normalized hydrogen partial pressure of 0.24, 1500 scf/bbl gas rate, and liquid hourly space velocities of approximately 3 .0 and 5.0 employing various catalyst bed lengths. Average reactor temperatures were 645°F a t 3 LHSV and 660°F at 5 LHSV. Very high sulfur removal rates, 90% or greater, were obtained. The experimental data are summarized in Table 11. As shown in this table, the product inspections, particularly the weight per cent sulfur and nitrogen, are essentially independent of the catalyst bed length for both sets of conditions of LHSV and average reactor temperature. The 5 LHSV, 660°F conditions were not examined in a 262-cm3 catalyst bed because of equipment limitations. Ind. Eng. Chem., Process Des. Dev., Vol. 14, No. 3, 1975
319
Table 111. Summary of Experimental Data for Vacuum Gas Oil Charge stock vacuum gas oil API 23.4 2.95 Sulfur, wt %; Nitrogen, wt %; 0.088 Distillation, vacuum: cor 760°F end point 1060 10% 666 20% 726 30% 77 1 40% 80 1 50% 836 60% 868 70% 908 80% 951 90%; 1005 95% 1043 Operating conditions Average reactor temperature, "F LHSV, vol/vol/hr Normalized hydrogen partial pressure" Gas circulation rate, scf/bbl
Run no. 2
1
27.3 0.69 0.067
3
27.3 0.67 0.060 1050 64 1 698 740 779 811 843 879 918 967 1002
1043 629 689 736 774 816 852 884 928 993 1026
1055 64 5 700 743 784 820 851 887 928 985 1022
650 1.01 0.40
64 9 0.97 0.41
1932
2028
27.0 0.64 0.059
650 1.01 0.40 1969
4 27.3 0.73 0.07
1055 639 690 742 7 84 811 84 5 886 924 984 1022 680 1.98 0.40 2015
5 27.1 0.65 0.071 1034 639 694 737 778 809 849 885 924 973 1009 680 2.01 0.40 1990
6 26.9 0.64 0.069 1050 63 9 695 740 777 815 843 879 922 969 1004 680 2.03 0.40 1941
Referred to same reference value employed in Table I
03
01
r1 t I
003 01
=740 F
i
/
/
temp = 780'F.
/
/
I
I
I
1
i
I I I I I
05
1
1
1
I l l l l l
50
10
l/LHSV
111. Analysis and Discussion
a'
'A
less, respectively, than the ones produced by the smallest bed. Light Gas Oil. A Kuwait light gas oil was desulfurized a t a normalized hydrogen partial pressure of about 0.32, 1800 scf/bbl gas, an LHSV of 2, and 650°F average reactor temperature. The experimental data are summarized in Table IV. The largest bed size again produced the highest quality product. The differences in sulfur and nitrogen concentrations in the products from the largest and the smallest beds were found to be 0.07 and 0.007 wt 70,respectively.
100
Ihrl
Figure 7. Second-order kinetic plots based on the holdup model for the nickel and vanadium removal reactions.
Vacuum Gas Oil. A Kuwait vacuum gas oil was desulfurized a t a normalized hydrogen partial pressure of about 0.40, 2000 scf/bbl gas rate and LHSV of approximately 1.0 and 2.0, also a t varying catalyst bed lengths. Average reactor temperatures were approximately 650°F a t 1 LHSV and approximately 680°F a t 2 LHSV. Desulfurization levels of 7570 or greater were obtained. The experimental data are described in Table 111. These data indicate that, a t the conditions of 1 LHSV and 650"F, the 262-cm3 catalyst bed gave sulfur and nitrogen concentrations in the product 0.05 wt % and 0.008 wt % less, respectively, than the ones obtained in the 87-cm3 catalyst bed a t the same conditions. For a LHSV of 0.2 and a temperature of 680"F, the largest catalyst bed produced a product with sulfur and nitrogen concentrations 0.09 and 0.01 wt 70 320 Ind. Eng. Chem., Process Des. Dev., Vol. 14, No. 3, 1975
