Ind. Eng. Chem. Res. 1990,29, 1901-1907
1901
Performance of a Large-Scale Packed Liquid-Liquid Extractor A. Frank Seibert, Bob E. Reeves, and James R. Fair* Separations Research Program, The University of Texas at Austin, Austin, Texas 78712
T h e performance of commercial-size high-efficiency packings was determined by using a 42.5-cmdiameter (i.d.) liquid-liquid extractor. Both random and ordered packings were studied to determine their effects on the fluid hydraulics and mass-transfer efficiency in a larger diameter extractor. These packings included no. 25 and 40 metal Intalox saddle (IMTP) random packing and Norton 2T ordered packing. Physical property effects were studied for three test systems, Isopar-M/water, toluene/ water, and toluene/acetone/water. The influence of packed height was also investigated and was observed to have a significant effect on the mass-transfer efficiency of the random packings. The experimental results were compared with models based on data obtained in small-scale columns. For the ordered packing, these models agreed favorably with the larger scale data; such was not the case with random packings. Based on these experimental results, it is evident that significant maldistribution can occur in larger diameter towers containing random packings. T h e apparent maldistribution was observed to increase with increasing packed height and column diameter. Maldistribution did not appear to be a problem with the ordered packings. As a result of improving chemical and petrochemical markets, debottlenecking of existing processes has become attractive. In this connection, many older liquid-liquid extractors are being reevaluated to take advantage of newer, more efficient contacting devices. This has created a need for more reliable flooding and mass-transfer correlations for commercial-scale columns, since previous correlations have been verified only with data from small-diameter columns. There has always been a great deal of uncertainty whether data obtained from small packed extractors can be scaled up reliably to commercial sizes, since conventional wisdom has dictated at least an 8 1 co1umn:packing diameter ratio if channeling is to be avoided. In order to meet this requirement, smaller packing elements, possessing a large surface area:volume ratio, must be used. Under these conditions, the type of drop movement may deviate substantially from that occurring in larger columns, where a low surface area:volume packing ratio is a logical choice. Accordingly, significant differences in efficiency can arise, which may not be predicted by presently available mathematical models. The objectives of this study were to obtain flooding and mass-transfer efficiency performance data from a 42.5-cm-diameter extractor containing commercial-size packing materials and then to compare this performance with that predicted by published flooding and mass-transfer correlations. Packings for extraction are the same as those used in distillation and absorption service. These packings are classified as random or ordered. The familiar ring and saddle packings are examples of random packings. They are generally dumped into the tower, whereas ordered packings are assembled and stacked in a particular arrangement. Ordered packings (also called structured packings) are usually in the form of corrugated metal sheets that are arranged carefully for controlling the flow of the phases. A particular objective of this work was to ascertain the performance of these packings relative to that of random packings. While earlier studies have shown that high-efficiencyrandom packings perform as well as ordered packings in small-diameter columns, there was concern over whether this would be the case a t larger scales of design.
