Ind. Eng. C h e m . Res. 1989,28, 1589-1596 wt % telluric acid/Si02 system, which was less effective than highly dispersed telluric acid/SiOz. Furthermore, Te02was also detected by XRD even in the lightly loaded catalyst when water feed was omitted. To confirm that telluric acid is the active catalyst for phenol formation, additional experiments were carried out using K2Te04/Si02and its acid form, H2Te04/SiOz.The results are shown in Table I. As expected, the K2Te04/ Si02 catalyst was almost inactive for benzene oxidation, whereas its acid form was active. It has been reported by Iwamoto et al. (1983) that group 5 and 6 metal oxides supported on silica gel, especially vanadium oxide, are active for the formation of phenol, with N20 as an oxidant, while only a trace amount of phenol can be formed when O2is used instead of N20. In fact, from our separate experiments with group 5 and 6 metal oxide catalysts supported on silica gel, selectivity for phenol production in the oxidation of benzene by O2 could not exceed 2% in most cases. Bi203/Si02and Se02/Si02were practically inactive. By contrast, the selectivity of telluric acid supported on silica gel exceeded 40%. At the present stage, the one-pass yield of phenol is less than 1% . Nevertheless, a much higher yield can be expected by further optimization of the reaction conditions, the catalyst design based on the telluric acid catalyst, and reactor design. For example, by use of a recirculating flow reactor, 7% yield of phenol at 16% conversion of benzene was attained over 5 wt 90 silica-supported telluric acid catalyst at 733 K (conditions; catalyst weight, 5 g; total flow rate, 56 mlsmin-'; partial pressure, N2, 0.2, 02,0.17, benzene, 0.03, H20,0.6 atm; flow rate of recirculating gas, 500 mL-min-9. Further details will be reported elsewhere.
1589
Registry No. Benzene, 71-43-2;telluric acid, 7803-68-1;water, 7732-18-5; phenol, 108-95-2.
Literature Cited Groves, J. T.; Nemo, T . E. Aliphatic Hydroxylation Catalyzed by Iron Porphyrin Complexes. J . Am. Chem. SOC.1983, 105,
6243-6248. Guengerich, F. P.; Macdonald, T . L. Chemical Mechanisms of Catalysis by Cytochromes P-450 A Unified View. Acc. Chem. Res.
1984,17,9-16. Hamilton, G. A.; Friedman, J. P. A Hydroxylation of Anisole by Hydrogen Peroxide Requiring Catalytic Amounts of Ferric Ion And Catechol. J. Am. Chem. SOC.1963,85,1008-1009. Ioffe, I. I.; Levin, Ya. S. Direct Oxidation of Benzene to Phenol. Sb. Statei, Nauchno-Issled. Inst. Org. Poluprod. Krasitelei 1961, No. 2,88-117; Chem. Abstr. 1962,57,1177. Iwamoto, M.; Hirata, J.; Matsukami, K.; Kagawa, S. Catalytic Oxidation by Oxide Radical Ions. 1. One-Step Hydroxylation of Benzene to Phenol over Group 5 and 6 Oxides Supported on Silica Gel. J . Phys. Chem. 1983,87,903-905. Kimura, E.;Machida, R. A Mono-oxygenase Model for Selective Aromatic Hydroxylation with Nickel(I1)-MacrocyclicPolyamines. J. Chem. Soc., Chem. Commun. 1984,499-500. Kinoshita, T.;Harada, J.; Ito, S.; Sasaki, K. Aerial Oxidation of Benzene in the Presence of Electrochemically Generated Cu+ Ions. Angew. Chem., Int. Ed. Engl. 1983,22,502. Orita, H.; Hayakawa, T.; Shimizu, M.; Takerhira, K. Catalytic Hydroxylation of Benzene by the Copper-Ascorbic Acid-02 System. J. Mol. Catal. 1987,42,99-103. Udenfriend, S.; Clark, C. T.;Axelrod, J.; Brodie, B. B. Ascorbic Acid in Aromatic Hydroxylation I. A Model System for Aromatic Hydroxylation. J. Biol. Chem. 1954,208, 731-739. Walling, C. Fenton's Reagent Revisited. Acc. Chem. Res. 1975,8,
125-131. Received for review December 29, 1988 Accepted July 5, 1989
Testing Catalysts for Production Performance and Runaway Limits Imre J. Berty, Jozsef M. Berty,* Paul T. Brinkerhoff, and Tibor Chovan Berty Reaction Engineers, Ltd., 100 Lincoln Street, Akron, Ohio 44308-1708
The results of laboratory catalyst tests, conducted in recycle reactors under fixed conditions of feed rate and composition and otherwise at average production conditions, permit the evaluation of catalyst performance for production reactors. These tests are performed in short steady-state runs a t stepwise increasing temperatures until a specified product concentration is reached. From these results, in addition to performance evaluation, the thermal stability criteria of the reaction can also be calculated. This information is needed to maximize production within the thermal runaway limit. Since the thermal runaway limit, estimated from the catalyst test, does not contain assumptions on kinetics, the experimentally evaluated runaway limit can be used as a benchmark to help discriminate between kinetic models that were developed from other data sets. The evaluation of the performance as well as the thermal runaway limit is shown on actual experimental measurements made for the production of ethylene oxide by oxidation of ethylene. Mathematical derivations are kept simple and explanations detailed so that the method could be used without much difficulty under practical conditions.
