Ind. Eng. Chem. Res. 1987,26, 726-731
726
Knopf, F. C.; Okos, M. R.; Reklaitis, G. V. Znd. Eng. Chem. Process Des. Dev. 1982, 21(1), 79-86. Murtagh, B. A.; Saunders, M. A. Technical Report SOL 80-14, June 1980, Stanford University Systems Optimization Laboratory, Palo Alto, CA. Schrage, L. User's Manual for LZNDO; The Scientific: Palo Alto, CA, 1981. Sparrow, R. E.; Forder, G . J.; Rippin, D. W. T. Znd. Eng. Chem. Process Des. Dev. 1975, 14(3), 197-203. Vaselenak, J. A. Ph.D. Dissertation, Carnegie-Mellon University, Pittsburgh, PA, 1985.
Literature Cited Avriel, M. Nonlinear Programming: Analysis and Methods; Prentice Hall: Englewoods Clifs, NJ, 1976. Balas, E. Lin. Alg. Its Appl. 1971, 4 , 341-352. Duran, M. A,; Grossmann, I. E. Technical Report DRC 06-68-84, 1984; Design Research Center, Carnegie-Mellon University, Pittsburgh, PA. Duran, M. A,; Grossmann, I. E. AZChE J . 1986, 32, 592-606. Garfinkel, R. S.; Nemhauser, G. L. Integer Programming; Wiley: New York, 197'2. Geoffrion, A. M. J . Opt. Theory Appl. 1972, 10, 237-260. Greenberg, H. J.; Pierskalla, W. P. Opns. Res. 1971,19, 1553-1570. Grossmann, I. E.; Sargent, R. W. H. Znd. Eng. Chem. Process Des. Dec. 1979, 18(2), 343-348.
Received for review December 2, 1985 Revised manuscript received September 29, 1986 Accepted December 6, 1986
Performance of a Continuous Solid-Liquid Two-Impinging-Streams (TIS) Reactor: Dissolution of Solids, Hydrodynamics, Mean Residence Time, and Holdup of the Particles Abraham Tamir* and Moshe Grinholtz D e p a r t m e n t of Chemical Engineering, B e n Gurion University of the Negev, Beer Sheva, Israel
T h e two-impinging-streams (TIS) reactor, already employed for many gas-solid and gas-liquid operations, was successfully implemented here for liquid-solid systems. The following performance parameters of the reactor were investigated: mean residence time and holdup of the particles as well as the dissolution of urea in water, while both phases are in a continuous flow. It was found that the T I S reactor is a very effective device that provides the highest mass-transfer coefficients among the reactors designed for continuous operation. Dissolution of solids is a very important operation in chemical engineering. It may be carried out in a batch device such as the agitated vessel (Hixon and Wilkens, 1933; Hixon and Baum, 1941; Johnson and Huang, 1956; Barker and Treybal, 1960) and the rotating dissolution cell (Bennet and Lewis, 1958) and in a semibatch apparatus such as a packed bed reactor (Ranz, 1952) or a trickle bed reactor (Hirose et al., 1976; Satterfield et al., 1978). Since the impinging-streamsreactor, applied so far for gas-solids and gas-liquids systems, has proven to be a very efficient device for effecting various processes in chemical engineering (Elperin, 1972; Tamir et al., 1984; Luzzatto et al., 1984; Tamir and Hershkovitz, 1985; Tamir, 1986; Tamir and Luzzatto, 1985; Tamir and Luzzatto, 1985a,b; Tamir and Sobhi, 1985, Tamir et al., 1985), an attempt has been made to apply this reactor also for solid-liquid systems. Therefore, the major aims of the present work were (a) to develop and to test the two-impinging-streamsreactor for dissolving solids in liquids, while both phases are in a steady continous flow; (b) to study the hydrodynamical behavior of the reactor and to find out the mechanical energy needed to transfer the solid-liquid suspension through the reactor; (c) to measure important properties needed for reactor design, such as the mean residence time of the particles in the reactor and their holdup; and (d) to measure mass-transfer coefficients in the dissolution process of urea in water in order to be able to evaluate the effectiveness of the reactor as compared to other commonly used devices. The above aims were completely achieved, indicating that the TIS reactor is a very useful tool for dissolving solids in liquids. ~~
