Salt removal from water by continuous ion exchange using thermal

I n d . E n g . C h e m . R e s . 1989,28, 1345-1351

1345

Panwalkar, S. S.; Iskander, E. A Survey of Scheduling Rules. Oper. Res. 1977, 25, 45-61. Pekny, J. S.; Miller, D. L. A Parallel Branch and Bound Algorithm for Solving Large Assymmetric Travelling Salesman Problem. Internal Report EDRC 05-27-88; Engineering Design Research Center, Carnegie-Mellon University: Pittsburgh, PA, May 1988. Reddi, S. S.; Ramamoorthy, C. V. On the Optimal Flowshop Sequencing Problem with No Wait in Process. Oper. Res. Q. 1972, 23, 323-331. Wismer, D. A. A Solution of the Flowshop Scheduling Problem with No Intermediate Queues. Oper. Res. 1972,20, 689-697.

Gupta, J. N. D. Optimal Flowshop Schedules with No Intermediate Storage Space. Nau. Res. Logis. Q. 1976,23, 235-243. Ku, H. M.; Karimi, I. A. Scheduling in Multistage Batch Processes with Finite Intermediate Storage. Part I I Approximate algori t h m ; Presented at the Annual Meeting of the American Institute of Chemical Engineers, Miami, FL; AIChE New York, 1986; paper 72E. Ku, H. M.; Rajagopalan, D.; Karimi, I. Scheduling in Batch Processes. Chem. Eng. Prog. 1987, 83(8), 35-45. Kuriyan, K.; Reklaitis, G. V. Scheduling Network Flowshops so as to Minimize Makespan. Comput. Chem. Eng. 1989, 23,187-200. Marsten, Roy. ZOOM User's Manual; University of Arizona: Tucson, AZ, 1986. Meeraus, A.; Brooke, T. GAMS; Development Research Department, The World Bank: Washington, DC, 1985.

Receiued for review December 12, 1988 Reuised manuscript received J u n e 6, 1989 Accepted J u n e 26, 1989

SEPARATIONS Salt Removal from Water by Continuous Ion Exchange Using Thermal Regeneration Tah-Ben HSU* and Robert L. Pigfordt Department of Chemical Engineering, University of Delaware, Newark, Delaware 19711

Thermally regenerable, bifunctional ion-exchange resins can produce potential energy savings for water purification. A cold loading column and a hot regeneration column with other essential mechanical devices constitute a continuous process in which resin beads are transported to achieve steady-state desalination. T h e solid phase and the liquid phase move countercurrently in both columns. Both a fluidized-bed mode and a descending, fixed-bed mode have been studied. T h e effects of product and water throughputs, feed pH, and regeneration temperature on the efficiency of salt removal are also investigated. Ion exchange has been an important separation technology to remove impurities from solutions for decades. The process development of ion-exchange resins depends heavily on the resin characteristics. Conventional monofunctional ion-exchange resins can exchange either cations or anions and usually need chemical regeneration with acidic, basic, or salty solutions. Amphoteric ion-exchange resins offer the advantage of simultaneous cation and anion exchange within the same exchanger. Thermal regeneration (usually below 100 "C) is superior to chemical regeneration from the viewpoints of energy conservation and pollution prevention. Bolto and co-workers (1968, 1970, 1977) pioneered the thermally regenerable processes of amphoteric resins and mixed beds with weak electrolytic, monofunctional resins. The concept of continuous operation has been attracting chemical engineers' attention for apparent reasons. Similar to operations like gas absorption and liquid extraction, most of the continuous ion-exchange processes had been developed in which the adsorbent and the adsorbate move countercurrently to achieve the advantage of less resin inventory. Along the reasoning above are the works of McCormack and Howard (1953), Hiester et al. (1954a,b),

Higgins and Roberts (1954), Gilmore (1955), Hancher and Jury (1959), Shulman et al. (1966),Kunz (1967),Bouchard (1970), Higgins and Chopra (1970), Letan (1973), Slater (1974), Gold and Sonin (19751, Gold and Todisco (1975), Kosaka et al. (1981), Savall(1982), Broughton (1984), and Khabirov et al. (1985). In this work, the amphoteric ion-exchange resins studied are Amberlite XD-2 and Amberlite XD-5 (products of Rohm and Haas Company, Philadelphia, PA). They both possess the carboxylic group and the amine group as bifunctional active sites. The resin properties of Amberlite XD-2 were studied by Ackerman et al. (1976). Knaebel et al. (1979) reported the ion-exchange rates of Amberlite XD-2 at 25 and 95 "C with step concentration change. Hsu (1979) studied the kinetics of the same resin by means of frequency analysis. Both resins had been investigated by Knaebel (1980) in a cycling zone adsorption process to remove salt from brackish water. In this paper, we study the salt removal with both resins in a truly continuous countercurrent process. The effects of operation conditions, such as throughput, feed pH, and regeneration temperature, on desalination are also presented.