Catalyst Tests in Recycle Reactors Evaluation of catalysts for production units requires answers to the following questions of economic importance: (1)What is the product concentration in the reactor effluent? (2) What is the production rate per unit reactor volume filled with catalyst? (3) What is the selectivity of the process? The answer to the first question is important to estimate the product separation cost and the cost of recycling any unconverted readants. The second answer helps to clarify the investment cost in the synthesis loop. The third 0888-5885/89/2628-1589$01.50/0
question addresses the raw material cost and the byproduct disposal charges. All these are considered in relation to the cost of the catalyst for its economical lifetime. The three questions mentioned above are interrelated; therefore, it is practical to introduce some standardization and simplification. For quality control tests and also for experiments to optimize a related, particular group of catalysts, one useful simplification is to require that in the tests a standard product concentration should be reached with the feed rate and feed concentration held at some standard level. This can be done by changing the temperature to reach the desired level of product concentra0 1989 American Chemical Society
1590 Ind. Eng. Chem. Res., Vol. 28, No. 11, 1989
tion. It was decided that the product concentration in the recycle reactor should be about half that expected from a tubular production reactor. This type of test was suggested by Bhasin et al. (1980) for ethylene oxide catalysts. In such a test, questions 1 and 2 are implicitly answered. The product concentration is given, and this multiplied by the standard flow rate gives the standard production rate. The temperature, required to achieve the production rate, is a measure of catalyst activity. The selectivity achieved at this temperature is the second important measure of catalyst quality. Other testing methods were published previously by Berty (1979a,b).
Example: Ethylene Oxidation In ethylene oxidation, three reactions progress: (1) ethylene oxidation to ethylene oxide (EO), (2) ethylene oxidation directly to carbon dioxide (DO), and (3) oxidation of ethylene oxide to carbon dioxide: rl: C2H4 + 0.502 C2H40, -AHr, = 117 kJ/mol of C2H4 (1) r2: C2H4 + 3.002
-
2.8
-
I
0 0 0 0 0 0 000000
2.6
2.4 2.2
I
I
-
-
2.0
-
1.8
-
1.6
-
141.2
,
:I
I
c++
+++++
onon
+
+
++
+
++++++
++
I
+++
I
+++
os{
I
04-+-, 0
1 10
30
20
+
EO mol&
I]
1
Time (Houn) COZ meld4
T/1OO
0
(C)
Figure 1. EO (mol %), C02 (mol W),and temperature vs time.
-
r3: C2H40+ 2.502
2C02 + 2H20, -AHr, = 1334 kJ/mol of C2H4 (2)
2C02 + 2H20, -AHra= 1217 kJ/mol of C2H40 (3)
Of these three reactions, only two are stoichiometrically independent and hence observable. Since the third reaction (r3) has only minor significance, the observable rates can be approximated by the first two. Neglecting the third reaction in the heat balance equation does not cause any error since
REO = rl - r3, RDo = 2(r2 + r3)
+ (-mr,)rz + (-Ws)r3 (-AHrl) + (-AHrJ = - M r Z qgen = (-AHrl)REo + 0 * 5 ( - ~ r , ) R ~ o
Qgen
3.0
= (-mrl)rl
(4)
OB
:a'
/
7 ~
05
0 4 ] 250
I
I
,
254
,
~,
258
$
262
,
266
, I
,
, 270
,
7 274
278
Temperature (C) 0
EO
mole%
+
c 0 2 mole%
(5)
Figure 2. EO and COB(mol %) vs temperature.
(6)
times the product concentration and the feed flow times the feed concentration were the bases for the material balance calculation. All instruments and the chromatograph were interfaced online to an IBM XT personal computer with a 20-Mbyte hard disk. Every 5 s, the computer read all continuous instrument signals, calculated in real time the running averages of the last 10 readings of all values, and displayed these in engineering units. Since the unit was built for unattended operation, an emergency shut-down system (ESD), independent of the computer, turned off the power, depressurized the unit, and flushed it with nitrogen if high temperature, high pressure, high oxygen concentration, or low shaft speed (rpm) was sensed. The chromatographic analysis, by the HACH-Carle Model 400 GC, took 30 min to finish;therefore, every hour, once feed and product analyses were completed, the computer calculated the material balances (for total mass and for C, H, and 0 atomic masses), the rates of reactions, the selectivity, and other performance variables and displayed the results on the CRT within a minute. This was also done during unattended operation. The results were stored on the hard disk memory of the personal computer and printed at request for off-line evaluation and correlation. A silver catalyst, similar to that described in the patent of Bhasin et al. (1980), was used in the experiments. Experimental conditions are given in column 3 ((Nielsen and La Rochelle, 1976) in a 5-in. Berty reactor) of Table I. The results of such tests are shown in Table I1 and in Figure 1 in time sequence and in Figure 2 with the results cross plotted as concentration vs temperature.
