dimensionless distance along the system 0 = dimensionless time 8, = residence time of the exit element O,i = the computed residence time a t time io, Os = sampling period 7 =
Nomenclature
A = magnitude of step change in inlet state A , = area of cross-section of the reactor C A = concentration of component A C A ~= initial steady-state inlet concentration A e = error, difference between set point and output of process go, g1 = discrete proportional-integral controller parameters kA = rate constant for the reaction A B K = controller gain K* = feedback gain setting a t which oscillations in the output occurs L = length of the system mi = manipulated variable value between the i t h and (i 1)st sampling instants m(0) = is the manipulated variable defined as C/o n = order of the reaction A B t = time u ( 0 ) = unit-step function, u(e) = 1 , O 2 0; u(O) = 0,O < 0 u = velocity through the system ui = velocity through the system during the time between i t h and (i 1)st sampling instants X = dimensionless state variable, equal to C A / C A ~ X L = lower limit of the output X a t the sampling instant Xu = upper limit of the output X a t the sampling instant X l ( 0 ) = inlet state of the process X*(O) = outlet state of the process X, = set point value of the loop z = distance along the reactor/exchanger
-
+
-
+
Subscript m
= refers to ultimate steady-state value
Others - = refers to initial steady state L i t e r a t u r e Cited COX, J. B.. Hellums, L. J., Williams, T. J., Banks, R. S., Kirk, G. R., Jr., /SA J.,
13,65 (1966). Dahlin, E. B.. Instrum. ControlSyst., 41(6),77 (1968). Hassan, M. A.. Solberg, K. O., Automatica, 6, 409 (1970). Koppei. L. B.. lnd. Eng. Cbem., fundam., 5, 403 (1966). Koppel. L. B., Kamman, D. T., Woodward, J. L., Ind. Eng. Cbem., fundam., 9,
198 (1970). Moore, C. F.. Smith, C. L., Murill, P. W , Instrum. Control Syst., 43(1),70
(1970). Mosler, H. A., Koppel, L. B., Coughanowr, D. R., Ind. fng. Cbem., Process Des. Dev., 5, 297 (1966). Mosier, H. A.. Koppel, L. B., Coughanowr, D. R., AlCbEJ., 13(4),768 (1967). Mutharasan, R.. PhD. Thesis, Drexel University, 1973. Palas. R. F., Ph.D. Thesis, University of Minnesota, 1970. Paraskos, J. A,. McAvoy, T. J., AlChEJ., 16(5), 754 (1970). .Se&.!&l. J H.. Lapidus, L., Cbem. Eng. Sci., 23, 1461 (1968). Seinfeld, J. H.. Gavalas, G. R., Hwang, M., lnd. Eng. Chem., Fundam., 9, 651
(1970). Vermeychuk, J. G..Lapidus, L., AlChEJ., 19(1),123 (1973).
Greek Letters fl = process parameter, defined as LkAc).i"-l/fi
Receiued for reuiew March 17,1975 Accepted July 30, 1975
Coal Liquefaction in Coiled Tube Reactors R. E. Wood' and W. H. Wiser Department of Mining, Metallurgicaland Fuels Engineering, University of Utah, Salt Lake City, Utah 84 112
Dry, powdered coal impregnated or mixed with 5 % ZnCI, can be converted to liquid and gaseous products in small diameter coiled tubes. Yields to 70% (50% liquid, 20% primarily CH4 gas) can be obtained at 5OO0C and 1800 psi hydrogen pressures in 3/16-in. i.d. stainless steel tubes 60 to 120 ft in length. The liquid portion is a complex mixture, principally aromatic, which could serve as a synthetic crude petroleum for refinery feed stock. Procedures for catalyst recovery and or recycle are indicated.
