Nt = over-all number of transfer units based on weight ratio
driving force in fluid phase = quantity of solute adsorbed between breakthrough point and exhaustion point, lb. = superficial velocity of feed, ft./hr. = W e - W b , lb. = quantity of solute-free effluent a t breakthrough point, lb. W e = quantity of solute-free effluent a t exhaustion point, lb. = fraction of voids within a sieve pellet x = concentration of solute on adsorbent in equilibrium X8 with feed of concentration Yo,lb. solute/lb. adsorbent Y = concentration of fluid stream a t any point in column, lb. solute/lb. solvent = concentration of fluid stream a t breakthrough point, y b lb. solute/lb. solvent Ye = concentration of fluid stream a t exhaustion point, lb. solute/lb. solvent Y* = concentration of fluid stream in equilibrium with adsorbent a t any point in column, lb. solute/lb. solvent Z = height of packed bed, ft. za = height of adsorption zone, ft. = density of solvent, lb./cu. ft. Pb = bulk density of activated sieves, lb./cu. ft. Ps
(2) Carson, D. B., Broughton, D. B., Petrol. ReJiner 38 (4), 130-4 (1959). \ . _ _ _
(3) Cheml Eng. 70 (19), 69 (1963). (4) Dryden, C. E., Strang, D. A., Withrow, A. E., Chem. Eng. Proer. 49 (4). 191-6 (1953). (5) Furanz, Wl‘R., Christens‘en, E. R., May, J. E., Hess, H. V., Petrol. Rejner 38 (4), 125-9 (1959). (6) Glueckauf, E., Coates, J. J., J . Chem. Sac. 149, 1315 (1947). (7) Griesmer, G. J., Rhodes, H. B., Kiyonaga, K., Petrol. ReJiner 39 (6), 125-9 (1960). (8) Henry, J. P., Chermakesavan, B., Smith, J. M., A.Z.CI1.E. J . 7 ( I ) , 10 (1961). (9) Johnson, P. A., Babb, A. L., Chem. Reus. 56, 387 (1956). (10) Kehat, E., Rosenkranz, Z., IND.ENG.CHEM.PROCESS DESIGN DEVELOP. 4, 217-20 (1965). (11) Michaels, A. S., Znd. Eng. Chem. 44, 1922-30 (1952). (12) Reid, R. C., Shenvood, T. K., “Properties of Gases and Liquids,” pp. 283-99, McGraw-Hill, New York, 1958. (13) Roberts, P. V., Ph.D. thesis in chemical engineering, Cornel1 University, Ithaca, N. Y . ,1966. (14) Schumacher, W. J., Ph.D. thesis in chemical engineering, Cornell University, Ithaca, N. Y . , 1964. (15) Treybal, R. E., “Mass Transfer Operations,” pp. 54-5, 504-8, McGraw-Hill, New York, 1955. (16) Tsuruzumi, Akie, Bull. Chem. Soc. Japan 34, 1457 (1961). (17) Wheeler, A., Advan. Catalysis 3, 249-327 (1951). (18) Ziegenhain, W. C., Refining Engr. 29, C-6-C-12 (August 1957).
literature Cited (1) Barry, H. M., Chem. Eng. 67 (3), 105-20 (1960).
RECEIVED for review July 25, 1966 ACCEPTEDFebruary 2, 1967
OLEFIN CHLORINATION IN HOMOGENEOUS AQUEOUS COPPER CHLORIDE SOLUTIONS M. L. SPECTOR, HEINZ HEINEMANN, AND K. D. MILLER
M . W. Kellogg Co., Piscataway, N . J . The application of liquid phase homogeneous catalysis to the oxyhydrochlorination of ethylene is characterized by virtually perfect selectivity. This is attributed to the ease with which a uniform temperature is maintained and to the mild reaction conditions necessary to achieve economic space time yield.
