780
Ind. Eng. Chem. process
1. The liquid-phase axial mixing is very low, and the plug flow can be safely assumed for all cases studied. 2. The volumetric mass-transfer coefficient is of the same order of magnitude as in a conventional stirred vessel for corresponding bubble sizes. 3. The volumetric mass-transfer coefficient kLa is favorably affected by the three velocities studied: Vg,VL, and (AF). 4. The liquid-phase flow significantly enhances the mass transfer by simultaneously increasing the mass-transfer coefficient, kL,and the interfacial area, a. This may imply that a gas-liquid Karr column is best suited for continuous operations with high throughputs. 5. Both the energy input and the gas flow rate greatly increase the interfacial area but have an adverse effect on the mass-transfer coefficient, k L , presumably due to the decreased activity a t the gas-liquid interface as a result of the reduced slip velocity.
Acknowledgment We thank the Natural Sciences and Engineering Research Council of Canada for financial support. Nomenclature AF = reciprocating speed, cm/s a = interfacial area, cm
Des. mv. iga8, 25. 780-786
= liquid-side mass-transfer coefficient, cm/s kLa = volumetric mass-transfer coefficient, L/s L = liquid height in column, cm V , = effective liquid velocity, cm/s V, = superficial gas velocity, cm/s t = distance, cm kL
Greek Symbols c =
gas holdup
u = fractional open area of plate
Subscript e = exit stream value 0 = reference value Literature Cited Chen, B. H.; Vailabh, R. Ind. f n g . Chem. Process D e s . Dev. 1970, 9 , 121. Danchwerts, P. V. Chem. fng. Sci. 1953, 2 , 1. Karr, A. E. AIChE J . 1919, 5 . 446. Karr, A. E. Sep. Sci. Techno/. 1980. 15, 877. Karr, A. E. AIChE J . 1984, 30,697. Kim, S.D.; Baird, M. H. 1. Can. J . Chem. Eng. 1978, 5 4 , 81. Lo, T. C.; Prochazka. J. "Handbook of Solvent Extraction"; Lo, T. C., Baird, M. H. I . , Hanson, C., Eds.; Wiley-Interscience: New York, 1983. Noh, S.H.; Baird, M. H. I.AIChf J . 1984, 30,120. Miyanami, K.: To@, K.; Yano, T. J . Chem. f n g . Jpn. 1973, 6 , 518. Miyanami. K.; Tojo, K.; Yano, T. Chem. fng. Sci. 1978, 33,601. Tojo, K.; Miyanami, K.; Yano, T. J . Chem. f n g . Jpn. 1974a, 7 , 123; 1974b, 7, 127. Voyer, R. D.; Miller, A. I. Can. J . Chem. f n g . 1988, 46, 339. Wang, K. B.; Fan, L. T. Chem. Eng. Scl. 1978, 33,945. Yang, N. S.;Chen, 8. H.; McMilian. A. F.; Shen. 2. Q. Ind. Eng. Chem. Process D e s . Dev., in press.
c = concentration, mol/L c* = concentration at the interface, mol/L DL = axial dispersion coefficient, cm2/s
Received for review April 29, 1985 Revised manuscript received November 11, 1985 Accepted January 21, 1986
Pyrolysis of Hydrocarbons in a Large Pilot-Scale Reactor. 1. Experimental Design Lester S. Kershenbaum" and Patrick W. Leaneyt Depafiment of Chemical Engineerhg and Chemical Technolw, Imperlel College, London S W7, England
An experimental technique is presented which eliminates many of the uncertainties associated with the determination of the kinetics of fast, complex reactions (such as the pyrolysis of light hydrocarbons) by minimizing the effects in the experknental system. Constructionof a welunstrumented of heat transfer, mass transfer, and large-scale reactor enabled measurements to be made within the reactor by the use of probes which could travel continuously through the reactor space, both radially and axially, measuring point compositions and temperatures. A detailed design and analysis of the probes ensured small but quantifiable errors for the measurements and for the disturbance to the flow pattern. The experimental measurements yielded a complete set of data (conversion and product distribution vs. residence time) from a single run in an experimentally verified isothermal zone and for a well-characterized set of operating conditions.