Coal Liquefaction in Coiled T u b e Reactors Since coal represents the great majority of the fossil fuel reserves of the United States it is expected to carry a significant portion of the burden of synthetic natural gas and synthetic petroleum requirements beginning before the 1980's (Hottle, 1971). Many problems relating to mining, preparation, and conversion to liquids and gases remain to be solved. Still, the needs are such that authorization has been given for construction and operation of several large pilot plants for production of synthetic natural gas from coal and lignite. Three of these plants, those of Consolidation Coal Co. (COz-acceptor process), Bituminous Coal Research Corp. (Bi-Gas process), and Institute of Gas Technology (Hi-Gas process), have been supported in large measure by the Office of Coal Research of the Department of the Interior. A fourth process is that of the U S . Bureau of 144
Ind. Eng. Chem., Process Des. Dev., Vol. 15,No. 1, 1976
Mines (Synthane process). The only coal liquefaction process currently funded for large pilot plant testing is the multi-stage coal pyrolysis process of FMC Corp. (COED process), This process also has had considerable Office of Coal Research support. Selection of candidate processes for large scale testing of coal hydrogenation to produce synthetic liquid fuels is yet to be made. One of the coal hydrogenation projects supported by the Office of Coal Research is that a t the University of Utah. This process has been through several bench scale phases, all of which have been directed toward a short contact time and high ratio of hydrogen to coal under moderate reaction conditions (Haddadin, 1968; Qader, 1969). The current reactor configuration (block diagram shown in Figure 1) is a series of coiled ?&-in. i.d. X ?&in, 0.d. no. 316 stainless steel tubes. The length of this tube system can
Table I. Comparison of Catalysts Tested in the Small Diameter Tube Reactop Salt
L I Figure 1. Block diagram of coiled tube coal hydrogenation reactor system.
be varied from 20 to 120 ft. The coal, ground and screened 100 mesh particle size, is entrained in a stream of to -40 hydrogen gas and is carried into and through the small tube by the force of the moving gas. Temperatures of 500 to 550°C and pressures of 1600 to 2000 psi hydrogen are sufficient for the reaction. The coal is impregnated with zinc chloride as the catalyst by drying a slurry mixture of the two and is metered into the reaction tube through the use of a variable speed star wheel device a t the bottom of the coal hopper. A second application procedure consists of tumbling dry coal and dry ZnClz together to obtain a uniform mixture. An impregnated catalyst is not essential. From the hopper, the coal passes through a 20 f t long preheater to bring it to the desired reaction temperature. The coal is then held at the reaction temperature for the desired time (1 to 6 sec) by the choice of tube length. Conversion of coal matter to liquid and gaseous products is measured by collecting and weighing the solid (char) and liquid fractions. Gases are measured by the weight difference between initial and final product weights. A normal trial consists of passing coal a t the rate of 100-200 g/min and hydrogen a t the rate of 30 scfm (standard cubic feet per minute), about 60 g/min, through the reactor. The coal charge amounts to 1200 g of dry coal with about 5.5% of dry zinc chloride. The solids fraction is the char and unreacted coal from the solids receiver after correction for toluene soluble liquids, ash, and catalyst. The liquids are considered to be the total materials carried into the condenser collecting vessels plus the toluene soluble liquids from the char. Gaseous conversion is obtained by difference. Total conversion is defined as the weight of MAF (moisture and ash free) coal minus the corrected char divided by the weight of MAF coal treated in the reactor. In the most rigorous sense the liquids should be corrected for benzene insolubles. However, the data on benzene insolubles in the liquids was obtained in only a few cases. A general survey of small diameter tubes as coal hydrogenation reactors was made with: (1)a 3-ft section of Ih-in. i.d. tubing, (2) a 3-ft section of l/4-in. i.d. tubing, and (3) a 6-ft section of %-in. i.d. tubing. Although powdered coal can be converted to liquids and gases in these short tubes the conditions are quite severe. Temperatures of 650 to 700°C are required to achieve 60-65% conversion. Some reactor plugging must be expected at these temperatures. A decrease in the fraction of coal converted follows (1) an increase in coal feed rate, (2) a decrease in average particle size, or (3) an increase in hydrogen flow rate. A higher coal feed rate means a lower hydrogen to coal ratio or less hydrogen availability for reaction. A smaller average particle size means that more of the coal is carried through with the gas and is not in the hot zone long enough to become heated to reaction temperature. An increase in hydrogen flow rate, Le., linear flow velocity, also means a reduced residence time in the reaction zone. Measurements of residence
+
ZnBr, ZnI, ZnC1, SnC1,-2H,O SnC14*5H,O LiI CrC1,
% Conversion 58.5 46.3 41.1 40.5 25.6 16.6 12.8 11.7 11.0
Salt Sn (powder) CUC1,*2H,O FeC 1,- 6H,O Zn (powder) ZnSO,. 7H,O ("4) sM@2,'4H@ FeC1, C aC 1., H,O N+C03.H20
%
Conversion 7.9
7.6 7.2 7 .O
5.4 5.4 3.3 No reaction No reaction
7.9 a Comparison Conditions. Last Chance, Utah coal, -40 + 100 mesh particle size; reactor tube: Ys-in. i.d. x 40 in. length; temperature: 650°C, pressure: 1750 psi; hydrogen flow rate: 3.5 scfm (standard cubic feet per minute); hydrogen gas velocity: 18.1 ft/sec; coal feed rate: 12.5 g/min; catalyst concentration: 10% by weight; catalyst application: impregnation from solution except for Sn and Zn metals in which cases the dry coal and dry powdered metal were mixed.