(7)
Experiments The test recommended by Bhasin et al. (1980) needs a search for the temperature in a recycle reactor (RR) where the desired product concentration is reached. This search consists of a series of short steady-state (SS) measurements after an initial SS was reached. Experiments were made according to the above recommendations in an internal RR designed by Berty (1974), but these can be made in any well-designed RR, like in the new ROTOBERTY reactor (Berty and Berty, 1987), or in any other nearly gradientless catalytic reactor. The experimental unit was built at our company for an export order. The experiments reported here were part of the acceptance test. The unit was equipped with six mass flowmeter controllers to prepare the feed from 02, C2H4, C2H6,C2H,C1 in CH,, COz, and N2 A seventh mass flowmeter controlled the total flow to the reactor, and the excess of the mixture served as a continuous feed sample. The reactor wall temperature was controlled by a solid-state temperature controller. The computer was used to cascade the inside temperature control on the wall temperature controller as was suggested by Silva (1987). The pressure of the hot product gas was lowered to 5 psig, right after the reactor for sampling by the gas chromatograph. The product gas line was heat traced to keep products from condensing. Product gas flow was metered by a turbine flowmeter. The product gas flow
Ind. Eng. Chem. Res., Vol. 28, No. 11, 1989 1591 Table I. Comparison of Union Carbide’s Air Test with Shell’s Oxygen Test and the Recommended Oxygen Test for Recycle Reactors UCC” Shell Shell in labb test conditions reactor pressure, psig temp (ref), “C rPm composition: ethylene, mol % ethane, mol % oxygen, mol % COz, mol % EO, mol % water, mol 70 nitrogen, mol % EDC inhibitor, ppm physical properties cp, kJ/kg MW, av, kg/kmol P , kg/m3, T,, P, I.C,kg/(m.s) catalyst charge vol, cm3 Wt, g shape size, cm 6 vol/surface, cm void vol fraction flow x sect, cm2 bed depth, l / d p flow conditions GHSV,h-l make-up feed, cm3/min, STP superficial lin vel, u,cm/s, TJ’, recycle ratio mass flow, m, kg/s mass vel, kg/(m2.s) Reynolds no. (p) Schmidt no. Prandtl no. mass-transfer coeff, m/s heat-transfer coeff, W/(m2.s) increase of CzH,O concn between top and bottom of the bed, mol % concn diff between catalyst surface and gas, % per pass adiabatic temp rise in the catalyst bed, K temp diff between catalyst surface and gas, K Bhasin et al., 1980. *Nielsen and La Rochelle, 1976.
5-in. Berty
microtube
5-in. Berty
275 260 2000
200 260
200 260 3000
feed
reactor
8.0 0.5 6.0 6.5
6.7 0.5 4.5 7.2 1.0 0.7 79.4 7.5
feed
reactor
8.0
30.0 0.5 8.0
62.0 5-8
61.5 10.0
28.2 0.5 5.8 1.0 1.5 1.0 62.0 10.0
1.19 29.31 13.25 22 x 104
1.46 28.34 9.16 22 x 10”
1.46 28.34 9.16 22 x 104
80.0***c 68.0*
2.57** 3.5***
40.0*** 34.0*
rings
crushed
rings
0.79 X 0.79 0.696 0.484 15.56 6.5
0.03 0.03 (?) 0.43 0.508 423.3
0.79 X 0.79 0.696 0.484 15.56 3.25
7,975 10,632 82.59 74.16 0.017 11.10 3410 0.857 0.606 0.039 768.0 0.014 0.012 3.105 1.585
3,300 2,200 1.58
6,555 4,370 116.8 185.0 0.017 11.04 3390 1.219 0.713 0.044 857.0 0.008 0.018 1.550 1.760
79.0 7.5
* = estimated, **
Thermal Stability Criteria The requirement for steady state in an ideal “isothermal” CSTR or RR with high recycle is q t r = qgen (W/m3) (8) The isothermal expression was introduced by Carberry (1976) to indicate that the feed is introduced to the reactor already heated to the reactor temperature; therefore, all heat is removed by heat exchange. Otherwise, all CSTRs have to be internally isothermal to be “gradientless”. The “slope condition” for thermal stability requires that dq,/dT 2 dqgen/dT (W/(m3W) (9) The need for this was explained graphically by Van Heerden (1958). This requirement simply states that with increasing temperature the heat-transfer rate should increase more than the heat-generation rate increases. This, in turn, is needed to overcome small upsets without any control action. The steady-state requirement and the slope condition can be deduced intuitively from fiit principles. A rigorous derivation was given by Aris and Amundson (1958). Another derivation by Carra and Forni (1974) can be found in Carberry’s book (1976). The same criterion applies to catalyst particles as to CSTRs, as was shown by Aris (1969).