*Current address: ARC0 Chemical Company, 3801 W. Chester Pike, Newtown Square, P A 19073. Deceased.

Resin Properties For the study of ion-exchange phenomena and for investigation of the process, it is necessary to understand the

0888-5885/89/2628-1345$01.50/0

0 1989 American Chemical Society

1346 Ind. Eng. Chem. Res., Vol. 28, No. 9, 1989 Table 11. Some Resin Proaerties Amberlite property unit XD-2 dP cm 0.070 PS g/mL 1.252 0.58 eP 0.33 eb Umf cm/min 2.5 4 cm/min 25

x

-

g/

~

opc

1

1

1

1

001

0

002

'

I

003

1

1

004

c , CONCENTRATION,

1

1

005

1

1

006

1

I

i

007

meq/ml

Ji-

1

Figure 1. Resin capacity of Amberlite XD-2 for NaCI.

:

5

\

AmberliteXD-2

35

10-

25

30 -I

13

j 10 "0

001

002 c

003

004

, CONCENTRATION,

005

007

006

008

meq/ml

Figure 2. Resin capacity of Amberlite XD-5 for NaC1. Table I. Langmuir Isotherma Parameters

resin Amberlite XD-2

solute NaCl

Amberlite XD-5

NaCl

aq*

= Ac/(l

temD. "C 25.8 31.7 49.5 69.0 25.9 47.4 68.8 89.7

A, mL/g drv resin 131.956 124.308 69.132 45.437 72.429 41.595 20.233 12.202

B, mL/mequiv solute 116.762 113.707 61.124 50.700 42.579 22.406 7.919 3.866

+ Bc).

resin properties, such as resin capacity, density, porosity, swelling, etc. The resin capacities of Amberlite XD-2 and Amberlite XD-5 for NaCl are shown in Figure 1 and Figure 2, respectively, a t four different temperatures. These resin capactiy data were obtained from batch equilibrium experiments with fully regenerated resin particles. They represent the maximum capacities of the resins, according to Helfferich (1962). The capacity data of these two resins show a clear Langmuir isotherm behavior. Table I lists the Langmuir isotherm parameters for both resins. In general, Amberlite XD-5 has larger resin capacity than Amberlite XD-2. Taking up salt, both resins swell. The extent of swelling is a strong function of equilibrium solution concentration but is much more weakly affected by the temperature change. Amberlite XD-2 is more porous but structurally stiffer than Amberlite XD-5 (Hsu and Pigford, 1981). Being spherical, both resin particles exhibit satisfactory mechanical strength against crushing and attrition. This

Amberlite XD-5 0.054 1.195 0.48 0.33 2.0 20

'K

Figure 3. Mass-transfer coefficients as a function of temperature. Square points were from response analysis, and spherical points were from breakthrough experiments (Hsu, 1981).

is an important property that allows each resin to be used in continuous processes (Grammont, 1970). Table I1 summarizes some resin properties which are needed later in this paper. Hsu (1981) has conducted the kinetic study for both resins to find the temperature dependence of mass-transfer coefficients, as shown in Figure 3. Since the major mass-transfer resistance lies in the resin phase, the mass-transfer coefficient is independent of the hydrodynamic conditions in the liquid phase.