The pyrolysis of light hydrocarbons has been studied extensively over the past 50 years with varying objectives. Early work sought to determine product distributions over a range of operating conditions (Frey and Smith, 1928; Schneider and Frolich, 1931; Schutt, 1947) or overall kinetic expressions which could predict the gross behavior of these complex reaction systems (Laidler et al., 1962; Martin et al., 1964; Kershenbaum and Martin, 1964). Despite a significant amount of experimental work, discrepancies between the various workers was widespread, t Current address: Chevron
Research Co., Richmond, CA. 0 7 96-4305f86f 1 125-078080 1.5010
both in terms of product distribution and apparent activation energies for the overall reactions. Recent work has attempted to correlate the conflicting experimental data through complex kinetic models which more closely reflect the chemical steps involved in these reactions (Powers and Corcoran, 1974; Allara and Edelson, 1975; Sundaram and Froment, 1978 Edelson and Allara, 1980, Purnell, 1980). In taking this approach, workers have either assumed a steady-state concentration for the radical species or, alternatively, solved the stiff set of differential equations for the concentrations of all species, a much more difficult computational chore. Some workers have chosen to use in their models, the rates of the various 0 1986 American Chemical Society
Ind. Eng. Chem. Process Des. Dev., Vol. 25, No. 3, 1986 781
free-radical reaction steps as determined from independent fundamental studies or theoretical predictions. Others have used these kinetic parameters, to a greater or lesser extent, as adjustable constants in their models in order to minimize deviations from experimental results. In any case, there is a fair amount of arbitrary choice in knowing which steps to include and which parameters to adjust; furthermore, it is clear that no unique model is likely to emerge although several models may fit experimental data extremely well. An intermediate approach, favored by Froment and co-workers (Buekens and Froment, 1968; Van Damme et al., 1975; Froment et al., 1977; Sundaram and Froment, 1977), has been to approximate the set of free-radical reactions by a simpler set of lumped molecular reactions, chosen to reflect the free-radical reaction steps. Kinetic parameters were then obtained by nonlinear parameter estimation using experimental data from integral reactors. Complications clearly arise because of the nonisothermal reactor data and the well-known correlation between preexponential factors and activation energies. In some cases, it was necessary to account for apparent inhibition effects at higher conversions by empirical linear or hyperbolic relations. The purpose of this paper is to illustrate that some of the conflicting results may have been caused by poorly defined experimental conditions which were inconsistent with the models used to interpret the data. Under these circumstances, it is inappropriate to attempt to overcome contradictions merely by complicating the kinetic model. Some of the apparent inconsistencies (such as variable apparent orders of reaction, frequency factors, and activation energies) may not require complex models or empirical equations for their explanation. Rather, they may be caused by a series of extraneous factors which can be eliminated by more carefully controlled experimental procedures. For example, in a tubular or annular flow reactor, the use of a reactor wall temperature rather than a measured gas temperature is especially difficult to justify. The errors involved can be significant for pyrolysis reactions with high activation energies and, more important, will not, in general, be constant. The errors will be functions of the flow rate and temperature because of the corresponding effects on the heat-transfer coefficient and will also be dependent upon the gas composition because of the heat flux associated with the endothermic reaction. The use of a nonisothenpal reactor further increases the difficulties of obtaining well-defined kinetic parameters. The uncertainties associated with the heating and quenching of the process stream are only partially redressed by utilizing nonisothermal profiles or concepts of equivalent volumes or temperatures. The problems are compounded by the method generally used to study pyrolysis reactions. Typically, in any one experiment, only the composition of the reactor effluent is measured. The effect of varying residence time, and hence conversion, is obtained by altering the feed rate and composition to the reactor and noting the change in the composition of the effluent. In an ideal, isothermal plug-flow reactor, such an approach is perfectly satisfactory. However, for situations involving complex reaction kinetics in a nonisothermal reactor with an uncertain flow profile and possible interactions with the reactor wall, this approach is fraught with sources of serious error. The unmeasured temperature profile-whether axial or radial-will certainly be nonuniform and, moreover, a function of the Reynolds number. Since each run has a