* To whom correspondence should be addressed. 0888-5885/87/2626-0726$01.50/0
Liquid-Solid TIS Reactor Properties The reactor is shown schematically in Figure 1, and it comprises the following main elements (1)two inlet pipes, 9, for the solid-liquid suspension; (2) reactor, 4, with an inner pipe, 5, which directs the streams toward the active zone for the dissolution process where the streams collide; (3) particles feeding system, 6, which separates the flow into two streams. The particles enter the shown hopper from another hopper, whose exit diameter could be varied to obtain the desired particles flow rates; (4) conical exit, 7. The device operates as follows: two streams of a solidliquid suspension are fed tangentially into the annular space between the inner pipe 5 and the external cylinder 4. Consequently, the two streams are allowed to impinge at a predetermined location in the upper portion of the reactor. Oscillation of the solid particles is obtained at impinging zone 11,at high flow velocities of the suspension in the inlet pipes, due to the centrifugal and inertia forces acting on the particles. Thus, a particle penetrating into the opposed water stream may acquire a relative velocity U = Up- (4,) = Up+ U,. If under extreme conditions Up = U, at the entrance of the reactor (point 2 in Figure I), and assuming that this velocity is maintained up to the impingement zone, then U = ZU,. Under practical conditions, the above value of U is probably not achieved. However, the relative velocity is increased with respect to Up to such an extent that a significant reduction in the external resistance to the dissolution process occurs as compared to a configuration where countercurrent flow between particles and water does not exist. Thus, an enhancement in the dissolution mass-transfer rate may be expected. Another advantage of the configuration of impinging streams is the possible increase in the mean 0 1987 American Chemical Society
Ind. Eng. Chem. Res., Vol. 26, No. 4, 1987 727 In the present hydrodynamical experiments, the operating conditions were such that kmax= 0.5 and pp = 1200; thus, Pm Pw (4)
PARTICLES
where the difference (pm - p w ) / p w < 1 % . Typical values of the terms appearing in eq 1 are (in the order from left to right) 0.40, -0.14, and 0.215 m. If 4 is considered, the pressure drop on the reactor is given by the relation (5) pf= f b w , vwt u w , Wp, Ww, g, dp, Di) where Di designates the various geometrical lengths of the reactor shown in Figure 1. Applying the Buckingham Pi theorem yields Eu, = f ( R e , P , Fr, d p / D r ,Di/Dr) (6) where Dr is a selected reference length which makes the various lengths nondimensional. In the absence of particle flow, namely P = 0, eq 6 gives Euw = f ( R e , Fr, D i / D r ) (7) In eq 6 and 7, Eu, and Euw are the Euler numbers in the presence and in the absence of particles flow, respectively. The Euler number is defined by
The Reynolds number given by
D UW Re = VW
Figure 1. Two-impinging-streams reactor for dissolution of solids.
residence time of the particles due to their oscillation. In addition, good mixing between the opposed liquid streams is obtained and hence equalization in the concentration of the dissolved material. Finally, due to the particular flow pattern in the reactor and the relatively small holdup of the particles (in the order of 2.5% of the reactor’s volume), the effective area for the transport processes is the actual surface area of the particles. This is not the case in a fixed or fluidized bed reactor, where the active area is less than the surface area of the particles because of their contact.