= calculated,
30.0
none 0.145 2.0
*** = given.
Dividing the steady-state requirement (eq 8) with the slope condition (eq 9), we get Qgen qtr < (10) dqtr/dT - dqgen/dT and this is the most general expression of the Slope Condition for thermal stability, which in essence is also the criterion to avoid sensitivity (Baloo and Berty, 1989). The meaning of these results, for thermal stability conditions, can be shown in a simple example where the heat-transfer coefficient does not depend on temperature. In this case, the left-hand side of the previous inequality can be expressed as dqtr qtr qtr = Ua(T - Ts), dq,/dT = Ua, - = dT T - Ts (11)
Therefore,
In general, to evaluate the right-hand side of this ine-
1592 Ind. Eng. Chem. Res., Vol. 28, No. 11, 1989 Table 11. Test Resultso,* time, h 1.67 2.37 3.05 3.70 4.37 5.78 6.45 7.15 7.95 8.58 9.27 9.92 *10.57 *11.90 *13.08 *13.72 24.62 25.40 26.20 26.92 27.47 28.53 29.57 30.08 30.63 31.18 *31.72 *32.48
EO, mol %
COz, mol %
0.858 1.023 1.057 1.104 1.207 1.269 1.284 1.318 1.364 1.389 1.416 1.448 1.480 1.482 1.508 1.500 1.720 1.745 1.730 1.664 1.656 1.639 1.587 1.572 1.603 1.565 1.532 1.514
0.440 0.567 0.600 0.637 0.721 0.758 0.767 0.792 0.833 0.853 0.870 0.900 0.925 0.921 0.942 0.932 1.153 1.170 1.159 1.086 1.081 1.064 1.019 0.987 1.012 0.983 0.952 0.946 eff
mean
0.7625
temp, "C 252.2 260.8 262.8 264.4 267.4 268.0 268.2 268.9 270.1 270.4 271.0 271.7 272.3 272.2 272.6 272.7 276.8 276.6 276.3 274.9 274.9 274.3 273.2 272.7 273.4 272.5 272.0 271.7
eff 0.796 0.783 0.779 0.776 0.770 0.770 0.770 0.769 0.766 0.765 0.765 0.763 0.762 0.763 0.762 0.763 0.749 0.749 0.749 0.754 0.754 0.755 0.757 0.761 0.760 0.761 0.763 0.762 temn.
EO rate, mol/h 0.1004 0.1197 0.1237 0.1292 0.1412 0.1485 0.1502 0.1542 0.1596 0.1625 0.1657 0.1694 0.1731 0.1734 0.1764 0.1755 0.2012 0.2041 0.2024 0.1947 0.1937 0.1917 0.1857 0.1839 0.1875 0.1831 0.1792 0.1771
kJ/h 46.06 58.24 61.26 64.84 72.78 76.52 77.42 79.82 83.69 85.60 87.26 90.01 92.39 92.12 94.14 93.24 113.49 115.14 114.15 107.49 106.98 105.43 101.22 98.56 100.93 98.12 95.23 94.51 eff
O C
272.25
COz rate, mol/h 0.0514 0.0663 0.0702 0.0746 0.0844 0.0887 0.0897 0.0926 0.0975 0.0998 0.1018 0.1052 0.1082 0.1077 0.1102 0.1090 0.1349 0.1368 0.1356 0.1270 0.1264 0.1244 0.1192 0.1155 0.1184 0.1150 0.1113 0.1106
std
heat generation W 12.79 16.18 17.02 18.01 20.22 21.26 21.51 22.17 23.25 23.78 24.24 25.00 25.66 25.59 26.15 25.90 31.52 31.98 31.71 29.86 29.72 29.29 28.12 27.38 28.04 27.26 26.45 26.25 tema. O
0.0005
kW/m3 319.84 404.47 425.43 450.28 505.43 531.39 537.67 554.32 581.20 594.42 605.97 625.04 641.61 639.72 653.75 647.49 788.12 799.57 792.70 746.49 742.90 732.16 702.91 684.44 700.94 681.39 661.30 656.35
C
0.37
aAt the same conditions as in the third column of Table I. Feed = 4510 cm3/min (STP);AHrl= 117 kJ/mol; AHr8= 1334 kJ/mol; cat vol = 40 cm3. bCatalyst performance evaluation (points marked with * are considered).