The data shown in Table I1 indicate that both sulfur and nitrogen removal from furnace oil was independent of catalyst bed length. This means that under the reaction conditions employed, plug flow conditions prevailed in the reactor. Reynolds number, R ~ L in , this case was of the order of 10. It appears that low viscosity, a relatively greater ratio of vapor to liquid, and large LHSV may be the causes of near plug flow conditions in the experiments with furnace oil. For Kuwait vacuum gas oil and light gas oil, the catalyst bed length was found to have an effect on conversion of both sulfur and nitrogen. The shorter bed gave poorer conversions. Assuming that the catalyst bed length effect was caused by either liquid holdup or catalyst wetting and that first-order kinetics apply to both sulfur and nitrogen removal reactions (Mears, 1971; Henry and Gilbert, 1973), log-log plots of In ( C M / C A ~vs. ) L for both sulfur and nitrogen were obtained. These plots are shown in Figures 8 and 9. All the plots were well represented by straight lines implying that backmixing effects were not significant in these experiments. There was some scatter in the data for the nitrogen removal reaction. However, the trend of the
Table IV. Summary of Experimental Data for Kuwait Light Gas Oil .-
Charge stock Kuwait light gas oil Sultur, wt% Nitrogen. wt % Distillation, vacuum cor 760°F End point
852 613 654 685 708 730 74 5 766 782 805 823
10% 2 0% 3 0%
40% 50% 60% 7 0% 80% 90% 95% Operating conditions Average reactor temperature, "F LHSV, vol/hr/vol Normalized hydrogen partial pressure" Gas circulation rate, scf/bbl Catalyst charge density, g/cm3 Bed size, cm3 of catalyst length. cm a Fteferred to same reference value employed in Table I. I
3 . 0 1 - 7
I
I
I
I
I
0
VACUUM GAS OIL. L H S V : 1 , T E M P i650"F
[?
K U W A I T L I G H T G A S O I L L H S V - 2 T E M P :650°F
i
___-
3
2
1
30.1 0.61
30.3 0.64 0.036
26.8 2.46 0.046
API
Run no.
30.3 0.57 0.029
...
847 605 647 674 695 713 733 754 772 800 820
849 605 64 2 674 699 714 73 6 759 780 805 821
84 5 615 654 679 701 716 73 5 756 779 807 826
650 1.96 0.32
650 2.01 0.32
65 1 1.97 0.32
I
V A C U U M G A S 0 I L . L H S V - 2 , T E M P . : 680'F
2 0
slope = O 111
-__
slope = 0 0876
i 10
0 20
30
40
50
60
70
80
VACUUMGASOIL L H S V . 1 TEMP:650
F
VACUUM G A S O I L L H S V : 2 TEMP :e80
F
90 100
L icml
Figure 8. Effects of catalyst bed length on the desulfurization of gas oils-first-order kinetic plot based on the holdup model,
1 0
I
1 10
20
30
, -~ 40
50
60
70
BO
90 100
L (em)
data did not indicate curvature in the plots. The slopes of the plots varied considerably (minimum of 0.0498 for sulfur removal of vacuum gas oil at LHSV = 1, temperature = 650"F, and normalized pressure = 0.50 to the maximum of approximately 0.55 for nitrogen removal of Kuwait light gas oil a t LHSV = 2, temperature = 680"F, and normalized pressure = 0.32). The magnitude of the holdup effects was thus found to be strongly dependent upon the reaction conditions (i.e., LHSV, temperature, hydrogen partial pressure, etc.) and the nature of the reaction. The high values of the slopes shown in Figure 9 for denitrogenation a t 650°F average reactor temperature versus that a t 680°F may be due to different controlling mechanism of reaction a t these two temperature levels. It has been postulated that the denitrification of ring compounds may be taking place through an initial saturation of the nitrogencontaining ring followed by breaking of carbon-nitrogen bonds. It may be that the initial saturation of the nitrogen-containing ring may be more affected by the mass transfer and the flow distribution of the liquid a t 650°F than a t 680°F.
Figure 9. Effects of catalyst bed length on the denitrogenation of gas oils-first-order kinetic plot based on the holdup model.
Conclusions The study outlined in this paper indicates that desulfurization, demetallization, and denitrogenation of 53% Kuwait residue oil and the desulfurization and denitrogenation of various gas oils can all be represented as first-order reactions. Their kinetics in standard pilot plant trickle bed reactors are best modeled by considering either the liquid holdup effect as suggested by Henry and Gilbert (1973) or incomplete catalyst wetting effect as suggested by Mears (1974). The slope of In (CA,/CA