Hydrodynamics of the Reactor The hydrodynamics of the reactor was determined by measuring the pressure drop between points 1 and 3 (Figure 1)as a function of the water and the particles flow rates as well as the type of the particles used (urea with a density of 1330 kg/m3 and a mean diameter of 1.7 X m and acetal with a density of 1200 kg/m3 and a diameter of 2.1 x m). In contrast with the gas-solid suspensions, where the measured pressure drop (P1- P3) is totally attributed to friction forces, in the present case it is obtained from Bernouli’s equation which reads f l f
The bar designates scaled quantities, and the relevant physical properties of the suspension are maintained constant due to the condition p = 1. Neglecting the effect of gravity because of the relative high inertia of the particles (i.e., Fr = 0) and considering eq 6 and 7 yields - Eu, = Euw =1 (12) Then, from relation 10 it follows that
v
= qe/qs = 1
(13)
u+l
w,/ww
It should be emphasized that the above relation unites the data in Figure 2 into a single equation. The application of the above information for scale-up, namely, for calculating Upfor a large-scale reactor, is performed as follows:
+ PmAZ, - 2 3 )
(1)
In Figure 1, 1 and 3 designate positions. The density of the solid-liquid suspension is given by
where =
The conditions for a complete hydrodynamical similitude between a large-scale reactor (subscript e) and a small-laboratory-scale reactor (subscript s) are = Re,/Re, = 1 I; = L e / L , = constant (11) P = PdP* = 1
The significance of eq 13 is that a correlation of the kind 7 = f ( p ) , determined from small-scale measurements, could be used to predict the pressure drop with particles flow in a large-scale reactor. It should be noted that relation 13 was verified by Tamir and Shalmon (1987). Figure 2 demonstrates the behavior of Eu vs. Re where the q-p relation for acetal and urea particles in water reads q = 1.008 + 1 . 2 6 8 ~ (0 Ip 5 0.06) (14)
= (PI - P3) + P m ( u w , l - uw,3)/2
P
is defined on the basis of the inlet diameter to the reactor, D = 0.01 m (Figure 1). On the basis of Eu, and Eu,, the following quantity is defined: EuAP-
(3)
728 Ind. Eng. Chem. Res., Vol. 26, No. 4, 1987
Wp
x IO2
00.0 o
062' I 0 0 Acetol
B I50 I
[kg/sl
Woter
i
1200
0
Acetal
1330 A b & U r e o B e s t guess i i n e
I3 U r e a
086 X
15
1% 104
15~104
R e =-
z X 304
D"W
VW
Figure 2. Hydrodynamics for TIS reactor for acetal and urea particles.
-
4-
Y
-
0
\
c,
>
W
a
w 2 w
0 LL -
3-
2-
-
0
w
a
v,
1-
0.I
0.2 W,
( Kg/
0.3
s)
Figure 3. Mechanical energy per unit feed vs. water flow rate.
For a determined value of p , the value 7 = AP /e, is obtained from eq 14 for the large-scale reactor. $he value of AP, for the large-scale reactor may be obtained from the Eu-Re curve for the small-scale reactor, in the absence of particle flow. In this case, the value of AP, for the large-scale reactor will be the maximal (Tamir and Shalmon, 1987), and therefore a safe value for design purposes. Another possibility is to measure AP, for the large-scale reactor, namely, in the absence of particles flow, and to use these data to obtain AP, from 7APw where 9 is obtained solely from small-scale reactor experiments, as verified by Tamir and Shalmon (1987). Another important parameter for design of the reactor is the specific mechanical energy ( E ) defined by
where hpf is in units of N/m2, Q, in m3/s, and W, and Ww in kg/s. The dependence of E on the water flow rate is shown in Figure 3 where it may be concluded that the effect of the particles flow on E is relatively small as compared to the energy needed to transfer only water through the reactor. This behavior is acceptable consid-
Ind. Eng. Chem. Res., Vol. 26, No. 4,1987 729 Table I. Range of Operating Conditions 0.66 X 10-*-1.95 X lo-' particles flow rate, W,, kg/s 1.77 X 10-4-2.95 X lo4 water flow rate, W,, m3/s water velocity in inlet pipe, m/s urea concn a t inlet,O C,, kg of urea/m3 of water urea concn a t outlet from the reactor, Co, kg of urea/m3 of water saturation solubility of urea in water, C*, kg of urea/m3 of water particles holdup, V, kg of urea surface area of spherical particles, A , m2 reactor volt (including inlet pipes), m3 h
wp
a02
A
A
I 13 z';ri2 k p l s
A
d
I 96 x I O - ~
01
Figure 5. Mean residence time of particles vs. w.
\
0-00 195
- Best
guess line
\
li-
s t
i
X
10-'-5
X
lo-'
1.174 at 25 "C 0.01-0.028 0.028-0.075 7.94 X
"The reactor comprised the inlet pipes designated by 12 in Figure 1.
greater by a factor of 7 than that of the gas phase, in comparison to a factor of about ir! the present liquidsolid system. However, the most important observation is the existence of a relative velocity between the phases which reduce the external resistance to the dissolution mass transfer.
IJ= W p / W w
t
1.2
X
i
a05
05
1.2-1.9 0.15 X 10-3-3
.