react with aqueous copper halide solutions to multihalogenated paraffins. T h e reaction of ethylene with copper chloride in aqueous solution to produce 1,2-dichloroethane has been examined in detail. Simultaneous introduction of oxygen and hydrogen chloride regenerates the system. Thus, the over-all system is a homogeneous catalyst for the oxyhydrochlorination of ethylene as described below : LEFINS
0 produce
+ 2CuClz 2CuC1 + 2HC1 + ‘ / z C2H4
+
CzH4C12
0 2
2CuC1
(1)
HzO
(2)
+ HzO
(3)
+ 2 C ~ C l zf
OVER-ALL. Aq. CuCI, CuCli
CzH4
+ 2HC1 + ‘/z
0 2 + CzH4Clz
I n view of the established process for thermally decomposing 1,2-dichloroethane to produce vinyl chloride and hydrogen chloride, the oxyhydrochlorination of ethylene as described in Reaction 3 is a vital link in the production of vinyl chloride from ethylene. This investigation has formed the basis for
further development, which in turn has led to a commercial route. Experimental
Materials. ACS reagent grade salts. C.P. ethylene and oxygen. Apparatus. T h e reactor system (Figure 1) consists of a glass reaction vessel, A , suspended from the top portion of a metal flange, B, connected to an exit line, C, by means of a metal ball socket welded to the head of the flange. T h e glass reactor is encased in an electrically heated metal shell, D. The heat input, and consequently the temperature, are controlled by an automatic temperature regulator. Gaseous hydrocarbon is introduced into the metal shell through an appropriate opening, E, and enters the glass reactor by passage through a glass tube, F, which is fitted with a fritted-glass foot. This arrangement provides for pressure equilization on both sides of the glass reactor. The hydrocarbon gas comes in contact with the chlorinating solution a t the bottom of the reactor and bubbles u p through the medium; volatile products, water vapor, and unreacted hydrocarbon leave via line C, equipped with a valve, H, which serves to reduce the pressure to atmospheric. Then the exit gases are passed into a n appropriate trapping VOL. 6
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D Ot
C Ft
INLET
6
W
7.0
8 5.0
0 I
A TIME (MINJ Figure 2. Determination of duration for typical run Run 80
Figure 1.
Experimental apparatus Discussion of Data
system to remove the product. The experimental apparatus is also equipped with a bypass arrangement, G, which permits passage of gas through the pressure zone without passing through the aqueous chlorinating medium. I n the case of simultaneous regeneration with oxygen and hydrogen chloride, the system was modified to accommodate a direct introduction port, F ' , through which a 4 to 1 HCl-02 mixture was fed. Procedure. T h e chlorinating solution was prepared by dissolving a weighed amount of cupric chloride and cuprous chloride in water with constant heating and stirring. Additional water was added to adjust the volume to obtain the desired molar concentration. The chlorinating solution was charged, while hot, to the glass reactor. The catalyst volume was 0.375 liter in all cases unless otherwise noted. The reactor and contents, with bypass open, were preheated for 16 hours at reaction temperature under sufficient pressure of nitrogen to prevent the solution from boiling. Prior to beginning a run, nitrogen was purged from the system with hydrocarbon feed which passed over, but not through, the aqueous medium by successively pressuring the system with the hydrocarbon and depressuring to remove the nitrogen. Upon completion of the purge, the bypass arrangement was closed, forcing the hydrocarbon feed through the chlorinating solution. A run was begun after an initial line-out period, necessary to equilibrate the system. R u n duration is defined as the period of time during which an essentially constant conversion is maintained. This period was determined by gas chromatographic analysis of the reactor exit gases. Figure 2 shows a plot of mole conversion us. time, which served as a basis for determination of run duration. Consequently, the results obtained represent the average value for the run period a t essentially constant reactivity, as illustrated by Figure 2. The relatively nonvolatile products in the exit gases were condensed by means of cold traps; the remaining gases were passed through a sodium hydroxide scrubber solution to remove