Ind. Eng. Chem. Process Des. Dev. 1904, 23, 557-561
557
(hexadecanes and similar boiling range alkanes) has been developed. When coupled with hydroglycolysis of waste water-blown foams, the process upgrades the waste to a very high quality polyether triol. Economic evaluation of this Ford process indicates that it will be attractive to waste foam generators that cannot find a market for their waste.
back time of 1.8 years after taxes, and a time adjusted rate of return of 47% after taxes. The two shift per day operation, however, assumes that 1800000 kg per year of waste foam are available a t no cost to the recycler. Obviously, the operating profit of the plant is highly dependent on the ratio of the value of the poly01 produced to the cost of the plant. This ratio has not varied very much during the past few years. This is also the reason why two shift per day operation is so much more profitable. In practice, large generators of uncontaminated polyurethane foam waste find a market for this material among carpet underlay manufacturers, packaging material users, and others. Mid-1982 price of uncontaminated waste flexible polyurethane foam varied between $0.22 and $0.44/kg. Thus, an in-house recycler would give up the income resulting from waste foam sales in order to produce polyol. When the value of the waste foam at $0.33/kg is considered as an operating cost for the process, the first year results given in Table V become a before tax loss of $154 000 for a one shift operation and a before tax profit of only $31OOO for the two shift per day operation. Clearly, there is very little incentive to recover poly01 at the price at which clean waste foam can be sold today. These results show the real value of the hydrocarbon extraction purification process. Contaminated waste foam does not command a significant market price. Indeed, disposal in a landfill (at a cost) is the usual fate of these wastes. As a result, it is likely that generators of dirty, contaminated polyurethane waste will find the recovery process described in this paper attractive.
Acknowledgment
The authors are indebted to W. E. Stevens, J. D. Lisius, J. E. Albright, and L. M. Briggs for their extensive experimental assistance. The help of D. E. Wilemski in evaluating the recovered polyols at the Utica Trim Plant and in reviewing the economic evaluation of the process is also gratefully acknowledged. T. W. Jaskolski’s help in the preparation of process cash flow analyses is also appreciated. Registry No. Diethylene glycol, 111-46-6. L i t e r a t u r e Cited Braslaw, J.; Pai, P. U S . Patent 4 159972,July 3, 1979. Gerlock, J. L.; Braslaw, J.; Mahoney, L. R.; Ferris, F. C. J . Povm. Scl. 1980. 78, 541. Geriock, J. L.; Braslaw, J.; Zinbo, M. Ind. Eng. Chem. Process Des. Dev. 1884, preceding paper in this issue. Helss, H. L. U.S. Patent 3 109824,Nov 5, 1963. Kinoshita, 0.U. S. Patent 3632530,Jan 4, 1972. Kinstle, J. F.; Sepulveda, L. E. J . Po!)” Scl. 1977, 75(8), 467. McElroy, W. R. U.S. Patent 3 300 417,Jan 24, 1967. TenBroeck, T. R.; Peabcdy, D. W. US. Patent 2937151, May 17, 1960. Tucker, B.; Ulrich. H. US. Patent 3 983 087,Sept 28, 1976. Ulrich, H.; Odinak, A.; Tucker, 8.; Sayigh, A. A. R. Po/ym. Eng. Scl. 1978,
78(1 l),844.
Conclusion
Received for review September 13,1982 Revised manuscript received August 29, 1983 Accepted October 14,1983
A simple and relatively inexpensive poly01 purification process involving extraction with high boiling hydrocarbons
Interpretation of the Leaching Kinetics of Pentlandite in a Complex System by the Shrinking Core Model Daniel A. D. Boateng and Colin R. Phllllps’ Department of Chemlcal Engineering and Applied Chemlstty, University of Toronto, Toronto, Ontario, M5S 7A4, Canada
Nickel was leached from pentlandite concentrates in a complex system in which an organic phase containing bis(2-ethylhexyl)phosphoric acid was added to the aqueous slurry, with oxygen u* pressure forming the gaseous phase. The shrinking core model was used to interpret the data, and the leaching rate was shown to be controlled by both chemical reaction and mass transfer processes. The relative importance of the ratecontrolling mechanisms varied over the duration of the leaching and to some extent with temperature. No direct dependence of the rate of leaching on the concentration of nickel in the solid phase was found. However, the rate was approximately first order with respect to oxygen pressure for the 377 to 791 kPa pressure range.