Previous Work Raschig rings and Berl saddles have been studied rather extensively under extraction conditions, but this has not 0888- 5885 / 9012629-1901$02.50 / 0
been the case for commercial-size high-efficiency random or ordered packings. Surprisingly, little has been reported on the efficiency of traditional commercial-size random packings. In the 1950s, Gayler and Pratt published an extensive group of studies of random packings in liquidliquid extraction. In 1957, they reported on the efficiency and holdup for the transfer of acetone between water and four immiscible solvents in a 10.2-cm column with 1.27-cm Raschig rings at bed heights of 61, 122, and 305 cm. Further data were obtained in 15.2- and 30.5-cm columns with Raschig rings ranging in size from 1.27 to 3.81 cm. Gayler and Pratt observed a significant effect of the direction of mass transfer, which they attributed to microoscillation of the interface for the transfer of acetone from the continuous (aqueous) phase and to enhanced droplet coalescence for the opposite direction of mass transfer. They concluded that for the case of mass transfer into a dispersed solvent phase, the ratio of the column diameter to packing diameter should be greater than eight if channeling were to be avoided. This was also verified by Leibson and Beckmann (1953). Gayler and Pratt observed a negligible effect of column diameter on the mass-transfer efficiency when comparing operations in 10.2- and 23-cm-diameter columns. However, they also observed a decrease in mass-transfer efficiency with increasing packing height, as observed later by Nemunaitis et al. (1971) and Eckert (1976). This phenomenon was attributed to end effects that provided a larger relative mass-transfer contribution with shorter packing heights. Leibson and Beckmann also studied the effect of column diameter and packing size on mass transfer in a countercurrent packed liquid-liquid extractor. They used 7.6-, 10.2-, and 15.2-cm-diametercolumns filled with 0.38-, 0.64-, 1.27-, 1.63-, 1.91-, and 2.54-cm unglazed ceramic Raschig rings. The transfer of diethylamine from a continuous water phase to a dispersed toluene phase was found to improve with increasing column diameter. Nemunaitis et al. and Eckert investigated the performance of 2.54-cm Raschig rings, ceramic Intalox saddles, and copper Pall rings in a 42.5-cm-diameter column. They observed that the column packed with saddles yielded the highest mass-transfer efficiency, followed by Pall rings and Raschig rings. The height equivalent to a theoretical stage (HETS) for the kerosene/methyl ethyl ketone/water system was found to decrease with increasing continuous-phase rates and to increase with increasing dispersed-phase flows. Nemunaitis et al. reported that 0 1990 American Chemical Society
1902 Ind. Eng. Chem. Res., Vol. 29, No. 9, 1990
condeF= cw
Hold-Up Tank
Steam
b I
Condensate
Reboiler
-
I
I
i .
Figure 1. Flow diagram of experimental equipment.
flooding of their packed column occurred at loadings only 20% of those predicted by the flooding correlation of Crawford and Wilke (1951). Recently, Seibert and Fair (1988) reported on the hydrodynamic and mass-transfer behavior of packings in a 10.2-cm extractor. Their packings included 1.27-cm ceramic Raschig rings, 1.59-cm metal Pall rings, no. 15 metal Intalox saddles (IMTP), of the random variety, and ordered packings of corrugated sheet metal and corrugated metal gauze. Mechanistic flooding and hydraulic models were presented and verified with data obtained in their small-diameter column. The present paper may be considered to be a companion to the one published in 1988 by two of the present authors.
Experimental Equipment and Procedure A flow diagram of the extraction/distillation system is shown in Figure 1. In the mass-transfer mode, distilled light-phase solvent was fed and dispersed near the bottom of the extractor into the heavy continuous phase. The operating interface was maintained at the top of the extractor by controlling the flow of the exiting heavy phase (raffinate). The exiting light-phase (extract) stream was sent to the distillation column feed tank. The solute was separated from the solvent by distillation; the purified solvent flowed to the extraction feed tank whereas the purified solute was collected in the overhead accumulator. The solute was blended with the aqueous feed to give a feed mixture in the vicinity of 6.0 wt % solute. The inside