different flow rate and, hence, Reynolds number, there will be a variable heat-transfer effect which partially obscures the kinetic steps. Similarly, the effects of mass transfer with the wall and nonideal radial and axial mixing will also vary; these will be incorporated into the overall kinetics in a biased manner and may give the appearance of abnormal behavior. For example, the apparent inhibition of propane decomposition at high conversions (high residence times) at a given temperature may be caused, at least partially, by the fact that in a conventional flow reactor, the increased residence time is achieved only by reduction of the throughput rate. This, in turn, leads to an alteration in the reaction conditions, a fact which cannot always be ignored. Recently, the experimental systems used by research workers have been designed to avoid these problems. The industrial-scale reactor described by Van Damme et al. (1975) and used subsequently by others had the provision for sampling of the process stream at certain points within the reactor. Hautman et al. (1981) studied the decomposition of propane in a cylindrical quartz reactor by using a water-cooled probe to withdraw samples from the reactor center line, and that configuration has also been used by other workers. Similarly, the objectives of this work were to carry out experiments in such a way as to eliminate or at least measure and minimize the effects of heat transfer, mass transfer, partial mixing, and nonisothermal profiles, all of which can be important in pyrolysis reactions. We constructed a well-instrumented, large-scale reactor from which measurements could be made of gas compositions and temperatures within the reactor space, both radially and axially. The ability to establish and experimentally illustrate an isothermal region somewhere within the reactor would then yield a complete set of information from one run on the composition vs. residence time for a well-characterized set of operating conditions. On this basis, temperature and gas sampling probes were designed to ensure that measurements from the reactor space, particularly from along the center line, were representative of well-determined sampling positions.
Experimental System Reactor Configuration. The primary objective in the design was to overcome the sources of error inherent in kinetic studies of pyrolysis reactions, as described in the previous section. Emphasis was placed on providing detailed measurements within the reactor, which might then justify the use of specific models in the analysis of the resulting data. A schematic diagram of the experimental system is shown in Figure 1. At the heart of the system was an inconel tubular reactor mounted vertically in an electrically heated furnace and tested up to a maximum temperature of 1000 O C . The reactor tube was 60.3 mm o.d., 52.5 mm i.d., and 1.825 m in length, although only 1.11 m of its length was within the furnace. The reactor wall temperature within the furnace was measured by 18 equally spaced platinum/platinum-rhodium thermocouples in inconel sheaths, cemented into the tube wall in holes drilled to within 1 mm of the inner wall and spot-welded along the outer tube wall. The main purpose of these thermocouples was to provide information for the controllers which adjusted the power input to the various sections of the electric furnace in an attempt to establish an isothermal region within the reactor. The electric furnace had a maximum power input of 10 kW, distributed within five separately controlled zones. The power input to each zone was controlled by thyristors
782
Ind. Eng. Chem. Process Des. Dev., Vol. 25, No. 3, 1986
Exhaust
I
Firma!
t-P A'.
A
*, ,
-
control Vdve
I FID
To
safety trip sys)em Nitrogen Supply U
N2
Figure 1. Diagram of experimental system.
onp position ~ o b e~
e
0
c
h
7
- 53cm
Temperature Robe Reach +45
Reactor
Inlet
5 8
9 31
4 23.9 26.2 94
3 23.4 25.5 92
2 22.9 24.8 91
I 19.2 21.0, 76
Heater No. . Length, cm
Resistance, R No. of turns
Figure 2. Reactor/furnace assembly.
