Ind. Eng. Chem. Process Des. Dev. 1984, 23, 538-545
538
x = mole fraction of methane in solution
A a = solubility constant, eq 4 A@ = solubility constant, eq 4 y = activity coefficient
AA = solubility constant, eq 4 4 = fugacity coefficient vij = moles of ith ionic species per mole of jth dissolved salt e = residual, eq 6 Registry No. Methane, 74-82-8.
Literature Cited Arnold, D. S.;Plank, C. A.; Erlckson, E. E.; Pike, F. P. Ind. Eng. Chem., Chem. Eng. Data Ser. 1958, 3, 253-256. Blount, C. W.; Price, L. C.; Wenger, L. M.; Tarullo, M. Proc. 4th U.S. Gulf Coast GeopressurebOeothermaI Energy Conference: Research and Development: Dorfman M. H.; w. L. Flsher, Ed.; Center for Energy Studles, University of Texas, Austin, TX, 1979; Vol 3, pp 1225-1262. Blount, C. W.; Price, L. C.; Wenger, L. M.; Tarullo, M. Summary Report for DOE Contract DE-ASO8-78ET12145 (avakbie from C. W. Blount, Department of Geology, Idaho State Untverslty, Pocatello, ID), 1980. Claussen, W. F. J. Chem. fhys. 1951a. 79, 259-260. Claussen, W. F. J. Chem. fhys. IQSlb, 79, 1425-1426. Claussen. W. F.; Polglase, M. F. J. Am. Chem. Soc. 1952, 74, 4817-4819. Cramer, S . D. “Corrosion Problems In Energy Conversh and GenerariOn”; Tedmon, C. S.,Jr., Ed.: Electrochem. SOC.: Pennlngton, NJ, 1974 pp 251-262. Cramer, S. D. Ind. Eng. Chem. Process. Des. Dev. 1980, 79, 300-305. Cramer, S. D. Ind. Eng. Chem. Rocess Des. Dev. 1984, Communlcatlon In this Issue. Culberson, 0. L.; Horn, A. 8.; McKetta, J. J., Jr. Pet. Trans. A I M 1950, 789, 1-6. Culberson, 0. L.; McKetta, J. J.. Jr. Pet. Trans. A I M 1951, 792. 223-226. Dcdson, C. R.; Standing, M. 8. ”Drllllng and Production Practice”; American Petroleum Institute: Dallas, TX, 1944; pp 173-179. DWglaS, E. J. fhys. C h m . 1964, 68, 169-174. Dum, J. R.; Smith, N. 0.; Nagy, B. Geochem. Cosmochem. Acta 1961, 2 4 , 23-31. Ellis, A. J.; Goldlng, R. M. Am. J. Scl. 1983, 267, 47-60. Franks, F.; Gent, M.; Johnson, H. H. J. Chem. Soc. 1963. 2716-2723. Frollch, P. K.; Tauch, E. J.; Hogan, J. J.; Peer, A. A. Ind. Eng. Chem. 1931, 2 3 , 548-550. Fronlng, H. R; Jacoby, R. H.; Richards, W. L. Proc. Ann. Conv. Natl. &soline Assn Am. 1963, Tech. Papers 42. 32-39. Glew, D. N. J. fhys. Chem. 1962, 66, 605-609.
.
Gokcen. N. A. “Thermodynamics”; Techsctence; Hawthorne, CA, 1975; pp 77-86, 180-183. Gokcen, N. A.; Chang, E. T. Denkl Kagaku 1975, 43(5),232-237. Green, E. J.; Carrttt, D. E. ScMce July 14, 1967. 757, 191-193. Haas, J. L., Jr. “Physlcal Properties of the Coaxlstlng Phases and Thermochemical Properties of the HO , Component In BoWhg NaCl Solutions”; Bulletin 1421-A, U.S. Geological Survey, Reston. VA, 1978; pp 4-8. Haas, J. L., Jr. “Methane SolubHity in Water and Sodium Chlorlde Solutions”; Open Flie Report 78-1004, U.S. Geological Survey, Reston, VA, 1978. Helgeson, H. C. Am. J. Scl. 1968, 266(3). 129-166. Hlmmelblau. D. M. J. fhys. C t ” . 1959, 63, 1803-1808. Hlmmelblau, 0. M. J. Chem. Eng. &i?a 1960. 5(1), 10-15. Hougen. 0. A.; Watson, K. M.; Ragatz, R. A. “Chemical Process Principles”; Wky: New York, 1959; Part 1, pp 252-256, 345-347. Jones, P. H. “Proceedings, 1st Geopressured Geothermal Energy Conference”; Dorfman, M. H.; Deller, R. W.. Ed.; CONF-750612, Center for Energy Studies. College of Engineering, University of Texas, Austin, TX, 1975; pp 15-89. Kobayashl, R.; Katz, D. L. Pet. Trans. AIM€ 1949, 786(3), Technlcal Paper 2579, 66-77. Lewls, G. N.; Randall, M.; Pltzer, K. S.; Brewer, L. “Thermodynamics”; McOrew-HIII: New York, 1961a; pp 584-587. Lewls, G. N.; Randall, M.; Pitzer, K. S.; Brewer, L. “Thermodynamics”: McGraw-Hill: New York, 1961b pp 242-252. Malinln, S. D.; Savelyeva, N. I. Geochem. Int. 1979, 7972(5-6), 410-418. McAullffe, C. Nature (London) DOC 14, 1963, 200 1092-1093. McLeod, H. O., Jr.; Campbell, J. M. J . Pet. Techno/. 1961, 73(6), 590-594. Morrlson, J. R.; Blllett, F. J. Chem. SOC.1952, 3819-3822. Naumov, V. 8.; Khaklmov, A.; Khodakovski, I. L. Geochem. Int. Dec 1974, 77(1), 31-41. Orcutt, F. S.;Seevers, M. H. J. Blol. C t ” . 1936, 777, 501-507. O’SuHhran, T. D.; Smith, N. 0. J. fhys. Chem. 1970, 74, 1460-1466. Pray, H. A.; Schweickert, C. E.; Mlnnlch, 8. H. I d . Eng. Chem. 1952, 44(5), 1148-1 151. Price, L. C. Am. Assn. Pet. Geol. Bull. 1979, 63, 1527-1533. Ruetschl, P.; Amlie, R. F. J. fhys. Chem. 1966, 70. 718-723. Setchenow, A. 2.fhys. Chem. (Leiprlg) 1889#4 , 117. Smith, J. M.; Van Ness, H. C. “Chemical Engineering Thermodynamics”; McGrawHlll: New York, 1959a; pp 341-357. Smith. J. M.; Van Ness, H. C. “Chemical Englneering Thermodynamics”; McGraw-HIII: New York, 1959b; pp 118-128, 144-147, 404-412. Van Slyke, D. D.; Nelll, J. M. J. B&l. Chem. 1924, 67, 523-573. Zoss, L. M.; Suclu, S.N.; Slbbltt, W. L. Trans. Am. SOC.Mech. Eng. 1954, 76(1), 69-71.