diameter of both the extraction column and the distillation column was 42.5 cm. In the non-mass-transfer mode, the light (dispersed) phase was collected at the top of the column and was recycled to the solvent feed tank, bypassing the distillation tower. The heavy (continuous) phase was fed near the top of the extractor. The operating interface was maintained a t the top of the extractor by controlling the heavy-phase flow out of the tower. The exiting heavy phase was recycled to the heavy-phase feed tank. Three test systems were chosen for study: (1)IsoparM/water, (2) toluenelwater, and (3) toluene/acetone/ water. The first two systems were used for non-masstransfer flooding studies only. The latter system was investigated for both mass transfer and flooding. The physical properties of each of the systems are given in Table I. The equilibrium and transport data for systems 2 and 3 were reported by Misek et al. (1985). The physical properties of Isopar-M were obtained from the manufacturer. Technical-grade solvents and solutes of at least 99.2 wt % purity were used in order to minimize contamination. Compositions of all streams were determined by using a gas chromatograph with a thermal conductivity detector. Flow rates were measured with calibrated orifice meters. Experimental mass-transfer runs with material balance closures of less than 92% were discarded. With acceptable closure, the extract composition was checked to ascertain taht the equilibrium approach of the extract acetone composition to that of the entering feed was less than 90%. The experimental runs were designed so that the operating
Ind. Eng. Chem. Res., Vol. 29, No. 9, 1990 1903 Table I. Physical Properties of the Systems Studied (Temwerature = 25 "C) toluene/ Isopar- toluene/ acetone/ Miwater water watef Aqueous Phase viscocity, CP 0.89 0.89 1.05 density, g/cm3 0.994 0.994 0.989 solute diffusion coeff, cm2/s 1.01x 10-5 Organic Phase viscocity, CP 2.24 density, g/cm3 0.784 solute diffusion coeff, cmz/s slope of equilibrium lineb interfacial tension, dyn/cm 24.8
0.55 0.864 30.0
no. 25 IMTP
TAW
yes
0.53 0.853 2.51 x 10-5 0.82 22.0
nConcentration of acetone in the feed = 6 wt %. b(wt % of solute in organic phase)/(wt % of solute in aqueous phase). Table 11. Packing Characteristics specific surface packing type area, cm2/cm3 no. 25 IMTP random 2.26 no. 40 IMTP random 1.71 Norton 2T ordered 2.13
Table 111. Flooding Data packing systemn redistribution Norton TAW yes
no
no. 25 IMTP
TW
no
no. 40 IMTP
TAW
yes
no. 40 IMTP
IMW
yes
void fraction 0.95 0.97 0.97
line was approximately parallel to the equilibrium curve to maximize the accuracy of the data. The distillation column contained two beds of corrugated sheet metal structured packing, each with a height of 330 cm. This type of packing was chosen because of its high-capacity characteristics. The typical feed to the distillation column contained between 3 and 6 wt 7% acetone. The column was operated at 2-3 times the minimum reflux to ensure that pure (99.8+ wt 5%) toluene and acetone were recovered. The feed to the column was a t the center, between the packed beds. Care was taken to prevent flooding of the distillation column before the extractor capacity was reached. The extraction column contained two beds of packing with a total packed height of 640 or 700 cm. Since the packings were of type 316 stainless steel, the organic phase was dispersed while the aqueous phase preferentially wetted the packing surfaces. Normally, two beds of 350-cm height were used, with intermediate redistribution of the dispersed phase. In one set of runs with no. 25 IMTP packing, a 640-cm (23-ft) continuous bed was used to determine the effect of redistribution. The main distributor was a perforated pipe type containing 110 holes with a hole diameter of 0.32 cm and fabricated from 304 stainless steel. The light-phase distributor was located 16.5 cm below the packing. The packed beds were supported by open stainless steel grids. The redistributor was an orifice riser type containing 152 holes with a hole diameter of 0.64 cm. It was also fabricated from 304 stainless steel. Four 7.6-cm (0.d.) risers allowed redistribution of the heavy (continuous) phase. Three packings were studied, the characteristics of which are shown in Table 11. These packings are classified as "high efficiency" for gas-liquid service because of their high void fractions and their ability to provide a large amount of interfacial area for mass transfer.
Experimental Results Flooding. Flooding in packed towers occurs when true countercurrent flow no longer exists. Examples of flooding situations include entrainment of drops by the continuous phase; the dispersed phase being unable to penetrate packing, resulting in buildup of the dispersed (light) phase below the packing support; and phase inversion. None of
UCacm/s
U,,, cm/s
1.58 1.27 1.14 1.01 0.90 11.49 1.23 1.14 0.88 0.66 1.40 1.58 1.31 1.05 1.75 1.58 1.40 1.05 0.79 2.28 1.93 1.58 1.40 2.50 2.23 2.05 1.79 1.61 1.56 1.34 1.16 1.12 0.85 0.71
0.31 0.62 0.72 0.93 1.13 0.41 0.72 0.93 1.23 1.44 0.72 0.41 0.93 1.32 0.51 0.72 0.93 1.32 1.44 0.31 0.62 0.93 1.13 0.45 0.57 0.71 0.87 1.00 1.12 1.28 1.41 1.54 1.66 1.80
TAW = toluene/acetone/water system. TW = toluene/water system (no mass transfer). IMW = Isopar-M/water system (no mass transfer). 2.0
1.5
I
-
'.