which received their signals from individual three-term controllers. The controllers, in turn, could receive input signals from any of the wall thermocouples within that heating zone. The set points and control constants were supplied manually; thus, there was considerable flexibility in the imposition of a desired axial temperature profileisothermal or otherwise. The reactor-furnace assembly is illustrated in Figure 2. The main feature of the configuration was a set of probes which could be inserted vertically into the reactor space to measure directly the temperature and camposition of the process stream as functions of both radial and axial position. It was primarily to accommodate the probes with a minimum disturbance to the flow pattern that the reactor diameter was set at 50 mm, considerably larger than that used by previous workers. The probes were inserted from the exit end of the reactor through a small hole in a platelet which was clamped to the bottom flange of the reactor tube. A special mechanism which provided horizontal movement of the entire probe assembly allowed continuous radial traversal of the reactor at any axial position. The mechanism is described in more detail in the work of Beshty (1978).
Design of the Probes. Design of the probes was based on measurements at the reactor center line under the conditions used in the experiments for butane pyrolysis. In these experiments, described in the following paper in this issue (Kershenbaum and Leaney, 1986), the feed to the reactor was generally 5-10 mol % butane in nitrogen at a constant reactor pressure of 1.5 bar. In these experiments, flow was in the laminar regime (Re -150). The system was designed to operate at Reynolds numbers up to 2500, but the cost of raw materials prevented us from working at higher flow rates in these early experiments. Temperature measurement within the reactor space was accomplished by a minerally insulated thermocouple probe. The probe consisted of a 3 mm 0.d. outer inconel sheath which covered yet remained electrically insulated from the thermocouple tip. The sheath was supported by an outer tube, 5 mm o.d., constructed from seamless 321 stainless steel which ended 100 mm from the probe tip to minimize the thermal disturbance. Corrections of the order of 15 "C (depending upon the flow rate and temperature) were made for radiation from the reactor wall onto the probe tip, using well-known heat-transfer principles (Leaney, 1982). Axial conduction along the ther-
Ind. Eng. Chem. Process Des. Dev., Vol. 25, No. 3, 1986 I
750
1
I
I
1
I
I
I
783
750
W L L TEMPERATURE
740-
-B---.--ICL-m-+
740 a-O--o--Q-"--
0
730-
,O w
5
5
730
GAS TEMPERATURE
720-
720
-
710
700-
700
-
690
710
w
0
;
690
sso: 50
I
I
I
I
1
60
70
80
93
100
I
110
I
120
,680
I
130
LW2
I
$ I
2 RW
RADIAL POSITW (cm)
AXIAL DISTANCE FROM REACTOR lNLET(cm)
Figure 3. Typical (A) axial and (B)radial measured temperature profiles. Table I. Construction and Testing dl,b 4, 4, 4, 4 , mm mm mm mm mm probe I 0.38 0.80 3.50 4.00 5.33
of the Gas Sample Probe" de, probe mm length,m details 6.35 1.392 converging-diverging nozzle -0.2-mm diameter constructed a t probe tip; repeated blockage of capillary after 2-min operation probe I1 0.38 0.80 3.50 4.00 5.33 6.35 1.392 no nozzle; repeated blockage of capillary after 2-3-h operation probe I11 0.55 1.02 2.69 3.51 4.52 6.35 1.392 no blockage after maximum period of test, -6 h; test repeatable
-
"Material of contruction: 321 stainless steel. Reactor operating conditions: 700 "C, 7.psig; 10% C4HIoin N2 at flow rate of *For nomenclature, see Figure 4.
mocouple probe was unimportant because of the isothermal zone downstream of the thermocouple tip. The ability to attain an isothermal gas region is illustrated by a typical axial and radial profile in Figure 3. The isothermal regions attained were usually 40 cm in length, for which the gas temperature both radially and axially generally varied by less than 4 "C, and continuous center-line measurements showed the local variation to be less than 1 "C. The gas sampling probe was designed to provide measurements which were representative of the composition within the reador space. Preliminary investigation showed three major sources of potential error. (i) Aerodynamic effects exist because the presence of the probe and the flow through the probe disturb the flow pattern at and around the probe tip. A theoretical analysis by Rosen (1954) presented an expression for the fractional flow distortion, as a function of the distance perpendicular to the sink and the dimensionless sampling rate. (ii) The heat sink caused by the presence of the probe can disturb the flow and temperature fields especially near the probe tip. The degree of thermal effect would depend primarily on the design features of the probe. (iii) Reaction occurring in the probe can lead to two problems: carbon deposition on a catalytically active surface and the continuation of a homogeneous reaction if adequate quenching is not achieved. The two classical techniques for gas sampling, principally associated with the study of flames, were analyzed but discarded as being unsuitable here. In one method, isokinetic sampling, the gas is removed without any change in its velocity, thereby ensuring no aerodynamic disturbance to the flow field; quenching of the reaction is achieved by a coolant surrounding the capillary. The low axial gas velocity would require an efficient cooling system which
m3/s.