times within the hot zone of the '/&in. i.d. reactor tube indicated the range of 2 to 6 sec for the coal. Calculation of gas residence time is of the order of 0.8 sec for the same situation. These data indicate that coal sticks to the hot tube wall to some extent or it would pass through at the same rate as the gas. However, the adherence of the coal is not such as to prevent a continuing flow of coal solids and liquids. A survey of catalyst materials, Table I, permitted a choice of zinc chloride as the most economical of the effective catalysts. Figure 2, obtained with tests from the I/d-in. i.d. reactor tube, shows the typical response to changes in pressure and temperature. An increase in either parameter provides an increase in conversion to liquids and gases to pressures of 2000 psi. At higher pressures the liquid conversion continues to increase, somewhat, but the conversion to gases is decreased. The use of longer reactor tubes has resulted in extension of the residence time. The most obvious consequence of the use of long small diameter tubes has been that lower operating temperatures are effective. Figure 3 shows the effect of increased reactor length on the conversion of coal samples at two different temperatures in a Yls-in. i.d. reactor tube. Progressing from 20 f t to 120 f t of reactor length the conversion increases from 20 to 85% a t 538°C and from 10 to 63% at 510°C. These measurements were made at coal feed rates of about 16 lb/hr. Higher temperatures cause reactor plugging and lower temperatures result in greatly reduced conversion. Reactor operation is very smooth and dependable in the 100-ft length at 538°C or the 120-ft length a t 51OOC. Conversions of 60 to 70% of the coal matter can be expected in this operating range. These data were obtained by immersing coiled 20-ft sections of tubing into a molten lead bath and connecting them together to provide the various lengths. Figure 4 shows the effect of increased coal feed rate on the fraction of coal reacted a t 510°C in the 10 to 30 lb/hr range. Measurements were made with a 100-ft, a 110-ft, and a 120-ft 3/ls-in. i.d. coil as the reactor. All three cases show a marked decrease in coal conversion with increased feed rate. The effect of increased feed rate can also be seen in Figure 5 where conversion (data of Figure 4 from a 3/ls-in. i.d. reactor of 100, 110, and 120 f t length) is plotted in terms of space rate measured in pounds of coal processed Ind. Eng. Chem., Process Des. Dev., Vol. 15, No. 1, 1976
145
0 0
650
675
0
20
IO
30
700
TEMPERATURE, ‘C
Figure 2. Pressure and temperature effects for the ‘h-in. i.d. tube reactor.