carbon dioxide and finally through a wet-test meter. Effectiveness of the trapping system was determined by gas chromatographic analysis of exit vapors before they entered the basic scrubber solution. The products isolated were weighed and analyzed by standard chemical and instrumental methods of analysis. Instruments used for gas chromatographic analysis were an F&M Model 500 and a Perkin-Elmer Model 154 Vapor Fractometer. The column used in both cases was didecyl phthalate on C-22 firebrick. I n a system where reaction occurs only in the liquid phase and it is necessary to dissolve one or more of the reactants in the liquid, there is danger of being diffusion-controlled. This was avoided in the present case by use of a coarse porous frit to disperse the gas and by running a t a total pressure of 300 p.s.i.g. 328
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P R O C E S S DESIGN AND DEVELOPMENT
The data obtained for the copper chloride-ethylene system are listed in Table I, and the following discussion is based largely on these data. Reproducibility of experimental results was checked for several runs in order to determine the degree of confidence which could be placed upon the experimental data. Several sets of duplicate runs were conducted (runs 53 and 67, 63 and 66, and 73 and 75) ; results agree well within 10% in all cases. Criterion for Reaction Rate. It was established that the best criterion for measuring rate was the space time yield (STY), here expressed as moles of product produced per liter of solution per hour. Runs 66 and 67, made under otherwise similar conditions, show that when the flow rate was changed by a factor of 2.5, the space time yields remained relatively constant a t 0.40 and 0.46 mole of product per liter of copper chloride solution per hour. Since the STY was also independent of conversion, 2,7% us. 8.1% in these two runs, it was used as the prime indication of reaction rate. Temperature. Ethylene reacted with a solution of initial composition 6.OM CuC12-0.2M CuCl at 140°, 150°, and 160' C. (runs 55, 56, and 57); a 7.OM CuC12-0.5M CuCl solution a t 150' and 160' C. (runs 53, 62, and 67); and a 7.OM CuC12-2.0M CuCl solution a t 140' and 150' C. (runs 71 and 72). I n all cases the reaction rate, as measured by STY, increased with temperature and an Arrhenius plot of STY us. reciprocal of T o K. indicates an activation energy of 22 kcal. per mole, which is considerably in excess of the 3 to 7 kcal. per mole commonly associated with diffusion-controlled reaction processes ( 7 ) . Since the percentage of CuClz depletion was not identical in all cases, and this has an appreciable effect on rate, the energy of activation must be regarded as only approximate, but demonstrates that the system is not diffusionlimited. Another effect of temperature was evidence of C O Zin the effluent gas when reaction temperature reached 177' C. While the CO1 produced was in trace quantity at this temperature, competing reactions are indicated at this and higher temperatures. Selectivity. The selectivity to 1,2-dichloroethane was about 99% in all cases (Table I). The balance is largely ethyl chloride. Theoretical. The following hypothesis is advanced to account for the chlorinating reaction: cucl
+ excess CuCl2 $ CuC1~2CuC12+ free CuCL
(4)
Table I.
Summary of Data for Aqueous Phase Chlorination Reaction on 0.375 Liter of Solution (2C~C12 C2H4 - CzHdClz 2CuCl)
+
+
%d
Run 12‘0.
Initial Composition of Chiof inating Solution, M CuCl 2 CUCI 7 0.5 6 0.2 6 0.2 0.2 6 7 0.25 0.16 4.5 7 0.5 7 0.5 7 0.5 7 0.5 7 0.5 7 2.0 2.0 7 7 1. o
c2H4 Calcd. CzH4, P.S.I.A.a 273 271 258 282 274 262 273 273 273 273 273 273 282 272
Tfmp,: C. 150 150 160 140 150 150 160 150 150 150 150 150 140 150
Calcd. Run Flow H z O , ~ Duration, Rate, P.S.I. Min. MoleslHr. 42 60 2.2 60 2.3 44 57 60 2.3 33 60 1.7 41 60 2.3 53 76 2.3 42 30 4.5 42 60 4.0 42 35 5.4 42 45 2.1 42 45 2.2 42 60 2.0 33 60 2.0 43 45 ...
Au. Mole 70 Conu.
8.4 53 4.3 55 5.7 56 2.5 57 6.0 59 1.2 60 7.1 62 3.3 63 2.7 66 8.1 67 5.6 70 7.3 71 4.2 72 ... 73 73 77 80 81 a Total pressure minus partial pressure of water. Based on extrapolation of atmospheric boiling point data. c STY = mole yoconv. X (CzH4) moles/hr./volume of chlorinating solution liters = molesproduct/liter/hr. * Output basis. * STY based only on product isolated.