would prove difficult to design without introducing serious thermal effects on the surrounding flow field. The alternative method, aerodynamic sampling, demands the ability to maintain supersonic flow downstream of a converging-diverging nozzle situated at the probe capillary tip. Quenching is obtained by the decrease in sensible heat to accommodate the corresponding increase in kinetic energy. For typical L I D ratios ( 1000) for the probe capillary, frictional effects caused by the capillary walls make it impossible to achieve supersonic flow downstream of the diverging portion of the nozzle (Shapiro, 1953). Furthermore, subsonic flow conditions would lead to only a small drop in temperature and require excessively small capillary and nozzle diameters in order to minimize the disturbances to the flow pattern; small diameter nozzles, however, rapidly become blocked with carbon under reactive conditions. Eventually, a compromise design was adopted which consisted of a straight bore capillary tube with two other concentric tubes for the flow of the nitrogen coolant as shown in Figure 4. Several such probes with different dimensions were constructed and tested under typical reactor operating conditions. Table I indicates some of the results; at temperatures above 700 "C, the probe capillary had to be greater than 0.5 mm in order to avoid the risk of blockage by carbon over the period of an experimental run. Analysis of the Sampling Probe. The uncertainties associated with the sampling probe were investigated by an analysis of models of its behavior. Temperature, composition, and pressure profiles within the probe could be obtained from the solution of energy, mass, and momentum balances in the three concentric tubes; these are expressed as a set of seven coupled ordinary differential N
784
Ind. Eng. Chem. Process Des. Dev., Vol. 25,
TUBE1
No. 3, 1986
I
TUaE3
TUBE2
\
1818 STNNLESS STEEL WIRE TO CENTRE TUBES I PND 2
INERT GAS (ARGON) WELDED
ARC
Figure 4. Design and construction of sampling probes.
equations with two-point boundary conditions and are presented in detail by Leaney (1982). The equations could be solved by integration from one end but required simultaneous iteration on initial conditions for three of the model variables including the unknown (but measurable) flow rate through the capillary. A consideration of the orders of magnitude of each of the terms in the original equations led to a partial decoupling of equations and to the following simplified model: -dT1 - - ho2rd3A1(T1 - 7'2) energy (1) dl GlC, dT2 -- dl
h01Td5A2(Tp
G2Cp dT, dl
momentum
- 7'2)
+
h02d3Az(T1 - 7'2) G2Cp
- ho3~diA,(Ti- Tc) GCCP
- - Gi2CF(d3 + dz) dl
%A1
dP2 - - -GZ2Cf~(d5+ d l )
dl
2P2A2
dp,
-=dl
(2)
(3)
mass
r.