Experimental System The experimental system for continuous desalination, shown in Figure 4, is described in this section. A deaeration system was applied to produce deaerated distilled water for the preparation of feed solution to minimize any potential oxidative effect on the resins. The feed solution was prepared in the cool feed tank, which was equipped with concentration, temperature, and pH controllers. The input and output pumps connected to the hot feed tank, of which only the temperature was controlled, were synchronized to keep a constant level in this tank. Both the loading column (LC) and the regeneration column (RC) were Pyrex glass tubes (60 cm long, 5 cm i.d., and 0.3 cm in wall thickness). The regeneration column was wrapped with a 300-W heating tape. This was the heat source in another controlled heating loop to prevent heat loss and to maintain the desired temperature in the regeneration column. We fed the same concentration, 0.03 mequiv of NaCl/ mL, of solution into each column from the bottom. The solution flowed upward and overflowed from the top of each column; resin beads moved downward by gravity. This achieved countercurrent operation in both columns. Two synchronized star valves controlled the resin flow rate and the height of resin bed to be 45 cm in each column. The resin beads were retrieved from the bottom of each

Ind. Eng. Chem. Res., Vol. 28, No. 9, 1989 1347

i!

2 Sample

Waste

Regeneration

Loading

Column

Column

Air

let

Q

4 Q Q

1 1 Q I

\ Roducl

I

.L

I

Figure 4. Continuous desalination process.

Figure 5. Instrumentation and steady-state monitoring system (modified from Knaebel (1980)).

column through the star valves and transported in a three-phase slug flow to the top of the other column by means of an air jet and a water jet. Through the water jet, small quantity of the waste solution and the product solution was utilized to transport the loaded resin beads and the regenerated resin beads, respectively, to their proper destinations. The resin particles adsorbed salt in the cool loading column and desorbed salt after being transported to the hot regeneration column. The steady state was monitored by a microcomputerized data-logging system, shown in Figure 5. Through a multiplexer interface circuit, a microcomputer based on the Motorola 6800 microprocessor controlled and recorded the signals from different thermocouples and conductivity cells. These probes were read by a digital microvoltmeter and a frequency counter, both equipped with BCD output. The microcomputer also gave a printout of the measured signals on a LA 36 Decwriter terminal. Both the waste and product streams flowed back to the cool feed tank to conserve NaCl in our laboratory experiments.

Results and Discussion The effects of four different operation conditions on the process performance have been investigated for both Am-

01 0

2 4 SdPERFIClAL VELOCITY I N L

6

I

8

C cm/mln Figure 6. Effect of throughput in loading column on separation factor (resin, Amberlite XD-2; U R = 3.5 cm/min; TR = 27 "C; feed pH = 5.6; R = 4.5 mL of wet resin/min; Cf = 0.03 mequiv of NaCl/mL).

berlite XD-2 and Amberlite XD-5: liquid throughput in the loading column and in the regeneration column, feed pH, and regeneration temperature. The process performance is evaluated with a separation factor. The "apparent separation factor" is defined as the ratio of waste concentration to product concentration. Since there is always some bypass of the feed solution through the star valves, the "apparent" waste or product concentration does not represent the genuine solution concentration just above the resin level in each column. By measuring the flow rate and concentration of the resin-transporting stream and the total effluent stream on the top of each column, the "true" waste or product concentration can be calculated with a simple mass balance. Hence, we can also define a "true separation factor" in a similar way. Figure 6 indicates the influence of the throughput in the loading column on the separation factor when Amberlite XD-2 is used. The hot feed flow rate is controlled such that fluidization always occurs in the regeneration column to obtain better temperature uniformity. The other operating conditions are also listed in the figure caption. As the throughput in the loading column increases, the separation factor decreases. Since the minimum fluidization velocity is 2.5 cm/min for Amberlite XD-2, only the leftmost run in Figure 6 is in fixed-bed operation. Operating the loading column in a fixed-bed mode gives better separation performance than that in the fluidized-bed mode. The back-mixing of resin beads in the fluidized-bed operation may be an important factor to cause the lower separation factor. The leakage through star valves reduces the separation factor by about 10-15%. After process characteristics, such as the resin properties and mass-transfer coefficients in Figure 3, were applied, the theory in the Appendix section gave satisfactory prediction for separation factors as shown in Figure 6 by the dashed line. One can, of course, fit the experimental separation factor data to find the best mass-transfer coefficient by means of the theory. However, in view of the pinched situation in the experiments, one does not expect to be able to determine a value of mass-transfer coefficient precisely. The close agreement may be fortuitous. Two sets of operating lines, shown in Figure 7, are also obtained from the theoretical consideration in the Appendix section. The solid lines represent the operation of the leftmost run (the fixed-bed mode) in Figure 6; the broken lines represent the run with liquid superficial velocity of 4 cm/min in the loading column (the fluidized-bed

1348 Ind. Eng. Chem. Res., Vol. 28, No. 9, 1989 I ?

t

i

c 7

! Waste

i I

00 oo

l no:

'

l 001

"

'

l

'

no6

C

'

008

.