I n this reaction all components refer to water-soluble species. Soluble complexes of CuCl are well known (2). I n an excess of chloride ion the complex CuC13’- would be expected to predominate over CuClQ-. “Free” CuC12 is defined as the molarity of the total cupric ion minus twice the molarity of cuprous chloride. Gaseous C2H4
soluble C2H4
Soluble C2H4 f. CuCl e2CuC12 eCuCl.2CuC12.C2H4
+ 2 “free” CuC12 CzHzC12
(6)
-f
+ C U C ~ . ~ C fU C2CuCl ~~
(7)
I t is further proposed that Reaction 7 is rate limiting. Under conditions where there is excess CuC12 and C2H4, CuCl becomes the minimum chemical and the concentration of complexed ethylene is proportional to CuCl. If the above postulation holds, the rate-limiting equation becomes
d(DCE) = k (CuC1) (CuC12 dt
-
2CuCl)’
where the latter term describes “free” or uncomplexed CuC12. Equation 8 may be treated algebraically as follows:
- y ) 0, - 2 t - y ) 2 = k ( l - y) (3y - 2t)2
STY = k ( t
where x =
CuCl, moles per liter
y = CuC12, moles per liter
t = total Cu, moles per liter
x = t - y
STYC 0.50 0.26 0.35 0.12 0.37 0.07 0.85 0.36 0.40 0.46 0.48 0.39 0.22 0.54e
Copper Utilization, yc 14.2 8.6 11.7 3.8 10.5 8.9 12.1 10.2 6.5 9.9 10.3 11 .o 6.4 11.7 11.7 7.4 14.4 5.3
STY is expressed in moles of 1,2-dichloroethane produced per liter of chlorinating solution per hour. Then: STY = kt3 (1
-y/t)
(3y/t
- 2)’
(9)
(5)
Reaction 6 describes the formation of a complex between solubilized cuprous chloride and dissolved ethylene and Reaction 7 proposes the reaction of this complex with 2 moles of “free” CuClz to produce 1,2-dichloroethane. CuCl .2CuC12 C2H4
Selectivity to 7,2Dichloroethane 98.4 99.0 99.0 99.5 99.3 99 .O 99 .O 99.1 99.0 99.4 99.0 99.6 99.8 99.1
Upon inspection of Equation 10, it is evident that STY approaches zero, when y / t , or mole fraction cupric ion, approaches 0.666 or 1.00. Figure 3 depicts the resultant curve when values of y / t between 0.666 and 1.O are substituted into the function (1 - y / t ) ( 3 y / t 2)2 and plotted against y / t . The ordinate is equal to relative activity, A , which is STY/kt3. The STY at any given mole fraction cupric content of a run is the product of the function plotted in Figure 3 and kt3. With the use of Figure 3, it is possible to test the validity of Equation 8 by graphically integrating between the initial and final mole per cent cupric content for all runs a t a partial pressure of ethylene of 270 i 12 p.s.i. and flow rates of from 2 to 4 moles per hour. Run 53 was selected as the datum point from which k at 150’ C. was determined. The values of k a t 140’ and 160’ C. were obtained from the value of k at 150’ C. and by use of the E, of 22 kcal. per mole. T h e agreement between calculated and observed STY may be seen from Table I1 and is plotted in Figure 4. It is apparent that the fit is good at all points except where the initial mole fraction of cupric ions is close to or below 0.75. This degree of fit lends confidence to the proposed reaction sequence ; however, this does not say that the proposed sequence is the only possible one. For instance, in the three cases (runs 71, 77, and 80) where the final mole fraction Cu2+/Cu is considerably below 0.75, where the predicted value is much lower than the observed, it is likely that the free CuCl2 is not as low as defined in Equation 4, but rather there may be a second complex such as CuC1.CuCl2, which does not bind as much CuC12. T h e equation and resultant curve are intended only to serve as a model and to be used as a basis for prediction and understand-