L
\
1.0
-
A
4
A
0.5 -
'
0.0 0
1
I
\
A
Norton2T
.\ \
\ I
I
1
1
2
3
Superficial Flooding Velocity of the Continuous Phase cm/s Figure 2. Flooding rates for the packings tested. Toluene/acetonelwater system with toluene dispersed. Column diameter 42.5 cm, with redistributor.
these phenomena can be tolerated and will eventually upset downstream processes. Windows located on the extractor allowed visual detection of flooding. Samples were also collected to verify that two phases were exiting together in either the raffmate or the extract streams. The effects of packing size variation, redistribution, and system properties were investigated in these studies. Flooding data are given in Table 111. The effect of packing size and type on the capacity of the liquid-liquid extractor is shown in Figure 2. The no. 40 IMTP packing yielded the highest capacity, which was expected because of its lower packing area per volume. This lower area results in a less tortuous path for the phase
1904 Ind. Eng. Chem. Res., Vol. 29, No. 9, 1990 2.0 0
0)
f
I
0
c
1.5
-
-
1.0
-
-
0.5
-
1.5
-
1.0
0.5
0.0
I
0.5
1 .O
1.5
2.0
2.5
3.0
"."
0.5 1.o Superficial Flooding Vebcity of the Continuous Phase CWS
0.0
Superficial Flooding Velocity of the Continuous Phase cm/s Figure 3. Flooding rates for no. 40 IMTP packing. Continuous phase = water. Column diameter 42.5 cm; single bed for IsoparM/water, two beds for toluene/acetone/water. I
I
I
1.5
Figure 5. Flooding rates for toluene/acetone/water system. Column diameter 2.5 cm for no. 25 IMTP packing; column diameter 10.2 cm for no. 15 IMTP packing (Seibert and Fair, 1988). 6
3.0 I
i
I
I
1
A/
1
A
:1/
0.5
0.0 L 0.0
I
I
I
I
0.5
1.o
1.5
2.0
Superficial Flooding Velocity of the Continuous Phase cm/s Figure 4. Flooding rates for no. 25 IMTP packing. Toluene/acetone/water system with toluene dispersed. Column diameter 42.5 cm.
flows, which promotes a capacity approaching that of a spray column. The Norton 2T structured packing and the no. 25 IMTP packing had about the same capacity, as would be expected from their similar surface areas. The flooding rates were observed to be approximately inversely proportional to the packing surface area. Figure 3 shows the effect of the chemical system. Greater capacities were obtained with the Isopar-M/water system. This is attributed to the greater phase density difference and interfacial tension, which results in a greater drop velocity and consequently a greater capacity. The toluene/acetone/water system had the lower capacity, which may be attributed to its lower interfacial tension. The use of a redistributor caused a lower flooding capacity as shown in Figure 4. The relative reduction in capacity was constant over the range of flow rates studied. It is apparent that there is relatively little static dispersed-phase holdup in the larger packings. The lack of holdup equates to greater volume availability for flow and thus a greater capacity. As shown in Figure 5, and earlier in Figure 2, the flooding rates for the larger packings are significantly greater than for the smaller packings, which confirms the conclusion of negligible static dispersed-phase holdup. Mass Transfer. The mass-transfer efficiencies of the no. 25 IMTP, no. 40 IMTP, and the Norton 2T packings are compared in Figure 6. The supporting mass-transfer data are given in Table IV. The overall volumetric