I
Dissolution of Urea in Water The two-impinging-streamsreactor was also applied for investigating the process of dissolution of solids. Water was introduced through the inlet pipes, and spherical particles of urea, with mean diameter d, = 1.7 X m and density pp = 1335 kg/m3, were fed from conical hopper 6 (Figure 1)into the two water streams. It should be noted that part of the water flow rate ( ~ 5 %enters ) through valve 8. This has been done in order to obtain a smooth flow of the particles into the reactor. The performance of the reactor was evaluated in terms of the mass-transfer coefficient defined by
I
1 ' ' ' 11300)
uRE-=-4 a 0 = D =
0
0
0.05
p i wp/
0.I
w,
Figure 6. Residence time ratio against mass flow rate ratio of particles to water.
between the particles when the water flow rate is increased, and hence their chance to leave the reactor is increased. The dependence of the mean residence time (7)is demonstrated in Figure 5 vs. the flow rate ratio ( p ) . The various kinds of particles have different behaviors, and presently, there is no consistent explanation for this. The only conclusion drawn from Figure 5 is that at p = constant, 7 a l / p , because of the increase of the buoyancy force effect when pp C p,. Figure 6 exhibits the mean residence time ratio rp/7, vs. the mass flow rate ratio W,/ W,. The prominent fact demonstrated is that for urea rP/7, C 0.5. Thus, solidliquid slip velocities are observed, confirming the expectations of high relative velocitiies between the phases and enhancement in the dissolution mass-transfer process as later discussed. For polypropylene an opposite behavior is observed, namely 7p/7, > 1,reaching 1.5, which means that the particles reside more time in the reactor. It may also be concluded that if pp/pw > 1,the solid particles will reside less time in the reactor than the continuous phase (water). This behavior is opposite to that of gas-solid systems where the mean residence time of the particles is
Equation 17 assumes plug flow. Indeed, the system is composed of the inlet pipe (between points 10 and 2 in Figure 1) and TIS reactor 4. Recent tests involving the pulse input of urea (in the vicinity of point 2 in Figure 1) into the reactor indicated that it behaves as a well-mixed reactor. Consequently, the application of AC,, in eq 17 is justified. Under the assumptions that ail the particles are spherical and with the same diameter and that the particles' surface area is the true area for mass transfer, 6V A=(18) PPdP
with V being the holdup of particles. Since the saturation concentration of urea in water (C*) is much higher than Co and Ci (approximately by a factor of 200), the mean logarithmic driving force of the process AC,, can safely be assumed to be equal to C* and eq 17 becomes
Because C* >> C,, eq 19 derived for a plug-flow reactor fits also the case of a perfectly mixed reactor. The ranges of the operating conditions in the experiments on dissolution of particles of urea in the reactor with impinging streams are presented in Table I. The concentrations of the dissolved urea at the inlet and outlet of the reactor were determined by taking samples with a syringe from positions 10 and 3. The analysis of urea concentration was made in a fully automatic ASTRA system at the chemistry laboratory of the Soroka Hospital in Beer Sheva.
730 Ind. Eng. Chem. Res., Vol. 26, No. 4, 1987 Tab1 11. Comparison of Mass-Transfer Coefficients for Processes of Dissolution exptl apparatus two-impinging-streams reactor (present work) trickle-bed reactord rotating dissolution cell”
chemical system urea-water benzoic acid-water benzoic acid-water benzoic acid-4% sucrose sol. lead-mercury tin-mercury benzoic acid, salicyclic acid-water salicylic acid-benzene succinic acid-n-butanol and acetone boric acid-water benzoic acid-water benzoic acid-benzene benzoic acid-methanol benzoic acid-ethylene glycol barium chloride-water sodium chloride-water naphthalene-methanol
agitated vesselb agitated vesselC agitated vessel‘
“Bennet and Lewis, 1958. Johnson and Huang, 1956. CBarkerand Treybal, 1960. ’Mean value of D at mean operating conditions with respect to temperature.
1 o 9 ~ m/s ,l 1.2 1.15 1.43 0.24 1.15 1.55
104k,, m/s 0.9-1.6 0.023-1.2 0.0081-0.46 0.048-0.22 0.17-0.75 0.088-0.75
10-~(k,/o), 7.5-13.3 0.20-10.4 0.057-3.2 2.O-9.2 1.5-6.5 0.57-4.8
0.28-1.85 1.29 1.15 1.40 1.40 3.5 X lo-* 1.18 1.38 1.75
0.036-0.88 0.092-2.3 0.15-0.55 0.42-0.58 0.32-0.63 0.013-0.019 0.54-0.61 0.54-0.87 0.63
1.3-4.8 0.71-17.8 1.3-4.8 3.0-4.1 2.3-4.5 3.7-5.4 4.7-5.2 3.9-6.3 3.6
Satterfield et al., 1978. e Hixon and Baum, 1941.
two-phase system of the kind considered here, the following correlation is applicable:
Wp(kg/si
00066
o 00113 0 00195
k is the number of groups corresponding to the geometry of the TIS reactor and a, cy, 6 , y, 6, and ei are adjustable parameters. For the present dissolution process, the above correlation is reduced to
0
al
l 09i 0 I5
’
0 ‘
1
‘
1 32
1
1
1
‘
1
0 25
1
’
I
1
! 03
WJ kg/s )
Figure 7. Mass-transfer coefficients vs. flow rate.