-
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a
Table II. Calculated Average STY from Equation STY = k[CuCI][CuCI~-2CuCI]* and E, = 22 Kcal./G. Mole Mole Fraction Cupric Run Initial Final Total Cu, Temp., Calcd. Obsd. No. (Y/t)l (Y/t)2 Moles/L. a c. k STY STY 55 0.968 0.884 6.2 150 0.0269 0.25 0.26 60 0.965 0.880 4.66 150 0.0269 0.12 0.07 59 0.965 0.865 7.25 150 0.0269 0.44 0.37 67 0.934 7.5 150 0.0269 0.844 0.54 0.46 70 0.935 0.837 7.5 150 0.0269 0.53 0.48 * 53" 0.935 0.803 7.5 150 0.0269 0.50 10.05 150 0.0269 80 0.878 0.752 0.89 0.86 8.0 150 0.0269 75 0.875 0.744 0.50 0.55 73 0.875 0.744 8.0 150 0.0269 0.50 0.54 0.685 9.0 150 0.0269 71 0.778 0.12 0.39 0.695 8.0 150 0.0269 77 0.750 0.09 0.30 0.631 10.5 150 0.0269 0.0 81 0.666 0.25 6.2 160 0.0460 0.47 0.852 56 0.968 0.35 0.85 62 0.935 0.824 7.5 160 0.0460 0.86 0.12 6.2 140 0.0141 0.12 0.930 57 0.968 0.17 0.23 9.0 140 0.0141 0.729 72 0.778 Run used to determine k at 1.50' C.; k values at 740" and 760" C. determined from E, = 22 kcal./mole.
.048--
,044,040
--
,036-
:;I/
,032,028-
.024--
.020--
0
0.I 0 Y : 0 0.1
. O 16-
I
:
0.2 0.3
:
I
:
:
I
:
0.4
0.5
0.6
0.7
0.8
0.9
1.0
STY CALCULATED .012--
Figure 4. Correlation of calculated with observed STY ,008-
0 0
0
.004--
.6 4
.?0
.76
.a2
.88
34
1.00
MOLES Cue* TOTAL MOLES CU
Figure 3. Function of calculated activity with respect to mole fraction of cupric ion
ing. The following may be deduced from Equation 8 and Figure 3. At a given mole fraction cupric ion, the activity increases as the cube of the total copper ion concentration. An 8 to 1 ratio of cupric to cuprous ion is optimum with respect to activity. The change in activity due to depletion of cupric ion may be anticipated from Figure 3 when the final mole fraction of cupric ion is 0.75 or higher. The above hypothesis assumes that the concentration of noncomplexed ethylene is negligible. As agitation and partial pressure of ethylene increase, the dissolved noncomplexed 330
l&EC PROCESS D E S I G N A N D DEVELOPMENT
14OoC. 150DC. 160° c.
ethylene may become appreciable relative to complexed ethylene. In this case, that portion of the over-all STY due to reaction of uncomplexed ethylene would be expressed by a different kinetic term. Regenerative Operation. The feasibility of simultaneous chlorination of ethylene via an aqueous CuCl&uCl reaction system and subsequent in situ regeneration of the copper chloride solution as described in Equations 1 and 3 was demonstrated. Ethylene was selectively chlorinated via an aqeuous 7 M CuClS-lM CuCl solution with in situ addition of 0 2 and HCl to regenerate the reduced cupric chloride. During a 24hour run, the chlorinating solution was passed through several regeneration cycles to maintain or restore the activity of the chlorination reaction. Product isolated from the run corresponded to greater than 100% utilization of the initial cupric chloride charged to the reactor. During this time water was added to maintain liquid volume. The selectivity to 1,2-dichloroethane was 99% except when the system became acid because of excessive buildup of HCl. Under these circumstances, the ethyl chloride make increased appreciably, but returned to normal when the excess was dissipated.