-0.0
1 .o
0.5
1.5
Superficial Velocity of the Dispersed Phase (cm/s) Figure 6. Volumetric mass-transfer coefficients for toluene/acetone/water system with toluene dispersed. Column diameter 42.5 cm, with redistributor. Table IV. Mass-Transfer Data for the Toluene/Acetone/Water System packing Norton 2T
redistribution Yes
no. 25 IMTP
Yes
no. 25 IMTP
no
no. 40 IMTP
Yes
uc,
ud
9
cm/s
cm/s
0.21 0.33 0.49 0.66 0.90 0.21 0.33 0.49 0.66 0.82 0.99 0.21 0.32 0.42 0.53 0.58 0.21 0.33 0.49 0.66 0.82 1.05 0.32 0.42 0.53 0.63 0.21 0.33 1.49 0.66 0.90
0.26 0.41 0.62 0.82 1.13 0.26 0.41 0.62 0.82 1.03 1.23 0.41 0.62 0.82 1.03 1.13 0.26 0.41 0.62 0.82 1.06 1.32 0.62 0.82 1.03 1.24 0.26 0.41 0.62 0.82 1.13
(HTU),, cm 125 116 95 119 162 186 204 241 277 305 308 171 183 189 189 207 277 329 357 387 445 390 292 253 290 280 302 235 232 214 220
Ind. Eng. Chem. Res., Vol. 29, No. 9, 1990 1905 Table V. Key Hydraulic a n d Mass-Transfer Models Hydraulic Sauter mean drop diameter, d,
&)
112
dv, = 1 . 1 5 4
--
= 1.0, c d, no mass transfer = 1.4, d c characteristic slip velocity, U, q q
0.0
0.5 1.o 1.5 Superficial Velocity of the Dispersed Phase (cm/s)
Figure 7. Volumetric mass-transfer coefficients for toluene/acetone/water system with toluene dispersed. Packing = no. 25 IMTP. Column diameter 42.5 cm. I
-p0.149 - - 3.42H0.441- 0.857, Re
-
. 0
4 -
.
3
-Re - - 0.94H0.757- 0.857, H I59.3
I
5 -
-1000
0.009 g/(cm s)
p0.149
6
K,a(sS1)
pw =
-
0
H > 59.3
Rw Uso= 2 dV&C
-
2 -
dispersed-phase holdup = a,d,/2
(7)
1 -
0
I
slip velocity
u, = u,, exp(-1.92&) mass-transfer coefficient increases approximately linearly with increasing dispersed-phase rate until flooding is approached. The improved efficiency is a result of a linear increase in interfacial area. The ordered packings performed very well, yielding significantly greater overall volumetric mass-transfer coefficients. The reason why the random packings performed so poorly is not completely understood. One likely explanation is that there was a significant maldistribution of the phases due to column size. Further evidence supporting this reasoning may be seen in Figure 7, where a redistributor placed in the middle of the random packing bed improved the efficiency by about 40%. Maldistribution in the ordered packing appears to be absent, since the mass-transfer efficiency of the ordered packing scaled up quite well with column diameter as shown in Figure 8. The slight differences in efficiency may be due to variations in surface treatment of the ordered packings; for the smaller column, a smooth surface SMV packing (Sulzer Brothers) with 3.2 cm2/cm3 area was used, whereas the Norton 2T has a surface area of 2.1 cm2/cm3.