The values of the mass-transfer coefficient calculated from expression 19 are reported in Table I1 and are also compared with available data. An additional quantity appearing in Table I1 is k , / D which is proportional to the ratio of the convective mass-transfer rate divided by the diffusive rate. Consequently, k , / D may also serve for comparison purposes among various kinds of devices for dissolving solids in liquids. The major conclusion drawn from the comparison of the values of the transfer coefficients presented in Table I1 is that the two-impinging-streams reactor provides the highest mass-transfer coefficient among the chemical reactors designed for continuous operation. The relatively high values of k , may be attributed to the unique hydrodynamic behavior of the solid-liquid suspension which was explained before. Only the values of the mass-transfer coefficient obtained from the batch agitated vessels reported in Barker and Treyball (1960) are higher than those of reactor with two impinging streams. Thus, the twoimpinging-streams reactor can be efficiently implemented for dissolving solids in liquids. Figure 7 demonstrates the variation of the mean masstransfer coefficient vs. the particles and the water flow rates whose extreme values are reported in Table 11. Although an increase is observed in k , for the highest particles flow rate, the trend, rather, is not clear considering the standard deviation in k , demonstrated for two data points. However, the efficient performance of the continuous reactor in dissolution of urea is unequivocally verified in comparison to other devices. For scale-up purposes, it is necessary to express the mass-transfer data by means of nondimensional groups. Using the Buckingham Pi theorem yields that for the
Sh = ap“Rep (21) On the basis of the data appearing in Figure 7 , the following correlation was obtained: Sh = 6,0”J.056Re0.45 (22) The mean deviation between the experimental data and the data calculated by means of eq 22 is 9.7’70,the minimal deviation is 3.570,and the maximal deviation is 21.5%. These deviations are within the accuracy needed for design purposes. Acknowledgment We thank R. Rifches, in charge of the Chemistry Laboratory of the Soroka Hospital in Beer Sheva, for his kind assistance in the analysis of urea in water. Nomenclature A = surface area of a particle given by eq 18 C* = saturation solubility of urea in water Co = concentration of urea at the reactor exit (position 3 in Figure 1) Ci = concentration of urea at the reactor inlet (position 10 in Figure 1) d, = particle diameter D = diffusion coefficient of urea in water D, = various geometrical lengths of the reactor shown in Figure 1
D, = reference length E = specific mechanical energy needed to transfer the sus-
pension through the reactor, defined in eq 15 Eu = Euler number defined in eq 8 Eu , E , = Euler number in the presence of particle flow and for water only, respectively = EuJEu,, a scaled quantity which is equal to the ratio of the Euler number for a large-scale reactor to the Euler number for a small-scale reactor Fr = Froude number, F I L g g = gravity acceleration h, = mass-transfer coefficient defined in eq 17 and 19 L = geometrical length
Ind. Eng. Chem. Res. 1987, 26, 731-737
L = scaled quantity, L,/L,
P,-P3 = measured pressure drop between positions 1and 3 in Figure 1 Q, = water volumetric flow rate Re = Reynolds number defined in eq 9 Re = scaled quantity which equals Re,/Re, S c = Schmidt number Sh = Sherwood number, k,d f D u = relative velocity, u = + U, U, = water velocity at the inlet pipe to the reador (9 in Figure 1) Up = particle velocity V = reactor holdup including the inlet pipes in Figure 1 W = mass flow rate of the particles w“, = mass flow rate of water Z1-Z3 = difference in height between positions 1 and 3 in Figure 1 Greek Symbols AC1, = mean logarithmic driving force APf = pressure drop on the reactor between positions 1 and 3 in Figure 1 due to friction which is calculated by eq 1 Up, AP,,,7 pressure drop on the reactor between positions 1 and 3 in Figure 1 with particle flow and in the presence of particle flow, respectively 7 = defined in eq 10 q = scaled quantity equal to ve/vs p = defined in eq 3 p = scaled quantity equal to p e / p s pm, pp, pw = density of the mixture defined in eq 2, density of particles, and density of water, respectively vw = kinematic viscosity of water T , T , T , = mean residence time defined in eq 16, mean resi8ence time of the particles, and mean residence time of water, respectively