quality, one has to measure experimentally q = f (T ) at a fixed residence time and feed concentrations. Then the function has to be differentiated and the function divided by its derivative. This procedure can be shown in the example of a reaction with first-order kinetics: dqgen E qgen= (-AHr)kOe-E/RTC,-= + -qgen - dC dT R T C~ d T ~ and for this example only the heat of reaction and all other thermodynamic and transport properties are considered to be independent of temperature in the narrow temperature range of interest. In the proposed experimental method, the effect of temperature on all properties is included in the results anyway. For the heat-transfer equation, U,,a should be considered at the limiting cross section. In laboratory recycle reactors used for catalytic studies, the equipment usually loses more heat than the reaction generates. Therefore, the limiting cross section and the critical T - Tois at the outside surface of the catalyst. In tubular reactors, the heat-transfer coefficient is equal or better at the catalyst than at the tube surface. In addition, for a unit length of tube, the heattransfer area is several times larger at the catalyst than at the tube wall; hence, the critical T - Todevelops at the inner tube wall. This then fits the case of the quasi-homogeneous models well, used most frequently in designs. If the feed rate and concentrations to the CSTR are constant, then with a given volume of catalyst, Co = constant, 8 = V / F = constant (14) Hence, from the material balance of a CSTR, CO d C - -Co ( E/ R T 2, 8koe-E/R _ c = 1 + 8koe-E/RT (15) ' dT [1 + 8 k o e - E / R T ] 2
Therefore, the total or ordinary derivative of the heatgeneration rate, as was shown by Perlmutter (1972) for first-order reaction, is
A more detailed explanation is given by Berty et al. (1982). ~ Now~dividing ~ the heat-generation rate by its total temperature derivative, we get
This is valid in this form only for first-order reactions in CSTRs. The general criteria in the form -< q t r dqtr/dT
Qgen
- dqgen/dT
(18)
is valid even if we do not know the explicit, analytical functions or their derivatives. It is enough to be able to experimentally measure qgen= (-AHr)r as the function of temperature at constant feed rate and feed concentration in a CSTR or RR. Then reading qW and the slope for the derivative at several temperatures and dividing these will give the quotient at all temperatures. The value at the temperature that gave the desired concentration is a reasonable estimate for the permissible temperature difference. The smallest of all values will give the most conservative, safest value to use. Since the heat of reaction multiplies both the function and the derivative thereof, it can be simplified from the numerator and the denominator for a single reaction, i.e., Qgen -- (-AZO=- r (19) dqgen/dT (-AHr) d r / d T dr/dT
Ind. Eng. Chem. Res., Vol. 28, NO.11, 1989 1593 Returning to the first-order reaction example, with the case where the heat-transfer coefficient does not depend on the temperature, the previous inequality simplifies to
9w
This inequality simply means that the temperature difference used for transfer of heat has to be less than a critical value to maintain SS and to overcome small, transient upsets without control action. If this temperature difference is limited, then considering the heat-transfer equation rearranged for showing the temperature difference explicitly, 3w
qtr
T - TS = -
Ua '
qtr
=
qgen
=Q
(21)
implies that, if two of the three variables U , a, and q are given, the third is limited not to require a higher than the permissible maximum temperature difference. For example, with an existing reactor of given design and flow conditions, U and a are fixed and known; therefore, the maximum heat load, q, can be calculated. Therefore, the maximum temperature difference estimated from the SLOPE CONDITION is an important guide for estimating the maximum production of a reactor as limited by thermal runaway. Normally this limit is approached only partially to the extent where the thermal sensitivity, expressed as dq /dT, is within the safe limit set by the control system. R e DYNAMIC CONDITION (Gilles and Hofmann, 1961; Berty et al., 1982),that sets the limits for oscillations is usually satisfied in all but a few unusual catalytic reactions. This is because the usually large heat capacity of the catalyst dampens out most transient changes. Only in processes where the heat capacity of the catalyst is relatively low and the reaction is very exothermic and fast, like ammonia oxidation on platinum gauze, is it important to consider the dynamic condition. The THERMAL RUNAWAY CONDITION is used for cases where no thermal feedback is present and only a sensitivity problem exists. This condition is essentially the same as the slope condition for a CSTR, as can be seen from the results of Froment and Bischoff (1979). The sharp distinction between cases of stability and sensitivity exists only in the mathematical models applied to various integral reactors. In real reactors, the thermal feedback, described by the second derivative in the axial direction, is always there; it is just not too important until the gradients become large. Therefore, the primary concern for well-controlled reactors is the avoidance of the onset of sensitivity. The limit for this onset in integral reactors can be estimated from the proposed test in RRs or in CSTRs. Once the sensitivity limit is exceeded, instability can set in. Conditions at instability are beyond the subject of this paper. The critical value for the temperature difference can be calculated from those test results, which were described in the first section.
Calculation of Thermal Stability Criterion from Catalyst Test Results Data shown in Figure 2, where the search results are shown as EO and carbon dioxide (DO) concentration against temperature, can be considered as corresponding rate vs temperature plots. The rate of reaction can be calculated from the material balance in a CSTR or a RR as F'C, - FbCi,, V
- ri
1 /
o.-----1 /
250
ZM
,
,
261
250
,
,
I
266
T.mp.mt"n
270
271
271)
(C)
Figure 3. Heat generation vs temperature (regression with secondorder polynomial).
where F' is the volumetric flow rate and Ci is the mole concentration of the subscripted component, measured at the temperature and pressure of the reaction. Since for every mole of EO made a half mole volume disappears and xE0 is between 0.01 and 0.03, the material balance can be approximated by and in these test
where xE0 is the mole fraction of EO made for each mole of feed and REO and RDo represent the rates of formation of the two products. In other words, the previously described catalyst test gives the rates as functions of temperature at fixed feed rates and feed concentrations around the average operating conditions. For two reactions (as in this case), the simplification of the heat-generation rate to the reaction rates is not practical, and the heat-generation rate was expressed as This is plotted against the temperature on Figure 3. The calculated maximum temperature difference is associated with a certain temperature and concentration or conversion at the given test conditions. At lower conversion, the maximum permissible value will be lower. One should limit the actual value to somewhat smaller than that obtained above to include a safety factor. The minimum value for the maximum temperature difference can also be experimentally determined. At any fixed temperature and at zero conversion, when C = Co, the total derivative reduces to the partial derivative of the heat-generation function:
and at C = Co,
d+
I"
=-
a+
IC
(27)
1594 Ind. Eng. Chem. Res., Vol. 28, No. 11, 1989 900
o
, 250
,
, 254
,
, 258
,
, 262
1"rotYre
,
, 266
,
, 270
,
, 271
,
,
,
!