'

l

ciio

l oi:

' VI'

meqml

Figure 7. Comparison of operating lines in fixed-bed and fluidized-bed modes. The solid lines represent the leftmost run in Figure 6, while the broken lines represent the run at uL = 4 cm/min. The regeneration operation lines of these two runs overlap (R = regeneration, L = loading, EL = equilibrium line, OL = operating line). 1

I

i

I 7

SdPERFIC,A~ $ i L O C i T Y I', F;

C

:m/m

-

Figure 9. Effect of liquid flow rate in regeneration column on true separation factor (resin, Amberlite XD-2; d, = 0.07 cm; T R = 86 "C; TL = 27 "C; uL = 1.6 cm/min; feed pH = 5.5; Cf = 0.03 mequiv of NaCl/mL; H = 45 cm; R I A = 0.078 g of dry resin/min/cm2).

i

a

a

I

31

2

L

SUPERFICIAL VELOCITY I Y

C

5 cWrr n

a

Figure 8. Effect of throughput in loading column on apparent separation factor (resin, Amberlite XD-5; uR = 2.2 cm/min; T R = 90 "C; TL= 28 "C; feed pH = 5.8; R = 4.5 mL of wet resin/min; Cf = 0.03 mequiv of NaCl/mL).

mode). The regeneration operating lines of both runs coincide with each other because of the same operating conditions. Since the slope of the operating line is the ratio of liquid flux to resin flux, the fluidized-bed run has a steeper loading operating line due to the higher liquid flow rate. Both columns are pinched a t their bottom ends. Mainly because of limitations on the capacity of star valves, higher resin flow rate could not be achieved in the conducted experiments to reduce the slope of operating lines. If this shortcoming can be improved, one may get both operating lines roughly parallel to their corresponding equilibrium lines; hence, better separation effect may be obtained. Figure 8 shows similar results for Amberlite XD-5, which seems to have poorer performance a t comparable operating conditions. From the discussion and the observation about Figure 7 , decreasing the liquid throughput in the regeneration column will reduce the slope of its operating line; therefore, the separation effect can be improved. Figures 9 and 10 confirm this for both resins. The solid lines in these two figures are again theoretical predictions with process characteristics discussed before. The channeling of liquid and solid phases may be significant even in fixed-bed operation. But it is difficult to determine just how important channeling is quantitatively,

Waste

Product

-m

U

a

I

0'

2

3

SUPERFICIAL JELOCITY

Ilk

R C

c~/rrn

Figure 10. Effect of liquid flow rate in regeneration column on true separation factor (resin, Amberlite XD-5; mixed-sized beads; T , = 90 "C; TL= 27 "C; u L = 1.0 cm/min; feed pH = 5.8; Cf = 0.03 mequiv of NaCl/mL; H = 45 cm; R I A = 0.092 g of dry resin/min/cm2).

because of the pinches in both columns. Any pinch can consume a large number of transfer units and, hence, depress the separation factor. The effect of feed p H on the separation factor can be found from Figures 11 and 12. For each resin, there is an optimal value of feed pH a t which the process has a maximum separation factor. It is about p H 5.6 for Amberlite XD-2 and p H 5.8 for Amberlite XD-5. Ackerman et al. (1976) found that Amberlite XD-2 shows a maximum resin capacity a t about pH 5.5. The data here seem very consistent with their results. Finally, the effect of the regeneration temperature on the separation factor is investigated. For both resins, the separation is improved when the temperature in the re-

Ind. Eng. Chem. Res., Vol. 28, No. 9, 1989 1349

t

m

LL

1

_I

3

7 1

i 0 5 70 80 7

O3

FEED

9

90

pH

TR

Figure 11. Effect of feed pH on separation factor (resin Amberlite XD-2; U R = 3.5 cm/min; U L = 1.5 cm/min; TR= 86 "C; TL= 27 "C; R = 4.5 mL of wet resin/min; Cf = 0.03 mequiv of NaCl/mL).

I

100

110

'c

Figure 14. Effect of regeneration temperature on separation factor (resin, Amberlite XD-5; TL = 28 "C; uR = 2.2 cm/min; uL = 1.2 cm/min; feed pH = 5.8; Cf = 0.03 mequiv of NaCl/mL; H = 45 cm; R I A = 0.092 g of dry resin/min/cm2).