consisting of an aqueous slurry, an organic phase, and an oxygen gas phase, was investigated. The organic phase consisted of bis(Zethylhexy1) phosphoric acid (DEHPA) solution in kerosene. Details of the process are described elsewhere (Boateng, 1979). Many models for describing leaching processes exist in the literature (Loveday, 1975; Madsen et al., 1975; Roach and Prosser, 1978; Letowski, 1980). The shrinking core model described by Yagi and Kunii (19551, Levenspiel
Introduction
Although the leaching of sulfide concentrates under oxidizing conditions in acid media has been widely studied (Forward and Vletman, 1959; Vizsolyi et al., 1967; Prater et al., 1970; Bjorlin, 1973), this leaching system has found only limited application in the hydrometallurgy of nickel sulfide comentrates, mainly because of lack of selectivity. In a search for a leaching system which would allow for easy isolation of the solubilized nickel, a multiphase system 0196-430518411 123-0557$Q1.50/0
0
1984 American Chemical Society
558
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 3, 1984
(1972), and Szekely and Themelis (1971) was used to treat the data obtained in the work reported herein. Basic Considerations Particles of the concentrates treated were assumed to have negligible or very small initial porosity. Oxygen, which acts as the oxidant, was available through the maintenance of an oxygen pressure within the reaction autoclave. The solubility of oxygen in the liquid phase is a function of temperature and pressure. The leaching process may be thought to proceed according to the steps: (i) transfer of oxygen from the bulk liquid phase across the film surrounding the solid particle to the particle surface; (ii) transport of oxygen through the product layer to the surface of the unreacted core; (iii) reaction of exposed reactive species at the reaction surface; (iv) transport of leached metal values through the product layer to the exterior surface of the solid; (v) transport of the leached metal values through the interfacial film into the main body of the fluid; and (vi) distribution (diffusion and reaction) of the metal values between aqueous and organic phases. Oxygen plays an oxidizing role through the reaction O2 4H+ + 4e 2H20 (1)
-
+
-
The oxidation of metal sulfide occurs by MS M2++ So 2e
+
(2)
Possible further oxidation of elemental sulfur could occur by 2S0 + 3H20 S2032- + 6H' + 4e (3)
-
and
S2032-+ 5H20
+
2SOd2- 10H'
+ 8e
(4)
However, as discussed elsewhere (Boateng and Phillips, 1978), reaction 3 does not occur readily in acidic media. If iron is solubilized, oxidation of iron(II) to iron@) may also occur by Fe2+ Fe3+ + e (5)
-
The reverse reaction, that is, the reduction of iron(III), may also occur leading to further leaching of metal sulfide (Boateng and Phillips, 1978), as described by NiS
-
+ 2Fe3+
Ni2+ + 2Fe2++ So
(6)
Rate Equations The leaching rate may be controlled by any of the process steps listed above, and, in practice, two or more of these steps may become important over the duration of the leaching process. The leaching reactions of eq 1and 2 may be assumed to be irreversible, and any resistance to leaching offered by the leaching products should be due to solid products which are not transferred from the reaction site. The rate of leaching may be governed by one or more of the following 3 equations. Case 1. Rate Control by Oxygen Transport across the Surface Film. The rate equation, assuming spherical particles, can be shown to be 1 ---
WNi
4nro2 dt
cw3d 1 - -2f -
dt 4nro2 = 2fkL(C(02)L - qo,,,)
(7)
where ro is the initial particle radius, kL is the mass t r m f e r coefficient (E D(0 /xL, where D o is the diffusivity of oxygen in a film odthickness xL), &(odL and C(odSare the oxygen concentrations a t the bulk liquid phase and the surface of the particle, respectively, and f is the fraction