= U,
(A2
~ r ' i . y ( h 2
+ 1)1/2
+ 1)V- X(y - 1) dA
(8)
where x is related to the fractional flow distortion, f (the ratio of the radial velocity of fluid elements to the nominal gas velocity), by
(4)
(5)
-2cfG: Pcdl
generates temperature, pressure, and composition profiles within the capillary and enables estimation of the amount of homogeneous reaction within the probe. A second source of error, the aerodynamic error, is introduced by the probe because it has sampled elements which have traveled at an increased velocity due to the sampling process and with a radial as well as an axial velocity component. Rosen (1954) has given a detailed analysis of the fluid dynamics; he has also presented a simplified result that the residence time, 7,of an element traveling to the probe from a distance x along the central axis is given by
(6)
(7)
where F b = GCCb/p, with boundary conditions at 1 = 0 (probe tip) T2 = T,; Pz = P,; Tc = TI; P, = P,
Fb = GcCbr/pr where Gc is unknown at 1 = L (probe end)
P, = P2 = Pexit (normally 1 bar) PI = P1 inlet 7'1 = TI inlet. With these equations, each set of boundary conditions leads to a stable solution and the iteration of the unknown conditions at one end can be performed sequentially rather than simultaneously. The solution of the model equations
and y is the dimensionless sampling rate, Q/4rr,2um.Thus, it was possible to estimate the mean velocity of elements approaching the probe tip, taking into account the aerodynamic disturbance. The third potential source of error (thermal error caused by the presence of the probe) could likewise be quantified in terms of a reduction in the temperature of the gas phase. However, calculations showed that the depression of the gas temperature w a only 1 OC as close as 1.5 mm from the probe tip; the effect was too small to have a significant influence on the measured composition. Simulated Results and Discussion In order to quantify the total error, each disturbance caused by the sampling probe was represented as that displacement of the probe tip from its axial position which would completely compensate for the error introduced. Fortunately, the two main sources of error tended to act in opposite directions and led to a partial cancellation-the reaction error moves the effective sampling point downstream of the probe tip because of reaction within the probe and the aerodynamic error moves the effective sampling point upstream because of the increased velocity of the sampled elements. In order to obtain numerical results, two sets of simulations were performed at nominal gas temperatures of 700
Ind. Eng. Chem. Process Des. Dev., Vol. 25, No. 3, 1986
785
Table 11. Principle Results from the Gas Sampling Probe Model simulation isothermal zone temp, "C 700 700 700 800 800 800
simulated position of the sampling probe tip top of the isothermal gas region" 200 mm below the top 400 mm below the top
aerodynamic error, mm 1.76
reaction error, mm -1.32
summation of sampling errors, mm 0.44
2.01
-0.33
1.68
2.36
-0.04
2.32
top of the isothermal gas region 200 mm below the top 400 mm below the top
1.69
-1.68
0.01
1.89
-0.45
1.44
2.22
-0.06
2.16
Assumed to be 500-mm length, situated 625 mm from the exit of the reactor.
1
Y
li 4//
- a 1.5
-1.4
2 W
5cQ
. E
- 1.2
11
- 1.1
200
-1.0 15
IO
05 DISTANCE FROM PROBE TIP (M)
0
Figure 5. Simulated temperature and pressure profiles within the sampling probe.
and 800 "C. In each set, sampling was simulated for three different axial positions along the center line within the isothermal region. The nitrogen coolant was used at a constant flow rate of m3/s (1bar, 300 K) for all the simulated sampling positions. A typical set of temperature and pressure profiles within the probe are presented in Figure 5 for a reactor temperature of 700 "C and for the probe tip situated at a point 0.345 m from the end of the isothermal region. From a knowledge of the temperature and pressure (and, hence, velocity) profiles, the amount of reaction taking place within the probe is easily calculated by using approximate values of the rate constant. As a check on the accuracy of the simple model, the predicted sample flow rate through the capillary was compared with that measured experimentally under identical conditions. The results differed by less than 10% despite the fact that the flow rate depended strongly upon the axial position in the reactor. For the case illustrated in Figure 5, the predicted flow rate through the capillary was 1720 m 3 / sand the measured value was 1580 mm3/s. The net effect of all sampling errors is illustrated in Table I1 for various axial positions within the isothermal region. It shows that the positional error along a total axial
length of 400 mm of reactor space is less than 3 mm; this is of the same order as the error in placing the probe in position and generally less than the error associated with chromatographic measurement of composition. Corrections for Laminar Flow. Under suitable flow conditions, the data resulting from the experimental system outlined above consist of a set of gas compositions as a function of axial position over an isothermal region of the reactor. When there is complete radial mixing and negligible axial mixing, the data can be simply analyzed by the straightforward treatment of an isothermal plug flow reactor. However, in the laminar flow regime (Re 150), there could be significant radial variation in both the temperature and composition at any axial position. Even for an isothermal radial profile, the composition would generally vary because of the radial distribution of the residence times. The amount of variation will depend to a large degree on the extent of segregation in the radial direction. Clearly, if radial mixing is good, then the behavior of the system will approach that of the plug flow reactor despite the existence of, say, a parabolic velocity profile. Leaney and Kershenbaum (1986) have investigated the estimation of rate constants from measurements in laminar flow reactors; they have found that the estimates are very similar to those obtained by using a plug flow model except under conditions which approach complete segregation. Specifically, in the case of the pyrolysis of hydrocarbons at 600-800 OC, the error introduced by using a plug flow analysis to estimate an overall rate constant is generally less than 5%. In the following paper in this issue (Kershenbaum and Leaney, 1986),we illustrated the measured radial composition profiles to show that they are consistent with the above approach.