i
c90
-60
’1
u
I
400
eo0
1200
1600
COAL FEED R A T E , L B / F T 3 / H R
Figure 5. Coal feed rate vs. coal conversion in terms of space rate utilization in the “coiled tube” reactor. REACTOR LENGTH, F E E T
Figure 3. Effect of reactor length in the “coiled tube” system. per hour per cubic foot of reactor volume. This figure indicates that a feed rate of 800 to 900 lb/ft3 hr will give 60% conversion. A feed rate of 500 lb/ft3 hr will yield 70% conversion. To show the real value of this process these space rate values should be compared with published rates of 30 lb/ft3 hr for the H-Coal process (Project H-Coal), and 20 to 60 lb/ft3 hr for the Bureau of Mines Synthoil process (Akhtar, 1971). The seemingly high space rates reported are a result of the turbulent motion caused by pumping large volumes of hydrogen through a small tube to push coal and catalyst through the hot zone. This motion makes for intimate contact between coal and hydrogen in the presence of a catalyst with a minimum of liquid present to act as a diffusion barrier. Whether this concept can be made to function in larger tubes, i.e., 1-in. i.d. as would be required for a Process Development Unit, remains to be proven. Tests which were terminated prior to completion of the expected feed period and before the flow of hydrogen could sweep the coal out of the tube indicate something of how the process operates. The coal is found in the form of an annular layer on the walls of the tube. The center of the tube is free for passage of gas. The hydrogen and product gas mixture moves down the center of the tube, impinging on and moving the fluid coal and coal product mass with it. Heat for the reaction is obtained through the wall of the tube from the molten lead heat transfer bath. Since coal hydrogenation reactions are mildly exothermic in character it is to be expected that heat flow from the coal layer through the tube to the lead would also occur. Plugging of the reactor occurs a t conditions where the flow of gas is low; i.e., the mass is not kept in motion sufficiently to prevent excessive devolitilization of the char with the resulting formation of a dry solid that plugs the flow. A second plug condition is found at high temperatures where the same excessive devolitilization can occur. In general, any set of conditions that will lead to conversions in excess 146
Ind. Eng. Chem., Process Des. Dev., Vol. 15,No. 1, 1976
of 75 or 80% of the coal matter can be expected to lead to reactor tube plugging. The desired goal of this project has been the production of a synthetic crude oil. This means an emphasis on production of liquid products with a minimum of gases. One measure of success then is the ratio of liquids to gases. Figure 6 (data from 3/16-in. i.d. reactor tubes) shows this measure as a function of the space rate factor and we see that the 800-900 lb/ft3 hr feed rate produces a ratio of 6 times as much liquids as gases. In terms of coal conversion, Figure 7 shows that at 60 to 65% conversion we can expect from 5 to 7 times as much liquids as gases. The 60 to 70% conversion range is of interest because it is expected that 30 to 40% of the coal will be needed for plant heat requirements and for production of process hydrogen. The data of Figure 7 were obtained in YI6-in. i.d. reactor tubes.
Quality of Products Gases. The recycle hydrogen gas used during the course of a coal hydrogenation test shows a continuous buildup of hydrocarbon gases as well as H2S, “3, and H2O. Table I1 shows a gas chromatographic analysis of samples taken at intervals during a 1-hr test where 15 lb of coal was tested. Hydrogen is the major component as would be expected. In the hydrocarbon portion of this gas at the conclusion of the test, methane predominates at 60%, followed by ethane at 20%. Propane is present at about 10% and normal butane is present at about 3%. Isomers and unsaturates show 1%or less each. The Cs hydrocarbons are found in trace quantities only. H& “3, and H20 make up the balance of the gas not shown in Table 11. Tests with hydrogen diluted with nitrogen indicate that inert gases may be present up to 50% of the recycle gas without a decrease in the coal conversion. Liquids. The liquid products from the coiled tube reactor are collected in two fractions. The material from the first condenser is a heavy, viscous product while that from