System Poisons. In an effort to detect potential poisons, runs were made in which Hz, CO, and ethane were substituted for 10% or more of ethylene. No adverse effect, other than a reduction in STY proportional to the decrease in partial pressure of ethylene, was noted. I n another run, the system was made 0.1M sulfide; here, too, no adverse effect was noted. Reaction with Other Halides and Olefins. The scope of the reaction has been extended to higher olefins, substituted olefins, and the bromide system, in that 1,2-dichloropropane, 1,1,2-trichloroethane, and 1,2-dibromoethane were produced from propylene, vinyl chloride, and ethylene, respectively.
Acknowledgment
The authors thank F. Caropresso and J. Craddock, who performed the bulk of the experimentation and contributed ideas pursuant to improving the system. literature Cited (1) Benson, S. W., “Foundations of Chemical Kinetics,” Chap.
XV, McGraw-Hill, New York, 1960.
(2) Vestin, R., Acta Chem. Scand. 8, 533-57 (1954).
RECEIVED for review July 9, 1966 ACCEPTED February 3, 1967 Division of Petroleum Chemistry, 152nd Meeting, ACS, New York,
N.Y.,September 1966.
PERFORMANCE CHARACTERISTICS OF SELF- ENT RA I N M ENT EJECTO R S FOUAD K H O U R Y , l M I C H A E L H E Y M A N , 2AND W I L L I A M R E S N I C K Department of Chemical Engineering, Israel Institute of Technology, Haifa, Israel
The one-dimensional analysis of ejector performance is extended to include the conditions expected to prevail for a self-entrainment ejector operating a refrigeration cycle. Performance data were obtained for two ejectors operating with two fluids, butane and hexane. Experimental curves are presented for the entrainment ratio as a function of motive and suction fluid pressures and are compared to the performance predicted from the one-dimensional analysis. Compression ratio efficiencies in the range of 56 to 74y0 were obtained.
NE
of the applications of self-entrainment ejectors, in which
0 the motive and suction fluids are identical, is the ejector refrigeration cycle in which the ejector replaces the mechanical compressor of the conventional vapor-compression refrigeration cycle. Vacuum refrigeration systems in which a steamactuated ejector serves as the compressor for a water vapor system are well known, but the cold temperature is limited to approximately 40’ F. If lower temperatures are to be attained, it is necessary to use other refrigerants. Although some theoretical and experimental investigations of selfentrainment ejectors and of ejector-actuated refrigeration cycles have been reported (7, 3-8), few data are available on the performance characteristics of ejectors operating with identical motive and suction fluids. This paper reviews briefly the theoretical one-dimensional analysis by DeFrate and Hoerl ( 2 ) and extends it to include the conditions expected to prevail for an ejector operating a refrigeration cycle. Experimental results are presented for the performance characteristics of two ejectors and the results compared to the theoretical performance. Analysis of Ejector
section with supersonic velocity. The suction gas, 6, is entrained and mixes completely with the motive gas. In the subsequent analysis the mixing section was assumed to be constant area. The analysis was based on the following assumptions : 1. The flow is adiabatic and there are no shear forces a t the wall. 2. The gases satisfy the perfect gas equation and the specific heats are constant. 3. Acceleration of the gases from rest to the section where mixing begins is reversible. 4. Mixing is complete and the velocity is subsonic a t the entrance to the diffuser. If the velocity is supersonic after mixing, a shock occurs before the diffuser entrance. 5. Deceleration in the diffuser is reversible. 6. The motive and suction gases a t section 1 are a t the same pressure. Identical suction and motive fluids are assumed in the following review of the DeFrate and Hoerl analysis. Nozzle Equations
When a gas expands reversibly and adiabatically from stagnation pressure PO to pressure PI, it will reach a velocity that can be expressed in dimensionless form as a Mach number:
T h e analytical ejector model used by DeFrate and Hoerl is shown in Figure 1. Operation is as follows: Motive gas a a t an elevated pressure expands through the converging-diverging nozzle and emerges into the mixing Present address, Department of Chemical Engineering, Rice University, Houston, Tex. * Present address, Central Research Department, Mobil Oil Co., Princeton, N. J.
I n accordance with the assumptions made, the motive fluid will be a t sonic velocity a t the nozzle throat, so that
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