731
Subscripts e = large-scale reactor m = of the mixture p = of the particle s = small-scale reactor w = of water
Literature Cited Barker, J. J.; Treybal, R. E. AIChE J. 1960,6, 289. Bennet, J. A. R.; Lewis, J. B. AIChE J . 1958,4 , 418. Elperin, I. T. Nauka Tekh. (Leningrad) 1972, 1. Grinholtz, M. “A Two-Impining-Streams Reactor for Solid-Liquid System”, Research Report, 1985. Hirose, T.; Mori, Y.; Sato, Y. J . Chem. Eng. Jpn. 1976,9, 220. Hixon, A. W.; Baum, S. J. Ind. Eng. Chem. 1941,33, 478. Hixon, A. W.; Wilkens, G. A. Ind. Eng. Chem. 1933,25,1196. Johnson, A. J.; Huang, C. J. AIChE J. 1956,2, 412. Luzzatto, K.; Tamir, A.; Elperin, I. T. AIChE J . 1984,30, 600. Ranz, W. E. Chem. Eng. Process. 1952,48, 247. Satterfield, C. N.; Van Eek, M. W.; Bliss, G. S. AIChE J . 1978,24, 709. Tamir, A. Chem. Eng. Sci. 1986,in press. Tamir, A.; Elperin, I.; Luzzatto, K. Chem. Eng. Sci. 1984,39, 139. Tamir, A,; Hirshkovitz, D. Chem. Eng. Sci. 1985,40, 2149. Tamir, A.; Luzzatto, K. AIChE J. 1985a,31, 781. Tamir, A.; Luzzatto, K. J. Powder Bulk Solids Technol. 1985b,9,15. Tamir, A.;Luzzatto, K.; Sartana, D.; Surin, S. AIChE J. 1985,31, 1744. Tamir, A.; Shalmon, B., submitted for publication in Ind. Eng. Chem. Res. 1987. Tamir, A.; Sobhi, S. AIChE J . 1985,31, 2089. Received for review December 20, 1985 Revised manuscript received September 22, 1986 Accepted December 6, 1986
Phase Equilibria in Supercritical Propane Systems for Separation of Continuous Oil Mixtures Maciej Radosz Exxon Research a n d Engineering Company, Annandale, New Jersey 08801
Ronald L. Cotterman’ and John M. Prausnitz* Materials a n d Molecular Research Division, Lawrence Berkeley Laboratory, and Chemical Engineering Department, University of California, Berkeley, California 94720
Experimental and calculated phase equilibria are reported for two systems containing propane and petroleum-derived oil mixtures near 400 K and pressures to 55 bar; these conditions are close to critical. In the first system, the oil is rich in saturated hydrocarbons; in the second system, the oil is rich in aromatic hydrocarbons. For both oils, number-average molecular weights are in the range 300-350. Solubilities in two equilibrium phases, measured with a flow-cell apparatus, are correlated with the perturbed-hard-chain equation of state wherein the composition of the heavy hydrocarbon is described by a continuous distribution function. A simple procedure is proposed to derive a molecular-weight distribution from boiling-point data. Calculated and experimental equilibria agree well when small empirical corrections are introduced into the perturbed-hard-chain equation of state to obtain characteristic potential-energy parameters for oil-propane interactions. For design and evaluation of supercritical-fluid-extraction processes, it is necessary to have quantitative phase-behavior data. We report here a study of vaporliquid equilibria for two propane-oil mixtures at temperatures near 400 K and pressures to 55 bar, near the mixtures’ critical conditions. The first mixture contains an Current address: W. R. Grace, 7379 Route 32, Columbia, MD 21044. 0888-5885/87/2626-0731$01.50/0
oil rich in saturated hydrocarbons; the second mixture contains an oil rich in aromatic hydrocarbons. Both oils have number-average molecular weights in the range 300-350. These phase equilibria are of interest because for many years, low-molecular-weightparaffin solvents (typically, propane) have been used to extract valuable components from heavy crude residua. In many cases, such a process is operated as a liquid-liquid extraction. More recently, 0 1987 American Chemical Society