250
278
262
266
270
-
274
278
Figure 5. Heat generation vs temperature (regression with expontential form). 50
The partial derivative of the rate can be evaluated experimentally in a CSTR from a series of measurements where at fixed feed concentration the temperature is changed and after each temperature change the feed rate is adjusted so as to give the same discharge concentration or conversion. The resulting curve gives the rate as a function of temperature at fixed discharge concentration, and differentiating will give the partial derivative of the rate with regard to temperature. The partial derivative will be a unique function of temperature for any order of rate, and it is the maximum value for the total temperature derivative. This divided into the heat-generation rate at the same temperature will give the most conservative, lowest estimate for the permissible temperature difference. The maximum rate in an industrial tubular reactor is not the initial rate, since feed is usually not quite preheated to the shell temperature of the reactor. Hence, the maximum rate will be reached where some level of conversion was already achieved. At this point, the rate is less than the initial rate would have been at the shell temperature. While a safe and conservative maximum temperature can be evaluated easily, a good knowledge of the technology has to be paired with the right interpretation of the experimental results to get a satisfactory but not too restrictive safety factor for the maximum permissible temperature difference.
251
Temprotwe (C)
(C)
Figure 4. Generalized slope condition (regression with second-order polynomial).
254
,
I
At T = 272.3 "C,qgen= 656.1 kW/m3 and dqgen/dT= 31.3 kW/(m3-K). Consequently,
Evaluation of Experiments Catalyst activity and selectivity results are given Table 11. In the example shown, to get better estimates of the results and errors, the temperatures and selectivities were averaged for ethylene oxide (EO) concentrations between 1.45 and 1.55 mol % , since the desired value was 1.50 mol %, and these are marked by an asterisk (*) in Table 11. From these results and for this catalyst, the required temperature was 272.3 "C, with *0.37 "C standard deviation (SD), and the selectivity was 76.25% with f0.05% SD to get 1.5 vol % ethylene oxide in the product gas. This concentration corresponds with the given space velocity to an average production rate r,, = 5.15 mol/(kg of catalyst-h) or = 1.22 mol/(m3-s). Runaway limits were calculated from all the results in Table I11 except the first three time points which may not have reached SS fully. The results of regression are summarized at the bottom of Table 111. To these data, a second-order polynomial was fitted by the least-squares method, and this is shown on Figures 3 and 4. The quadratic function of q, then, was differentiated with re-
Figures 5 and 6 show an alternative evaluation made by fitting q with a zero-order rate form, i.e., where the rate depends on T alone. Then taking the derivative and the quotient of the function and its derivative, the maximum permissible temperature difference gives about the same value as before:
At T = 272.3 "C,qgen= 654.7 kW/m3 and dq,/dT W/(m3.K). Consequently,
= 31.49
In the last results, the coefficient 14310 should not be interpreted as the real E / R , since, neglecting the multiplying affect of Co/C in the correlation, a smaller E resulted that compensated for the error caused by the simplification.
Ind. Eng. Chem. Res., Vol. 28, No. 11, 1989 1595 Table 111. Evaluation of Experiments experiment time, h 1.67 2.37 3.05 3.70 4.37 5.78 6.45 7.15 7.95 8.58 9.27 9.92 10.57 11.90 13.08 13.72 24.62 25.40 26.20 26.92 27.47 28.53 29.57 30.08 30.63 31.18 31.72 32.48
temp, OC 252.2 260.8 262.8 264.4 267.4 268.0 268.2 268.9 270.1 270.4 271.0 271.7 272.3 272.2 272.6 272.7 276.8 276.6 276.3 274.9 274.9 274.3 273.2 272.7 273.4 272.5 272.0 271.7
heat gen, kW/m3 319.84 404.47 425.43 450.28 505.43 531.39 537.67 554.32 581.20 594.42 605.97 625.04 641.61 639.72 653.75 647.49 788.12 799.57 792.70 746.49 742.90 732.16 702.91 684.44 700.94 681.39 661.30 656.35 0
dqldt a
b C
regression 1
regression 2
heat gen, q/(dq/dT), kW/m3 dq/dT, kW/(K.m3) K 10.20 23.44 239.02 19.01 365.54 19.23 19.04 406.10 21.33 23.01 19.19 441.57 19.70 26.16 515.32 19.83 531.20 26.79 27.00 19.87 536.58 20.04 555.73 27.73 28.99 20.34 589.77 29.31 20.42 598.52 29.94 20.59 616.29 30.67 20.78 637.50 20.96 656.10 31.30 20.93 652.97 31.20 21.05 665.53 31.62 21.08 668.70 31.72 22.42 36.03 807.59 35.82 22.35 800.41 35.50 22.24 789.71 34.03 21.77 741.03 21.77 741.03 34.03 720.80 33.40 21.58 32.25 21.23 684.69 668.70 31.72 21.08 32.46 21.29 691.17 31.51 21.02 662.38 30.99 20.87 646.75 20.78 637.50 30.67 a + b T + cTZ b + 2cT 139787 -541.45 0.5250