of transported oxygen available for the leaching of nickel. Case 2. Rate control by Diffusion through the Product Material in the Reacted Shell. The rate equation for spherical particles is given by
where De(o) is the effective diffusivity (= D(o) ( € I T p ) , where e is t i e porosity and Tpis the tortuosity {actor), rc is the radius of the reacting core, and C(02),is the oxygen concentration at ita surface. Case 3. Rate Control by Chemical Reaction. A generalized rate expression may be written as
-- 1 4nr:
W02)
X
2f dt = 2f
kmnCNiSmC(02),n
(9)
It can be shown that the time, t , for a given conversion is related to the particle size as follows t a ro (for chemical reaction control) (10) t a Po2
(for diffusion through reacted shell control) (11) t a r01.5-2.0 (for film diffusion control)
(12)
Experimental Section The leaching testa were carried out in a glass-lined autoclave. The oxygen pressure was maintained by a highpressure oxygen source connected to the reaction vessel. The concentrate to be leached was first sized by means of agitated sieves, and the size distributions of the various size fractions were determined by graphic analysis by use of a TEKTRONIX 4051 microcomputer system. The pentlandite concentrate fraction to be leached was made into a 1kg 15% solids aqueous slurry, to which was added a calculated volume of 1 M bis(2-ethylhexyl) phosphoric acid (DEHPA) in kerosene. The volume of DEHPA added depended upon the estimated leachable metal content of the concentrate and was such as would provide a slight excess. About 1.9 mL of sulfuric acid was added to adjust the pH to the 1 to 2 range. A typical leaching experiment was carried out by first heating the empty autoclave to about 10 OC above the desired temperature. The slurfy, prepared in the removable glass liner, was then placed in the autoclave. This procedure usually resulted in the desired temperature at the commencement of the leach. Agitation and timing were begun at the introduction of oxygen at the desired pressure. Elemental analysis was carried out by neutron activation with gamma counting with the University of Toronto SLOWPOKE nuclear reactor. The solid phases were also analyzed by X-ray diffraction before and after leaching. Results and Discussion The rate of leaching was found to be proportional to the oxygen partial pressure for pressures up to 653 kPa. However, the fiit-order dependence of the rate on oxygen pressure decreased as the oxygen pressure was increased further to 791 kPa. One explanation for this result and similar results obtained for other sulfide minerals (Roman and Penner, 1973) may be the fact that, a t a given temperature, the oxygen concentration in the bulk liquid phase increases with the partial pressure, as may be estimated from Henry's law. At high pressures the decreased dependence of the leaching rate on the oxygen pressure may indicate a condition of excess oxygen in the bulk of the
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 3, 1984
559
8 40 i t W
30Q (li
20-
f
11
I
1
1
1
1
2 4 6 Oxygen pressure, 10' kPa
1
1
1
1
8 1 0
Figure 1. Effect of oxygen partial pressure on the leaching rate (concentrate DB12, particle size 63.5 pm).
0
40
20
60 Time, min
80
100
120
Figure 3. Effect of particle size on leaching of nickel (DB12). 70--I
$
60-
,
.
316 K
v
333 K 363 K
1398 K
'A -*-
-I
10-
20
I
I f
60 Time, min
40
80
100
120
Figure 4. The effect of temperature on the concentration of nickel in the leachate (concentrate DB5, size 28 Fm).
rosity could result from the inability of solid products to attain the compact structure of the original particle. The variation of the leaching rate with time may indicate change of the rate-controllingmechanism with the progress of the leaching process. The relations of Equations 10, 11, and 12 may be used to identify the rate-controlling mechanism by determining the value of z in the expression
~~0L$020 I
Time, min
Figure 2. Effect of concentrate grade on leaching (81 pm).