cos
}{ ;
+ (1 - cos
{ :})( &) (9)
continuous-phase flooding velocity, Ucf
2Ucf = E( u s 0 1 + 0.925(
$)[
cos
interfacial area a. =
‘
6cbd -
(11)
dv,
Mass Transfer continuous-phase-film mass-transfer coefficient (12) dispersed-phase-film mass-transfer coefficient
kd
=
O.O0375U,
@ 42.5 cm
(18)
Conclusions For packed liquid-liquid extraction columns, data presented here should aid the process designer in predicting flow capacity and mass-transfer efficiency of commercial-scale installations. Flow capacity tests have been run for two high void fraction random packings and one structured packing in a 42.5-cm-diameter column, and a model has been developed that permits estimation of flow capacity for systems, column sizes, and packing types other than those tested. Unlike the case for small extraction columns, where packing openings can restrict flow of dispersed-phase drops, the packing sizes appropriate for larger columns permit higher phase flow rates. In general, for different packing sizes, the capacity appears to be inversely proportional to the specific surface area of the packing. On this basis, the ordered, or structured packing, had about the same capacity as the random packing. Mass-transfer rates for the packings tested were found to fit models developed earlier on the basis of smaller scale tests. There was a slight effect of column diameter on efficiency, with a constant factor needed to account for the reduced efficiency at larger column sizes. Thus, unlike the case for flooding, the diameter did appear to have an influence on mass transfer. It was found, however, that for the random packings there is a height effect on masstransfer efficiency, with deterioration for longer packed beds correctable through the use of dispersed-phase redistributors. This problem of maldistribution did not appear to be present when structured packings were used. Acknowledgment This work was supported by the Separations Research Program at The University of Texas at Austin.
Nomenclature ai = interfacial area, cm2/cm3 a p = total packing area per volume of column, cm2/cm3 cf = correction factor for flooding (eq 19) C = concentration, g/cm3 C * = equilibrium concentration, g/cm3 d, = column diameter, cm d,, = Sauter mean drop diameter, cm D = diffusion coefficient, cmz/s g = gravitational constant, cm/s2 HTU, = overall height of a transfer unit based on the continuous phase, cm k , = individual film coefficient for the continuous phase, cm/s hd = individual film coefficient for the dispersed phase, cm/s K , = overall mass-transfer coefficient based on the continuous phase, cm/s mdc = distribution coefficient based on a concentration driving force, dCd*/dC, Re, = drop Reynolds number based on the continuous phase, P c Usdvsl M c Sc, = Schmidt number based on the continuous phase, p c /
(PP,)
(19)
Scd = Schmidt number based on the dispersed phase, pd/
(20)
Sh, = Sherwood number based on the continuous phase,
No correction was necessary for the mass-transfer efficiency prediction for the ordered packings. The resulting efficiency models were found to predict the experimental
TPdDd)
kcdvslDc
U = superficial velocity, cm/s Vi, = interstitial velocity of the continuous phase, cm/s U s = slip velocity, cm/s
Ind. Eng. Chem. Res. 1990,29, 1907-1914
U , = characteristic slip velocity, slip velocity at low dispersed-phase flow rate, cm/s Greek Symbols Ap = density difference, g/cm3 e = void fraction of packing { = dimensionless tortuosity factor, a,d,,/2 p = liquid viscosity, g/(cm s) 7 = drop size correction factor p = liquid density, g/cm3 u = interfacial tension, dyn/cm & = fraction of dispersed-phase holdup in the contacting
section 3 = criterion for determining the applicability of either the
Handlos and Baron or the Laddha and Degaleesan dispersed-phase-film mass-transfer coefficient model Subscripts c = continuous phase d = dispersed phase w = water
Literature Cited Crawford, J. W.; Wilke, C. R. Limiting Flows in Packed Extraction Columns. Chem. Eng. Prog. 1951,47(8), 423. Eckert, J. S.Extraction Variables Defined. Hydrocarbon Process. 1976,55 (31,117.
1907
Gayler, R.; Pratt, H. R. C. Liquid-Liquid Extraction V. Further Studies of Droplet Behaviour in Packed Columns. Trans. Inst. Chem. Eng. 1953,31,69. Gayler, R.; Pratt, H. R. C. Liquid-Liquid Extraction X. Overall Mass Transfer Coefficients on an Area Basis for Extraction of Acetone in Packed Columns. Trans. Inst. Chem. Eng. 1957,35, 273. Grace, J. R.; Wairegi, T.; Nguyen, T. H. Shapes and Velocities of Single Drops and Bubbles Moving Freely Through Immiscible Liquids. Trans. Inst. Chem. Eng. 1976,54,167. Handlos, A. E.; Baron, T.Mass and Heat Transfer from Drops in Liquid-Liquid Extraction. AZCh J. 1957,3,127. Laddha, G. S.;Degaleesan, T. E. Transport Phenomena in Liquid Extraction; McGraw-Hill: New York, 1978. Leibson, I.; Beckmann, R. B. The Effect of Packing Size and Column Diameter on Mass Transfer in Liquid-Liquid Extraction. Chem. Eng. Prog. 1953,49 (S), 405. Misek, T.; Berger, R.; Schroter, J. Standard Test Systems for Liquid-Liquid Extractions; Institution of Chemical Engineers: Rugby, England, 1985. Nemunaitis, R. R.; Eckert, J. S.; Foote, E. H.; Rollison, L. H. Packed Liquid-Liquid Extractors. Chem. Eng. Prog. 1971,67 (ll),60. Seibert, A. F.; Fair, J. R. Hydrodynamics and Mass Transfer in Spray and Packed Liquid-Liquid Extraction Columns. Znd. Eng. Chem. Res. 1988,27,470.