-
Conclusions A careful experimental design and an accurate analytical basis have been developed for studying the kinetics of fast complex reactions. The main feature of the system is the use of probes to sample the temperature and composition within an isothermal section of the reactor. Analysis of the sampling errors involved shows them to be less than those associated with the measurement of the chemical composition by chromatography. I t is felt that this experimental reactor system can eliminate some of the abnormal behavior often associated with pyrolysis reactions. Conflicting data obtained by several workers over apparently similar operating conditions may well have been caused by experimental condi-
706
Ind. Eng. Chem. Process Des. Dev. W06, 25, 706-794
tions which have not been well-characterized and which may not be similar at all. Elimination of many of the experimental uncertainties can be achieved by direct measurement of the temperature and composition within the reactor space. Data generated by such experiments should have signiticantly leas scatter and may be correlated by simpler rate laws, presumably because some extraneous mass- and heat-transfer effects have been eliminated. The alternative approach of varying the process stream flow rate in order to vary the residence time and measuring only the reactor wall temperature and the composition of the reactor effluent may lead to significant errors in the apparent rate constants and has the additional problem of uncertain and definitely nonisothermal reaction zones. In the following paper in this issue (Kershenbaum and Leaney, 19861, the techniques presented here are applied in a study of the pyrolysis of butane at 650-750 OC.
Nomenclature A,, A2, A, = cross-sectional areas of probe tubes (see Figure 4)
= butane concentration in the probe and reactor C = specific heat capacity = friction factor dl-de = probe diameters (see Figure 4) Fb = flow rate of butane in the probe capillary f = fractional flow distortion G1, G2,G, = mass velocities in sections 1,2, and c of the probe, respectively hot,hoz,hO3= overall heat-transfer coefficients k b = pseudo-first-order rate constant for butane pyrolysis 1 = length L = total length of the probe P,, P2,P,, P, = pressure in sections 1, 2, and c of the probe and in the reactor, respectively Q = volumetric sampling rate r, = radius of probe capillary = d1/2 T,, T2,T,, Tr = temperature in sections 1,2,and c of the probe and in the reactor, respectively cb, cbr
4
Tp = outer wall temperature of probe u, = nominal linear velocity in the reactor x = distance from the sample probe tip
Greek Symbols dimensionless sampling rate X = dummy variable of integration p l , p2, p,, p r = gas density in sections 1,2, and c of the probe and in the reactor, respectively r = residence time of sampled fluid elements y =