fluid. Under such conditions, the availability of reactive mineral components at the reaction surface may control the reaction rate. The approximate first-order dependence of the rate of leaching on the oxygen pressure is shown in Figure 1. If the rate were controlled by chemical reaction, the value of n in eq 9 would be 1. The results shown in Figure 2 for the leaching of three concentrates DB1 (0.9% Ni), DB5 (5% Ni), and DB12 (13% Ni) do not show any simple dependence of the leaching rate on the nickel concentration in the solid matrix. The rate may be considered to depend, in general terms, upon the effective accessibility of the reactive metal species to the leaching reagents and not on the concentration of the metal in the solid phase per se. Accessibility depends upon the proportions of associated minerals in the concentrates and the nature of leaching byproducts. X-ray diffraction analysis showed that in all three concentrates elemental sulfur was a constituent of the reacted solid. In DB1, geothite [a-FeO(OH)]was also found while carphosiderite or hydronium jarosite was found in reacted particles of DB12. These two species might also have been present in DB5, in amounts not sufficient for positive detection by X-ray diffraction. The composition of the particles helps to explain the leaching results shown in Figure 2. Solid byproducts tend to block access to the reactive nickel mineral and thereby decrease the leaching rate. The net change in porosity can be estimated for a given conversion from densities of concentrate particles, reactive mineral, and solid byproducts, and from the concentration of reactive mineral in the solid matrix. Although difficult to account for, further increase in po-
t
(Y
Toz
(13)
The results from leaching various size fractions of DB12 at 353 K and 791 kPa oxygen pressure are given in Figure 3. The calculated values of z were: 1.8 for the time interval 5-15 min, 1.2 for a 12.5-30 min interval, 1.8 for a 15-60 min interval, and 1.4 for a 45-115 min interval. These values may be interpreted as: film diffusion may be rate-controlling for a short time at the start of leaching, presumably until a situation of excess oxygen at the reaction surface is attained, after which chemical reaction becomes important. As the reaction surface moves toward the core of the particle, diffusion through the reacted shell (consisting of inert solid and solid byproducts) becomes rate-controlling. Leaching of nickel becomes very slow (almost ceasing) as the diffusion through the reacted shell becomes dominant. It is possible that about 15% of converted sulfur, which was solubilized, was oxidized by eq 3 at this stage, leading to a further increase in porosity. Beyond this stage, the value of 1.4 for z would indicate a possible interaction of chemical reaction and diffusion through the reacted shell in controlling the rate. The effect of temperature within the range 318-398 K is shown in Figure 4. Varying the leaching temperature did not appear to affect the general behavior described above. Increasing the temperature up to 363 K significantly increased the leaching rate in the final stages of reaction. However such a final stage increase in rate was not observed at 398 K presumably because of increased resistance to oxygen diffusion as sulfur melts in the in-
560
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 1-
I
-
3. 1984
,
45r 40 r
-
a
DB 12 (- 170
+ 200)
87 3 5 c
.3
n
2 25
10'' _-
E
-
E
< ... m f
-
f
20
.
60
a,
100
140
180
Size UA
z
t z ,U'-300
-
320
340
~
-
380 400
360 Temperature K
420
Figure 5. The effect of temperature on the rate of leaching of nickel (concentrate DB5, particle size 28 rm): line AB, if chemical reaction were controlling; line CD, if film diffusion were controlling; line EF, if product layer diffusion were controlling.
terstices of the reacted shell. The variation of experimental leaching rates with temperature may be compared to theoretical behavior based upon various rate-controlling mechanisms. For generation of the theoretical curves, the value o f f , the fraction of transported oxygen available for metal leaching, was evaluated from analysis of the leach products for the amount of sulfur converted to sulfate. The proportion of oxygen available for metal leaching (that is not consumed by the conversion of S to S042-)varied from 0.81 to 0.90 for the cases plotted in Figure 5. The dotted curve AB was generated by assuming that chemical reaction represented the dominant mechanism, by extrapolating from the data at 318 K through eq 9. Oxygen concentrations at various temperatures were estimated by means of Henry's law with proportionality constants obtained from Perry and Chilton (1973). An activation energy of 27.3 kJ, estimated from data when reaction control was indicated, was used for the extrapolation to higher temperatures. The dotted line CD was generated by assuming eq 3 and the molecular diffusivity of oxygen was calculated by assuming behavior in dilute solutions through eq 14 (Perry and Chilton, 1973)
where M,, V,, and X are the molecular weight, molar volume, and association number (assume X = 2.6 since the liquid phase is mostly water) of the liquid medium, respectively. The average viscosity, pav, was estimated through equations given by Perry and Chilton (1973) for immiscible liquid and particulate systems. For example, at 353 K, D(02)was estimated at 6.3 X lob5cm2/s (1.06 X loy4cm2/s for a pure water system). The dotted line EF was generated by assuming eq 14. Initial porosities of various fractions of the concentrate were determined by noting the weight of fluid required to fill any pore species. However, a microscopic examination showed that the particles were compact and the measured values were actually due to cracks which were random. Nevertheless, the average initial apparent porosity, to,was
20
60
100
140
180
Size UA
Figure 6. Effect of leaching on particle size (DB12 -170, +200). (a) size distribution of concentrate before leaching; (b) size distribution of leached particles.