Table 11. BuilduD of Hvdrocarbons in Recvcle Gas during a 1-hr Coal Hvdrocenation Trial Sample no. Sample taken
1 Pre-run
3 20 min
4 30 min
5 40 min
6 50 min
7
a
10 min
60 min
Post-run
95.7 0.65 0.38 0.32 0.27
94.2 1.29 0.59 0.43 0.30
94.4 1.30 0.63 0.45 0.31
91.7 2.24 0.90 0.57 0.33
87.7
87.8
2
~~
% H2 % CH, % C2H6 % C3H8 % N-CJHi,
97.8 0.28 0.25 0.26 0.25
3.15 1.10 0.68 0.35
87.8
3.02 1.13 0.65 0.34
3.06 1.11 0.65 0.34
Table 111. Chemical and Physical Analysis of Oil Products Filtered Oil Samples
B
A
29.5 37 1.ova 42.1 SFS (58.80~) 129,179 0.54 0.17 2.2 74.7 7.5 (15.2)
Flash point “C Pour point “C ~ p g. r . g/cm3 Viscosity Btu/gal Ash, YG %S
%N %C %“cH
c/c
0 (by difference)
C
34.5 35 1.070 56.8 SFS (500~) 133,305 0.61 0.20 1.6 70.3 7.5 (20.6)
27.O 31.5 1.073 70.53 SFS (50°C) 136,550 0.37 0.17 1.54 74.25 7.35 (16.3)
Whole sample (Heavy Liquid)
22.3 16.6 43.1 15.9 2.1 __ 100.o
As phalte ne s Residue Soluble Oils Water Loss
15.7 24.2 36.2 ia .7 5.2 100 .o
12.62 14.27 43.32 26.33 3.46 100.0
59 .a 34.2 5.5 0.5 100.o
57.3 22.4 13.7 6.6 100.o
Soluble Oils
87.6
Neutral Oil T a r acids T a r bases Loss
11.1
1.3 0 -
L o , d
~
100 .o
I
I
I
COAL
FEED
RATE,
LB/FTS/HR
W
Figure 6. Ratio weight liquiddweight gases in terms of space rate utilization in the “coiled tube” reactor. C O A L CONVERSION, P E R C E N T
Figure 7. Correlation of weight liquiddweight gases with percentage of coal converted in the “coiled tube” system. the second and third condensers (combined) is a light, low viscosity material. The heavy liquid fraction is separated first by removal of water, then filtered through a 100-mesh screen to separate char and converted coal. The oils are washed from the char with benzene and the benzene soluble material washed with cyclohexane after removal of benzene. The result is a residue, an asphaltene (cyclohexane insoluble) and an oil fraction. The liquid portion is also treated with benzene, filtered, benzene removed by evaporation to produce a residue or benzene insoluble fraction
and soluble oils. The soluble oils from both char and liquid portions are combined and then separated into tar acid, tar base, and residual or neutral oils by the use of 10% soluHzS04. The complete analysis of the tions of NaOH three samples treated in this fashion, together with physical and elemental analysis on the same samples is given in Table 111. The unusually low flash point of these samples is
+
Ind. Eng. Chem., Process Des. Dev., Vol. 15, No. 1, 1976
147
Table IV. Proximate and Ultimate Analvses of Starting Coal and Some Selected Chars Proximate analysis (as is basis) Sample
% Moisture
% Ash
Volatile matter
Fixed carbon
Btu/lb
Coal (Hiawatha, Utah) Char 1263 85.6% Conversion Char 1 2 7 1 61.4% Conversion Char 1272 59.67, Conversion
2.43
6.70
44.24
46.63
11,450
0.34
34.40
15.86
49.40
11,020
0.48
17.63
29.90
51.99
12,050
0.24
17.54
29.95
52.27
12,200
Ultimate analysis (dry, ash free basis) Sample
% Carbon
% Hydrogen
% Nitrogen
% Sulfur
Coal Char 1263 Char 1 2 7 1 Char 1272
70.1 63.3 70.1 69.6
5.4 4.1 5.6 5.8
1.6 1.7 1.6 1.5
0.88 3.53 1.42 1.31
brought about because condensable pentanes and hexanes are included. These materials are easily volatilized and ignited. The liquids show about 20% water, 20% benzene insoluble residue, 15% asphaltenes, and 40% cyclohexane soluble oils. The soluble oils show 60% or more neutral oils, 20% tar bases. Some unconverted coal or char is carried over as indicated by the 0.5% ash content. Low sulfur content in the liquids indicates that the sulfur remains in the char (inorganic sulfur) or is converted to H2S (organic sulfur). Char. The coal used in the tests discussed herein was obtained from the U S . Fuel Co. Hiawatha, Utah Mine a t Hiawatha, Utah. Table IV gives an analysis of the coal and three char products resulting from different tests. This table shows that a t 60% conversion the char still contains 30%volatile matter and 16% volatile matter a t 85% conversion. Carbon, hydrogen and nitrogen contents are almost unchanged from the original coal, while sulfur is greatly increased in the char. These observations tend to bear out the