Received for review January 8, 1990 Revised manuscript received May 29, 1990 Accepted J u n e 9, 1990
Extraction of Mercury(I1) with Sulfurized Jojoba Oil Jaime Wisniak,*it Gal Schorr,t Dov Zacovsky,+and Sofia Belferl Department of Chemical Engineering and The Institutes for Applied Research, Ben-Gurion University of the Negev, Beer-Sheva, Israel
Sulfurized jojoba oil containing 12% by weight S has been tested as an extractant for Hg(1I) from aqueous solutions. Experiments have been performed with the extractant dissolved in a solvent (liquid-liquid extraction) or adsorbed in an appropriate resin matrix (solid-liquid extraction). The extraction characteristics of both systems have been measured and show that sulfurized jojoba oil exhibits very good possibilities as an extractant. T h e performance of several resins treated with sulfurized jojoba oil for adsorbing mercury(I1) was studied. T h e morphology of the different resins was examined by using scanning electron microscopy. The sulfurized oil is attached to the resin sites through the sulfur atoms; it is estimated that there are about 2 mol of S active sites per kilogram of resin.
Introduction Mercury metal and most of its compounds are poisons that can be fatal to all living organisms. In the context of present-day governmental regulations, the control, recovery, and disposal of mercury-bearing wastes are as important as the manufacturing process. EPA regulations normally require that liquid effluents contain no more than 5 ppb. The most significant source of pollution is the mercury-containing brines from chloralkali industries. Cleaning procedures include the use of activated carbon impregnated with silver, contact of the brine solution with a strong anion-exchange organic resin of the quaternary ammonium cross-linked type, precipitation of mercury salts from alkaline solutions with soluble sulfides or water-soluble reducing agents, etc. Solvent extraction and separation techniques using liquid membranes are considered to be the most effective and energy-saving separation techniques for the purpose mentioned above (Kirk-Othmer, 1981). Swanson et al. (1973) have suggested that treating solutions containing Hg(I1) ion with starch xanthate-polycation complex can reduce the residual mercury content to extremely low levels (3.8 ppm)
without introducing large amounts of other contaminants. In the past, solvent extraction of mercury(I1) has been investigated using extractants like trioctylphosphine oxide, tributyl phosphate, trioctylamine, and tricaprylmethylammonium chloride (Aliquat 336). Some of these extractants contain nitrogen or oxygen as a donor atom and have only poor selectivity to mercury(I1)over other metals; much effort is being spent today on the possible use of extractants containing sulfur as a donor atom, since some such as dialkyl sulfides and trialkyl thiophosphate are known to be very powerful and selective extractants for mercury( 11). The medium-to-low selectivity of neutral oxygen-containing extractants of metal complexes with inorganic ligands is due to the fact that these extractants may extract the metal both in the form of coordinately solvated compounds and in the form of complex metal acids. The selectivity of sulfur-containing neutral extractants must be higher. They are usually protonized with difficulty; therefore, their solutions in inert diluents do not extract complex metal acids. On the other hand, these extractants give coordinately solvated neutral complexes only with
0888-5885/9012 629- l907$02.50/0 0 1990 American Chemical Society