heat gen, q/(dq/dT), kW/m3 dq/dT, kW/(K.m3) K 19.29 239.95 12.44 18.68 19.92 372.09 20.49 20.07 411.23 445.25 22.05 20.19 25.28 20.42 516.14 20.46 531.51 25.97 536.73 26.21 20.48 555.37 27.05 20.53 20.62 588.72 28.55 20.65 597.34 28.93 29.72 20.69 614.94 20.74 636.07 30.66 654.71 31.49 20.79 31.35 20.78 651.57 31.91 20.81 664.22 667.42 32.06 20.82 38.40 21.13 811.49 38.06 21.12 803.84 792.50 37.57 21.10 741.49 35.33 20.99 35.33 20.99 741.49 20.94 720.57 34.41 683.63 32.77 20.86 667.42 32.06 20.82 690.21 33.07 20.87 20.81 661.04 31.77 645.33 31.07 20.77 30.66 20.74 636.07 exp(a b/T) -b/T2 exp(a + b/T) 32.720 -14 310
The two results are essentially identical. From the two different correlations, the second, using the exponential form (after logarithmic linearization), is preferred. This is because the derivative of the quadratic form close to the zero value can result in numerical instability. Other models for fitting the results can also be explored. The experimental result should not be extrapolated much, especially toward the low-temperature limits because the differentiation method can introduce significant errors. Otherwise the experimental method itself is not limited to any particular range of parameters. The main interest is in the range where production reactors work. The method will be useful for homogeneous reactions also; it is not limited to heterogeneously catalytic reactions since a "quasi-homogeneous" model was used. The thermal runaway limit calculated from catalyst test measurements can be compared to the same developed by computer simulation of tubular reador performance. This can help to discriminate models since the runaway limit is a well-defined, sharp value. Predictions can also be compared with results of experimental runaway limits in plant-size single-tube pilot plants, and in this way, models can be judged at the "worst" or most sensitive conditions, i.e., right at or before runaway, as was proposed earlier (Berty, 1979a,b).
Conclusions In summary, the thermal stability or runaway criterion was evaluated from catalyst test results in addition to the usual determination of activity and selectivity and without additional measurements. This can be done with other catalysts too, if those are evaluated by the method suggested by Bhasin et al. (1980). In cases where highly skilled personnel or a feedback controller (Pan et al., 1988) brings
+
the unit to the desired performance quickly, a few additional measurements may be needed to introduce more variations in the temperature and corresponding performance to generate enough change in the data for evaluation. The heat-generation rate divided by its temperature derivative is the most general expression to estimate the runaway condition and thermal stability criteria. The heat generation rate can be measured experimentally and no knowledge of the kinetics in any form is needed, and this then assures that various approximations included in the kinetic model will not add any error to the estimation of the runaway conditions. In addition, runaway criterion calculated from the proposed catalyst test can be compared to that calculated from proposed kinetic models. This then can help in the task of model discrimination. Registry No. EO, 75-21-8; ethylene, 74-85-1.
Literature Cited Aris, R. On Stability Criteria of Chemical Reaction Engineering. Chem. Eng. Sci. 1969, 24, 169. Aris, R.; Amundson, N. R. An Analysis of Chemical Reactor Stability and Control-I. Chem. Eng. Sci. 1958, 7, 121. Baloo, S.; Berty, J. M. Simulation of Experiments for the Determination of the Thermal Stability Criteria for the "UCKRON-I" Test Problem. Chem. Eng. Commun. 1989, 76, 73-91. Berty, J. M. Reactor for Vapor-Phase Catalytic Studies. Chem. Eng. Prog. 1974, 70, 78. Berty, J. M. Testing Commercial Catalysts in Recycle Reactors. Catal. Reu.-Sci. Eng. 1979a, 20, 1, 75. Berty, J. M. The Changing Role of the Pilot Plant. Chem. Eng. Prog. 197913, Sept, 48-51. Berty, I. J.; Berty, J. M. New Concepts for Shaftless Recycle Reactors. AIChE Annual Meeting, New York City, 1987; Paper 32D. Berty, J. M.; Lenczyk, J. P.; Shah, S. M. Experimental Measurement of the Stability Criteria for the Low Pressure Methanol Synthesis. AZChE J. 1982, ,%(No. 6), 914.
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Ind. Eng. Chem. Res. 1989,28, 1596-1600
Bhasin, M. M.; Ellgen, P. C.; Hendrix, C. D. Catalyst Composition and Process for Oxidation of Ethylene to Ethylene Oxide. United Kingdom Patent GB 2,043,481 (to Union Carbide Corp.), 1980. Carberry, J. J. Chemical and Catalytic Reaction Engineering; McGraw-Hill: New York, 1976. Forni, L. Aspetti Cinetici della Teoria del Reatore-ChimCarra, S.; ico. Tamburirni, Milan, 1974. Froment, G. F.; Bischoff, K. B. In Chemical Reactor Analysis and Design; John Wiley and Sons: New York, 1979. Gilles, E. D.; Hofmann, H. Bemerkung zu der Arbeit: An Analysis of Chemical Reactor Stability and Control. Chem. Eng. Sci. 1961, 5, 328. Nielsen, R. P.; La Rochelle, J. H. (to Shell Oil Co.) Catalyst for Production of Ethylene Oxide. U.S. Patent 3,962,136, 1976.