Received for review October 12, 1982 Revised manuscript received October 6, 1983 Accepted October 29, 1983
Limestone/Dolomite Sulfation in a Vertkal Pneumatic Transport Reactor Llang-SMh Fan* and Sun11 SatlJa Department of Chemical Engineering, The Ohio State Unlversity, Columbus, Ohio 432 70
Byung C. Kim and Herman Nack Battelle Columbus Laboratory, Columbus, Ohlo 4320 7
The sulfation characteristics of “estone/dolomite sorbents in a vertlcal pneumatic transport reactor operated with or without the spent sorbent recycle are examined In this study. Experiments are performed in a 25.4-cm (10 in.) pilot scale reactor where limestone/ddomlte reacts with sulfur dioxide generated from coal combustion. A mathematical model is employed to account for the SWdkxide conversian and the sorbent utilization in the vertlcal pneumatic transport reactor as a scrubber for flue gas desulfurization with limestone/dolomite sorbents. The comparison between the experimental data and model prediction is conducted.
Introduction The combustion processes utilizing the pneumatic transport or circulating bed reactors are available in industry, e.g., Lurgi’s Process (Peterson et al., 1980) and 0196-4305/84/1123-0538$01.50/0
Battelle’s Process (Nack et al., 1977). Based on Battelle’s process, which is known as the Multisolid Fluidized Bed Combustion Process,Nack et al. (1976,1977) reported that more stringent sulfur dioxide emission standards can be 0 1984 American Chemical Society
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 3, 1984 530
met with a moderate limestone requirement. Furthermore, higher sorbent utilization was observed in this reactor compared to that in the conventional bubbling or turbulent bed reactors (Nack et al., 1977). The major factor contributing to high sorbent utilization in such a reactor is the sorbent size. The sorbent size for the pneumatic transport or circulating bed reactor is as small as 20 pm, while that for the bubbling or turbulent bed reactor is as large as lo00 pm or over. In addition, substantial amounts of spent sorbent are recycled in the operation of the pneumatic transport or circulating bed reactor. A key to widespread commercialization of fluidized bed combustion technology is the ability to accurately predict the retention of sulfur dioxide by a given sorbent. Limestone/dolomite are common sorbents used for fluidized bed combustion. Variables dictating utilization of limestone/dolomite sorbents in the combustor are numerous including reactivity, pore size distribution, surface area, and porosity of the calcined limestone sorbent, and transport properties of the combustor. Characterization of sorbents obtained from over 30 different locations in the country have recently been made by Fee et al. (1980). Extensive analysis of the sorbent utilization in the bubbling or turbulent fluidized bed combustors has also been explored (e.g., Chen and Wen, 1981; Marroquin et al., 1982; Fee et al., 1983). However, little analysis (Fan et al., 1984) has been done on sorbent utilization in the pneumatic transport or circulating bed combustors. Recently, Fan et al. (1984) developed a heterogeneous model to account for general characteristics of noncatalytic gas-solid reactions in a vertical pneumatic transport reactor. The model was also partially verified with experimental data on sulfur dioxide retention by limestone sorbents taken from a bench-scale multisolid pneumatic transport reactor. This paper is intended to examine the specific characteristics of sulfation of limestone/dolomi& sorbents in a vertical pneumatic transport reactor operated with or without the spent sorbent recycle. Experimental data are obtained from a 25.4-cm (10 in.) i.d. pilot scale reactor where limestone reads with sulfur dioxide generated from coal combustion. The design and operation of the reactor is