0.005, and the estimated change in porosity, A€, due to leaching was 0.009 for DB5 and 0.006 for DB1 and DB12. These values and a tortuosity factor (T,)of 2.5 were assumed in estimating the effective diffusivities from
For example, at 353 K, the effective diffusivity of oxygen for DB5 was estimated as 3.53 X lo-' cmz/s. The analysis is subject to the following overall assumptions: (1)Particles were not spherical as assumed, although this assumption should not affect the general conclusions reached. (2) The constant f was introduced to account for the fact that oxygen consumption could occur by reactions other than that involving the leaching of nickel. Its value was determined, after the fact, from a sulfur balance. However, the value o f f might not be constant over the entire leaching period. (3) The particle size was assumed to remain constant during leaching. The particle size distributions shown in Figure 6 show that the proportion of small particles increased.
Conclusions Despite its obvious simplicity, the shrinking core model provides a satisfactory interpretation of the leaching of nickel in an oxygen-DEHPA system. Acknowledgment This work was supported by a grant from the Natural Sciences and Engineering Research Council of Canada. Registry No. Ni, 7440-02-0; DEHPA, 298-07-7; oxygen, 7782-44-7.
Literature Cited Bjoriing, G. "International Symposium on Hydrometallurgy", Evans, D. J. I.; Shoemaker, Ed.; A I M : Chicago, 1973 pp 701-717. Boateng. D. A. D. Ph.D. Thesis, 1979, University of Toronto, Toronto, Ont., Canada. Boateng, D. A. D.;PhllHps, C. R. Min. Sei. Eng. 1978, 70(3), 155-162. Forward, F. A.; Vletman, H. J . Met. 1959, 1 1 , 836-840. Letowski, F. Hydrometallurgy 1980. 6 , 121-133. Levenspiel, 0. "Chemical Reaction Engineering", 2nd ed.;Wiley: New York, 1972; Chapter 12.
Ind. Eng. Chem. Process Des. Dev. 1984, 23, 561-565 Loveday, B. K. J . S . AI?. Inst. Mln. Metall. 1075, 78(2), 16-19. Madsen, B. W.; Wadsworth, M. E.; roves, R. D. Trans. Soc. Min. Eng. A I M March 1075, 258, 69-74. Perry, R. H.; Chllton, C. H. "Chemlcai Englneers' Handbook", 5th ed.; kicC;rew-HIII: New York, 1973; Section 3. Prater, J. D.; Queneau, P. B.; Hudson, T. J. J . Met. 1970. 22, 23-27. Roach, 0. I. D.; Prosser, A. P. Trans. Inst. Mln. Metall. Sect. C Mlneral PIOcess Extr. Metall. 1078, 87, Cl29-Cl38. Roman, R. J.; Brenner, B. R. Mln. Scl. Eng. 1973. 5(1), 3.
501
Szekeiy, J.; Themeiis. N. J. Wiley: New York, 1971; Chapter 17. Vlzsolyi, E. et al. J . Met. 1067, 19, 52-59. Yogi. S.; Kunii, D. "5th Internetionel Symposium on Combustion"; Reinhold: New York, 1955; p 231.