Pan, D.-F.; Schnitzlein, K.; Hofmann, H. Design of the Control Scheme of a Concentration-Controlled Recycle Reactor. Ind. Eng. Chem. Res. 1988,27,86-93. Perlmutter, D. D. Stability of Chemical Reactors; Prentice-Hall: New York, 1972. Silva, J. M. Cascade Temperature Control System for the Berty Reactor. Ind. Eng. Chem. Res. 1987,26, 179-180. Van Heerden, C. The Character of the Stationary State of Exothermic Processes. Proc. of the 1st European Symposium on Chemical Reaction Engineering; Pergamon Press: Amsterdam, 1958; p 133.
Receiued for reuiew November 18, 1988 Accepted June 7, 1989
Naphtha Reforming Capacity of Catalysts with Different Metallic Functions Javier M. Grau and Jose M. Parera* Instituto de Inuestigaciones en Catdlisis y Petroqulmica, INCAPE, Santiago del Estero 2654, 3000 Santa Fe, Argentina
The aromatizing capacity and the deactivation produced by coke and sulfur of metals supported on A120,-C1 catalysts were studied for the reforming of n-paraffins (C7-Cl0) and naphtha cuts (light and heavy) under commercial conditions. The metallic components of the catalysts were Pt(0.3%), Pt(1.14%), and Pt(0.30% )-Re(0.30%)-S(0.04% ) for balanced metal and Pt(0.22 % )-Re(0.44 % )-S(0.064%) for skewed metal. Both Pt(0.3%) and Pt-Re balanced-metal catalysts are more selective toward aromatics with n-C7 and n-C8, while Pt(1.14%) and Pt-Re skewed-metal catalysts are better with n-C9 and n-Clo. Regarding the naphtha cuts, the balanced-metal catalyst produces a higher increase in octane number for the light cut than the skewed-metal catalyst, while the skewed is better for the heavy cut. T h e skewed catalyst produces more gas with both cuts. This catalyst is more stable and has a greater activity recovery after being deactivated by coke deposition. The thiophene deactivation is similar in both catalysts. Naphtha reforming is a catalytic process that increases the gasoline octane number mainly by an increment in the aromatic hydrocarbons concentration. A new catalyst composed of platinum supported on an acidic oxide was introduced after World War 11. The first catalyst of this generation was patented by Haensel (1949). When the processes is operated at high pressures (30-35 kg cm-2), these catalysts have a good selectivity toward aromatics and an acceptable stability with operation cycles of about 10 months. Nearly 20 years later, bimetallic catalysts appeared. Kluksdahl (1968) patented the addition of rhenium as a second metal to platinum. The most common concentrations of Pt and Re used in the process are of similar values (generally 0.3%), and the process (Rheniforming E) allows us to reduce the working pressure to 8-15 kg cm-2,with greater activity, selectivity, and stability than the monometallic catalyst (Gates et al., 1978). This equal-metal or balanced-metal catalyst was rapidly adopted in most of the reforming plants, and since it is more sensitive to poisoning by sulfur (Menon and Prasad, 1976),a more severe feed hydrodesulfuration is necessary. About 7 years ago, the same company, Chevron (US.), introduced a new catalyst with more Re than Pt (Rheniforming F), which is called a “skewed-metal” or skewed catalyst. According to Larsen (1986),this high Re catalyst presents greater activity and stability than the normal balanced-metal catalyst. These properties could allow for operation at a lower pressure and a lower hydrogen to hydrocarbon ratio, conditions that favor aromatization. Because of its high stability, Moorehead (1986) recom-
mends using the skewed catalyst in the last reactor. Knowing the behavior of catalysts with different compositions of the metallic component, mono- as well as bimetallic catalysts, for the reforming of different feeds could be useful in the selection of the catalyst as a function of the reformer charge. Changes in the selectivity as a function of feed composition using catalysts with the same metallic components (0.3% Pt, 0.3% Re) but different acid functions (alumina with different chlorine concentrations) were studied in a previous paper (Grau et al., 1988). In this paper, the activity, selectivity, and stability of four catalysts with the same acid function but different metallic functions are studied with n-paraffins of C7-Clo. The catalysts used are two monometallic catalysts with different Pt charges and two commercial bimetallic catalysts, Pt-Re, one with the same concentration of Re and Pt and the other with skewed concentrations, twice as much Re as Pt. Two hydrodesulfurized naphtha cuts doped with n-paraffins are also tested with the bimetallic catalysts.
Experimental Section Materials. 1. Catalysts. The monometallic catalysts were prepared by impregnation of r-Alz03 CK 300 supplied by Ketjen Cyanamid (Amsterdam) with an aqueous solution of HzPtC&-HCl,according to the technique of Castro et al. (1981). The bimetallic catalysts were of a commercial origin. The chlorine content of the catalyst samples was adjusted by passing a gaseous stream of air-HC1-water through the sample, following the method of Castro et al. (1983). To avoid the initial hyperactivity, the bimetallic
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