premise that the coal particles are reacted a t the outer surface only and that the remaining char is largely unreacted coal. This idea is further supported by the fact that the char can be recycled through the reactor, without further addition of catalyst and more conversion to liquids and gases will take place. The char, following an initial pass through the reactor, has a reduced softening point from that of the coal. The char is quite fluid below 200°C while the coal does not enter the plastic zone until about 350°C is reached. However, the char is readily devolatilized and the softening temperature increases rapidly when heat is applied. The chars, as collected in the solids receiver a t the end of the reactor tube, have not been exposed to solvents. Some of the volatile matter present is due to low boiling condensed products that collect with the solid residue. These products also contribute to the low melting character of the char. Catalyst Recovery. The role of ZnCl2 in coal hydrogenation in small tube reactors is not fully understood although some progress has been made (Bodily, 1974). The ZnClz melts a t 283°C and boils at 732°C a t atmospheric pressure. Under the reaction condition of 500-550°C it would be a liquid with a fairly high vapor pressure. I t coats the coal surface readily and apparently has high mobility. 148
Ind. Eng. Chern., Process Des. Dev., Vol. 15, No. 1, 1976
Table V. Recovery of ZnClz Catalyst from Reaction Products Coal charge Products
1 0 7 0 g Coal, 60 g ash, 70 g ZnC1,
a. 428 g Char, 60 g ash, 34.4 g ZnC1, b . 550 g coal liquids, 35.6 g ZnC1, c. 9 2 g g a s Liquid Treatment Separation of water from oil in liquid product. d . 465 g oil, 3.3 g Zn, 8.7 g C1, e . 85 g water, 23.6 g ZnC1, First water rinse of oil (d). f . oil, 1.4 g Zn, 4.7 g C1, g. water, 5.9 g ZnC1, Second water rinse of oil (d). h. oil, 0.19 g Zn, 1.16 g C1, i . water, 4.75 g ZnC1, Char Treatment One hour leach with cold concentrated HNO, on Char (a). j . Char, 0.43 g Zn One hour leach with cold concentrated HNO, plus 1 hr leach with boiling "0,. k. Char, 0.14 g Zn One hour leach with cold 6 N HNO, plus two 1 hr leach treatments with boiling 6 A' "0,: 1. Char, 0.39 g Zn
I t seems to move freely to form or to assist in the formation of reaction sites. At an application rate of 5.5% ZnCln and a cost of $0.40 per pound the catalyst cost is $44.00 per ton of coal. This means that the catalyst is twice as valuable as the coal used in the process. If we can attain 95% recovery of the zinc and can produce 2.5 barrels of synthetic crude petroleum per ton of coal the direct catalyst cost would be reduced to about $0.90 per barrel. This is still a high and undesirable expense. If a 99% recovery could be realized the direct catalyst cost would be reduced to $0.20 per barrel, a figure that
is much more acceptable. This catalyst cost represents only the cost due to inability to recover the zinc and does not represent the costs involved in various recovery schemes. Following the hydrogenation step the zinc is distributed about equally between the liquid and solid products. In the liquid, the zinc is primarily, but not exclusively, in the water phase. The zinc content of the organic phase is greatly reduced by water washing. The zinc associated with the char is much more difficult to recover. Some of the char associated zinc is water soluble and even more is soluble in hydrochloric acid. Still more is rendered soluble by the use of cold concentrated nitric acid or hot 6 N nitric acid. Table V lists the overall recovery of the zinc catalyst by the application of two water rinse steps on the liquid portion and three different char treatment procedures for zinc removal. The first char treatment proceduae; combined with the water rinse treatment of the liquid, results in a 98.2% recovery of the zinc. The second char procedure represents a 99.0% recovery o f t h e zinc, and the third a 98.3% recovery of the zinc. More stringent acid treatment procedures would result in more complete catalyst recovery, but a t the expense of greater char oxidation. Because of the small scale of the operation, it has not been possible to make a realistic evaluation of the cost factors in catalyst recovery as yet.