Received for review June 30, 1981 Revised manuscript received February 22, 1983 Accepted September 27, 1983
Improving Sieve Tray Performance with Knitted Mesh Packing Dlno A. Spagnolo' and K. Tzatang Chuang Atomic Energy of Canade Limlted, Research Company, Chalk River Nuclear Laboratories, Chalk River, Ontario KW 1J0, Canada
Sieve trays have become popular contacting devices for mass and heat transfer applications because of their relative simplicity and low cost. As a result, much has been published in recent years on their hydraulic and mass transfer performance. However, little has appeared on improving contact performance by adding a shallow bed of packing on the trays. The hydraulic and mass transfer performance of sieve trays containing a 25-mm bed of a knitted mesh packing was studied for the Glrdier-Sulfide (GS) heavy water production process operating at 32 OC and 2.17 MPa. With packing, tray pressure drops increased, tray-to-tray entrainment decreased, and efficiency increased by 3 to 20 percentage points. The greatest efficiency gains occurred in the low gas flow range where the tray operated in the bubble regime.
Introduction Canada produces all of its heavy water by the dual temperature Girdler-Sulfide (GS) process (Galley, 1981). Here deuterium is chemically exchanged by contacting water and hydrogen sulfide gas countercurrently in large sieve tray columns which operate at 2 MPa pressure. The columns are divided into cold and hot sections that operate at approximately 30 and 130 O C , respectively. The exchange reaction which occurs in the liquid phase can be written as HDS HzO + H2S HDO
Table I. Column and Tray Geometry
The equilibrium of this reaction is such that in the cold section deuterium is transferred from the gas to the liquid phase while the reverse occurs in the hot section. This reaults in accumulation of deuterium between the cold and hot sections. Tray design for the Canadian GS plants was originally based on technology derived from systems other than GS (i.e., air-water, petrochemicals) and on the operating experience of smaller GS plants in the USA (Proctor, 1963; Garvin and Norton, 1968). Soon after the Canadian plants were started up, it became apparent that tray efficiencies were lower than design values (Neuburg and Chuang, 1982a,b) resulting in considerable loss of production. An R&D program was initiated with the objective of developing reliable correlations that could be used to predict, design, and improve the performance of sieve trays. During the course of the program, several novel ideas for improving the contact performance of conventional sieve trays were studied. One such idea was the installation of a shallow bed of mesh packing on the tray deck. This paper deals with experiments where a knitted mesh packing was placed on sieve trays. The purpose was to determine the influence of such packings on the hydraulic and mass transfer performance of sieve trays.
at 2 MPa pressure in the saturator before flowing to the three-tray test column. The water leaving the test column via the surge tank is depressurized from 2 MPa to 0.2 MPa before entering the flash tank where approximately 90% of the dissolved H2S is removed. Complete removal is accomplished by stripping at 108 OC in a 1.5-m packed column. The recovered H2S gas from the flash tank and stripper overheads is pressurized and returned to the circulating gas circuit. Gas leaving the test column passes through an entrainment separator to protect the blower. The test column was 0.311 m in diameter and consisted of three sieve trays spaced 0.381 m apart. Column and tray geometry are summarized in Table I. Five layers of a knitted 316L stainless steel mesh packing (Yorkmesh Type 421) were placed on each tray such that the rib direction between layers alternated giving a total packing thickness of approximately 25 mm on each tray. Tray performance was monitored by taking the following measurements. (1) Total Tray Pressure Drop ( hT). The total tray pressure drop is the head required to drive gas through the tray perforations ( h h ) and through the aerated liquid and packing on the tray ( h L ) . hT = h h hL (1) Measurements of hT on the middle and bottom trays were made with differential pressure cells. The accuracy in reading the recorder was f6.2 Pa (0.025 in. H20). Since
+
+
Experimental Description Figure 1 shows a simplified schematic diagram of the GS pilot plant. Treated feedwater is saturated with HZS 0196-4305/84/1123-0561$01.50/0
column diameter total column cross-sectional area tray active area downcomer area hole diameter open hole area tray thickness outlet weir height inlet weir height weir length downcomer flow clearance tray spacing
Published 1984
+
by the American Chemical Society
0.311 m 0.0760 m 2 0.0471 m 2 0.0145 m 2 12.7 mm 0.00683 m 2 1.6 mm 102 mm 0 0.268 m 6.4 mm 0.381 m