R. C.

!i

/\

021

3

5

9

7

FEED

pH

Figure 12. Effect of feed pH on separation factor (resin, Amberlite XD-5; uR = 2.2 cm/min; U L = 1.2 cm/min; T R = 91 "C; TL= 28 "C; R = 4.5 mL of wet resin/min; Cf = 0.03 mequiv of NaCl/mL).

c,------J

q*dq

-

C-dC

1-- 1 - -~, 1-T-

i 1-

q

C

R

F

-f-r L-.-f

Figure 15. TCCA process scheme.

downward by gravity countercurrently against the upflowing liquid phase. The higher the regeneration temperature, the better the separation. Better separation is obtained when the liquid flow rate through the loading column is low (fixed bed) than when it is high (fluidized bed). This suggests that the process is limited by ion-exchange mass transfer. Both the bypass of feed solution through star valves and the channeling in both columns may suppress separation. The continuous process should be operated at the pH value at which the ion-exchange resin has its maximum capacity. Amberlite XD-2 seems to be better than Amberlite XD-5 in the present continuous process under the opera0I tion conditions presented in this paper.

'I 70

80

90

100

110

'C TR , Figure 13. Effect of regeneration temperature on separation factor (resin, Amberlite XD-2; TL = 27 "C; uR = 2.5 cm/min; uL = 3.8 cm/min; feed pH = 5.6; Cf = 0.03 mequiv of NaCl/mL; H = 45 cm; R I A = 0.078 g of dry resin/min/cm2).

generation column increases, as shown in Figures 13 and 14.

Conclusions A continuous desalination process has been studied with thermally regenerable, amphoteric ion-exchange resins. The solid resin phase takes up salt in a cool loading column and gives away salt after being transported into a hot regeneration column. In both columns, resin beads move

Acknowledgment The ion-exchange resins studied in this work were donated by the Rohm and Haas Company (Philadelphia, PA). This generosity and financial support from the National Science Foundation (Grant ENG 7719951) are greatly appreciated. The authors are also indebted to Prof. George C. Hsiao of the Department of Mathematical Sciences, University of Delaware, for his precious suggestions in numerical methods.

Appendix Figure 15 shows a schematic picture of a continuous countercurrent process. The adsorbent phase (ion-exchange resins, say) is fed from the top of each column,

1350 Ind. Eng. Chem. Res., Vol. 28, No. 9, 1989

retrieved at the bottom, and transported to the top of the other column. The fluid streams fed to the bases of both columns have the same composition, but that entering the regeneration column is hot while the other is cold. Since the resin releases salt at high temperatures, the resin flowing from the regeneration column can be fed to the cold loading column to adsorb salt. The following four boundary conditions are apparently true: Cc(0) = Cf CH(0) = cf qc(0) = qH(H) qH(O) = qc(H) (1) Because of the nonlinearity of the equilibrium relation, a numerical approach is presented here to simulate the process mathematically. The system of equations from the material balance has to be solved numerically with the proper constraints. Following the definitions of Ames (1977), this is a problem of mixed type which has some features of the initial condition problem and of the boundary condition problem. The first two conditions in (I),in this sense, are initial conditions, for they give the constraints when the independent variable, z , is equal to zero. With the assumption of plug flow, the following system of equations can be established in a differential height of the fixed bed:

to find CF,, and Cg,, where n = 2, 3, ..., N

+ 1:

+ ACY

cyn=

(9)

where

ACY = -(ko’ h 6 KO’

+ 2k1’+

2122’ + 123’)

= f j ( ( n - l ) h , Cyn-1, 0 )

+ hko’/2, C&-l + hk1‘/2, k3’ = fj(nh, Cyn-l + kz’h, 0)

+ h/2, k2’ = f j ( ( n- 1)h + h/2, kl’ = f j ( ( n - l ) h

Crn-l

0) 0)

and j = l as i = C

j = 2

as i = H

h = H/N (3) To solve q&, and qf?,, where n = 1, 2, ..., N + 1, we apply (1) C& and Cg,,,which are known from step 2, and (2) boundary conditions @,l

= q;,N+l

q?!,N+l = qg,l

(10)

The following system of algebraic equations is established from eq 4 and 5 for each pair of neighboring points along

(11)

We assume that R remains the same in each column and Fc equals F H . Also with the assumption that c is in equilibrium with q , the modified resin capacity is defined by q=q*+