Literature Cited Akhtar, S..Feedman. S., Yavorsky. P. M.. Technical Progress Report No. 35. Pittsburgh Energy Research Center, Pittsburgh, Pa., Bureau of Mines Coal Desulfurization Program, 1971. Bodily, D. M., Lee, S.,Wiser, W. H., 167th National Meeting of the American Chemical Society, Los Angeles, Calif., Mar 1974. Haddadin, R. A,, Anderson, L. L.. Hill, G. R.. Technical Report, Department of the Interior, Office of Coal Research, Contract Number 14-01-0001-271. University of Utah, Salt Lake City, Utah, "Dilute Phase Hydrogenation of High Volatile Bituminous Coal", 1968. Hottle, H. C., Howard, J. B., "New Energy Technology, Some Facts and Assessments", p 103, MIT Press, Cambridge, Mass.. 1971. Hsieh, B. C., Wood, R. E., Anderson, L. L., Hill, G. R.. Anal. Chem., 41, 1066 (1969). "Project H-Coal, Report on Process Development", Research and Development Report No. 26, Department of the interior, Office of Coal Research, Contract Number 14-01-0001-477. Hydrocarbon Research, Inc., Trenton, N.J., 1967. Qader, S. A., Hill, G. R., Hydrocarbon Process., 48, 141 (Mar 1969). Qader, S. A,, Haddadin, R. A,, Anderson, L. L., Hill, G. R.. Hydrocarbon Process., 48, 147 (Sept 1969).
Received for reuiew March 24, 1975 Accepted July 2,1975
This work has been supported by a grant from the Office of Coal Research, U.S. Department of the Interior (now Energy Research and Development Administration) and the State of Utah.
Free Energy Parameters for Reverse Osmosis Separations of Undissociated Polar Organic Solutes in Dilute Aqueous Solutions Takeshi Matsuura, J. M. Dickson,' and S. Sourirajan' Division of Chemistry, National Research Council of Canada, Ottawa, Canada, K 1A OR9
The concept of free energy parameter (-AAG/RT) governing reverse osmosis separations is extended to include nonionized polar organic solutes in aqueous solutions. Data on (-AAG/RT) are derived for many solutes with respect to both cellulose acetate and aromatic polyamide membrane materials. A general equation for solute transport parameter is given. For a given membrane material-solvent system, the above equation includes four quantities representing respectively (i) the average pore size on the membrane surface, (ii) the free energy parameter for the solute, (iii) the steric parameter for the solute, and (iv) a parameter which cancels the overlapping effect of pore size on the above three quantities. The above equation enables the prediction of reverse osmosis separations for a large number of organic solutes from reverse osmosis data for the system sodium chloride-water only. The prediction technique is extensively illustrated in the operating pressure range 250 to 1000 psig.
Introduction The concept of free energy parameter governing reverse osmosis separations of ionic solutes in aqueous solutions has been discussed (Matsuura et al., 1975b; Dickson et al., 1975). Using a modified form of the Born expression for free energy of ion-solvent interaction to both the bulk solution phase and the membrane-solution interface where water is preferentially sorbed, a parameter has been ob-
Co-op student, Department of Chemical Engineering, University of Waterloo, Waterloo, Ont., Canada.
tained to express the free energy change (AAG) involved in the repulsion of the ion at the interface. This parameter, called the free energy parameter for the ion, is represented by the symbol ( - A A G / R T ) i where R is the universal gas constant, T is the absolute temperature, and the subscript i represents the particular ion under consideration. It has been shown that in the case of the completely ionized inorganic solutes in aqueous solutions discussed earlier (Matsuura et al., 1975b), the solute transport parameter ( D A M / K ~ ) is related to (-AAG/RT)i for the ions involved by the expression In ( D A M / K = ~ )In C*
+ r(-AAG/RT)i
Ind. Eng. Chem., Process Des. Dev., Vol. 15, No. 1, 1976
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