Literature Cited Allara, D. L.; Edeleon, D. Int. J . Chem. Kinet. 1975, 7 , 479. Beshty, B. S. Ph.D.Thesis, University of London, London, England, 1978. Buekens, A. G.; Froment, G. F. Ind. Eng. Chem. Process D e s . D e v . 1988, 7 , 435. Edelson, D.; Ailara, D. L. Int. J . Chem. Kinet. 1980, 12, 605. Frey, F. E.; Smith, D. F. Ind. €ng. Chsm. 1928. 20, 948. Van den Berghe, P. J.; Goossens, A. G. Froment, G. F.; Van de Steen, B. 0.; AICM J . im.23.93. -&a&< 0 . J:; Santwo, R. J.; Dryer, F. L.; Glassman, I . Int. J . Chem. Kinet. 1981. 13. 149. Kershenbaum, L. S:; Leaney, P. W. Ind. Eng. Chem. Process. D e s . Dev., foHowhg paper In this Issue. Kershenbaum, L. S.; Martin, J. J. A I C M J . 1987. 73, 148. Laldler. K. J.; Sa@, N. H.; Wojclechowski. B. W. R o c . R . Soc. London, Ser. A . 1982, A270, 242. Leaney, P. W. PhD Thesis, University of London, London, England, 1982. Leaney, P. W.; Kershenbaum, L. S. Ind. Eng. Chem. Process D e s . Dev., submitted. Martln, R.; Dzierzynskl, M.;Nlcleua, M. J . Chem. Phys. 1984, 6 1 , 286. Powers, D. R.; Corcoran, W. H. I d . Eng. Chem. Fundam. 1974. 13, 351. Purneii, J. H. I n "Frontiers of Free Radical Chemlstry"; Academic Press: London, 1980; p 83. Rosen, P. "The Potential Flow of a Fluid into a Sampling Probe"; John Hopklns UniversJty: Baltimore, 1954; Applied Physics Laboratory Report APLIJHU, CF-2248. Schnelder, V.; Frollch, P. K. Ind. €178. Chem. 1931, 2 3 , 1405. Schutt. H. C. Chem. Eng. Prog. 1947, 4 3 , 103. Shaplro, A. H. "The Dynamics and Thermodynamics of Compressible Fluld Flow"; Ronald: New York, 1953; 2 Vols. Sundaram, K. M.; Froment. G. F. Chem. Eng. Sci. 1977, 3 2 , 601. Sundaram, K. M.;Froment, G. F. Ind. €ng. Chem. Fundam. 1978, 17, 174. Van Damme. P. S.; Narayanan, S.; Froment, G. F. AIChE J . 1975, 2 7 , 1065.
-.--.
Received for review October 26, 1983 Revised manuscript received December 18, 1985 Accepted January 22, 1986
Pyrolysis of Hydrocarbons in a Large Pitot-Scaie Reactor. 2. Pyrolysis of n-8utane Lester S. Kershenbaum' and Patrlck W. Leaneyt Department of Chemhl Engineering end Chemhl TednWbgy, Impenlcl Colhsge, London SW7, England
A study of the kinetics of the pyrolysis of n-butane in the temperature range 640-740 O C has been conducted in a pilot-scale Inconel reactor (1.5 m long X 52.5 mm 1.d.). When a furnace with five independently controlled heating zones was used, an isothermal gas zone could be achieved at any desired temperature. Probes inserted into the reactor space gave detailed axial and radlal composition and temperature profiles within the Isothermal section. Sets of experhnents were performed at constant thrwshput rates for differentisothermal gas temperatures leading to conversions between 10% and 80% for feed compositions of 5-10 mol % butane in nitrogen at 1.5 bar. The global rate of the n-ne decomposition was welcfltted as a first-order expresdon with a single activation energy of 212 kJ/mol, independent of the conversion up to values as high as 70%. The reaction was modeled by an appropriate set of free-radlcal steps using literature values of the rate constants. Excellent agreement was obtained between experimental and predicted product distribution as a function of n-butane conversion, although predicted decomposition rates were lower than those observed experimentally.
Despite the significant amount of work done on butane pyrolysis, uncertainties remain in the kinetics of the re+Current address: Chevron Research Co., Richmond, CA. 0196-4305/86/ 1125-0788$01.50/0
action and in the factors which affect the distribution of important producta such as ethylene and propylene. Early work (Pease and Durgan, 1930; Cambron, 1932; Steacie and Puddington, 1938) drew attention to the parameters which needed special consideration: reactor surface-to-volume 0 1986 American
Chemical Society