EP C

PA1 -

tP)

Then, c can be expressed as a function of q: + 4DBq]’/2 - ( A + D - Bq) [ ( A + D c = (6) 2DB where A and B are Langmuir isotherm parameters and (7) D = tp/[(l - 4 p S l

(12)

The algorithm includes the following steps: (1) Imagine that the column is divided into N sections. There will be N 1 points (1, 2, ..., N , N 1) from the bottom to the top of the column. “m” represents the number of iterations. Then, C& represents the cold bulk liquid concentration at point n in the mth iteration. The same convention applies to q. (2) Solve eq 2 and 3 by means of the fourth-order Runge-Kutta method (Kuo, 1972; Carnahan, 1969) with (1) initial assumption

There are 2N equations from eq 11and 12. With two more boundary conditions, eq 10, it is sufficient to solve 2N + 2 unknowns (q?,, and q;,, where n = 1, 2, ..., N + 1) by a trial-and-error procedure. For example, make an initial guess for qgfl+l to calculate qgfl from the last equation of eq 12. The calculations move upward in eq 12 until qg,l is computed. Then the second boundary condition in eq 10 is applied to continue the calculation for q& in an upward sequence in eq 11until q?,l is computed. With the help of the first boundary condition in eq 10, the value of this newly calculated q& is used to replace the old guessed qgfl+l. This iteration continues until the difference between two consecutively calculated qf?fl+l’s is within a desired tolerance. (4) m = m + 1.

+

+

+

q&=O qb,n=O (2) initial conditions

n = l , 2,..., N + l

CF,l = CE,l = Cf

(8)

Ind. Eng. Chem. Res., Vol. 28, No. 9, 1989 1351 (5) Equations 2 and 3 are used to solve new C&, and CG,n by means of the fourth-order Runge-Kutta method with the initial condition, eq 8, and newly calculated qg;’ and q;;;. The same expressions as in eq 9 are applied except that proper qr;!l replaces the “0” in the functions for k i f ( j = 0-3). (6) If the absolute value of the difference between Cgfi+, and CF$+l is smaller than a designated tolerance, the computation terminates. A set of satisfactory Crn and qrn (where i = C or H, n = 1, 2, ..., N + 1) is obtained. (7) If the requirement in step 6 has not been met, then use step 3. In this algorithm, step 3 has special importance. Since two boundary conditions are included in this step to solve qrn “implicitly”, computational stability is achieved to guarantee convergence. The common “shooting method” suitable for some boundary value problems simply produces either instability or divergence and thus is inappropriate here.

Nomenclature a = mass transfer area per unit bed volume, cm2/mL A = cross-sectional area of column, cm2 c = average salt concentration in the resin phase, mequiv/mL C = concentration of solution, mequiv/mL d = average diameter, cm F = liquid throughput, mL/min H = column height, cm k o = mass-transfer coefficient, cm/s q* = resin equilibrium capacity, mequiv/g of dry resin q = modified resin capacity, mequiv/g of dry resin R = resin flow rate, mL of wet resin/min T = temperature, O C u = superficial velocity, cm/min z = distance from the bottom of column, cm Greek Symbols t

= porosity density, g/mL

p =

Subscripts b = fixed bed

C = cold e = when entrainment begins f = feed H = hot L = loading column mf = minimum fluidization p = resin particle R = regeneration column s = solid resin phase Superscripts C = cold H = hot

Registry No. Amberlite XD-2,63513-73-5; Amberlite XD-5, 76296-15-6.

Literature Cited Ackerman, G. R.; Barrett, J. H.; Bossler, J. R.; Dabby, S. S. Industrial Deionization with Amberlite XD-2: Thermally Regenerable Ion Exchange Resin. AZChE Symp. Ser. 1976, 73(166), 107-111. Ames, W. F. Numerical Methods for Partial Differential Equations; Academic Press: New York, 1977. Bolto, B. A.; Warner, R. E. Ion-Exchange Process with Thermal Regeneration. Desalination 1970, 8, 21-34. Bolto, B. A.; Weiss, D. E. The Thermal Regeneration of Ion-Exchange Resins. In Zon Exchange and Solvent Extraction (7); Marinsky, Marcus, Eds.; Marcel Dekker: New York, 1977.

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Received for reuiew J u n e 6, 1989 Accepted J u n e 8, 1989