analogous to Battelle's Multisolid Fluidized Bed Combustion Process. The mathematical model of Fan et al. (1984) is utilized to extensively simulate the conversion of sorbents and sulfur dioxide in the reactor. This simulation accounts for the sorbent utilization in the vertical pneumatic transport reactor as a scrubber for flue gas desulfurization using limestone/dolomite sorbents. The model is also utilized to elucidate the pilot scale experimental data of the sulfur dioxide conversion obtained in this study. The Model It is assumed that the pneumatic transport reactor is operated in an isothermal condition. The limestone particles react with SO2 in the presence of oxygen to form CaS04 with the reaction represented in two steps by CaC03(s) CaO(s) + C02(g) (1) CaO(s) + SOz(g) + 1/20,(g) CaS04(s) (11)
-
-
At high temperatures, reaction I is relatively fast compared to reaction 11. Thus, only reaction I1 is considered in the determination of overall rate of reaction for sulfation. The rate of sulfation, ?,, can be expressed by
YE = kvC2nCEm (1) where k, is the volumetric rate constant. 7, has been shown to be first order with respect to the sulfur dioxide concentration (Borgwardt, 1970; Fee et al., 1980). The first
order with respect to the solid reactant concentration for 9, has been employed by Fan et al. (1981). In the present work, the first order with respect to both the sulfur dioxide concentration and solid reactant concentration are used for 9,. It is also assumed that the entrance effects in the pneumatic transport reactor can be neglected, the flows of the gas and solid follow the plug flow pattern, and the solid particles are spherical and uniform in size (Fan, 1981). A material balance of the gas reactant (SO,) in the gas phase in dimensionless form gives rise to (Fan et al., 1984).
The boundary condition for eq 2 is y = 0; f1 = 1 (3) A material balance of the gas reactant and the solid reactant in the solid phase yields, respectively (Fan et al., 1984)
(5) The boundary conditions for eq 4 and 5 are y = 0; f, = 0; g, = 1
(6) (7)
x = 1; DA
afz
= NSha(f1 - f2) ax
(8)
The reaction of CaO with SO2 is accompanied by a substantial expansion of the volume of the grains which constitute the CaO particles. The volume expansion is due to the fact that the molar volume of CaSO, is about three times that of CaO. The reacting particles usually retain their original gross external volume, and thus the reaction with SO2 causes decrease in the porosity. The relation between the void fraction and the solid concentration can be assumed to follow the equation (Wen, 1968; Fan et al., 1977) given by
' = '0 - Y ( 1 - g1)
(9)
where eo is the void fraction of the particle at the initial condition and y is a positive constant which characterizes the decrease in the void fraction with the solid conversion. Dimensionless diffusivity DA in eq 4 accounts for the overall rate of gas diffusion in the porous particle of the solid reactant. It is a complex entity constantly changing with variations in the characteristics of the pore network such as pore shape, pore size, and pore size distribution and with variations in the physical conditions such as temperature and pressure. For the diffusion of sulfur dioxide in calcium oxide particles, a simple relation for the effective diffusivity, DeA, was used by Hartman and Coughlin (1976) and Fan et al. (1981), which has the form DeA = D,' (10) where D, is the molecular diffusivity divided by the tortuosity. DeA can be related to DeAO by DeA
= DeAO
(11) '0
Equation 11 is used in this study to account for the dif-
540
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 3, 1984
fusivity variation with the void fraction in the particle. The conversion for the solid reactant is defined as 1
glx2 dx
x, = 1- 3
(12)
The SOz conversion in the reactor outlet can be evaluated by
x = 1 - ( f i + hf,)
(13)
where f 2 is the average dimensionless SO2 concentration in the sorbent which can be expressed by 1
72
= 3 1 € f i x 2 dx
Evaluation of Hydrodynamic a n d Mass Transfer Properties for the Model The mass balance for the solid particle and empirical correlation available in the literature for the solid friction factor, f,,and the linear particle velocity, U can be used to obtain the hydrodynamic properties for t i e pneumatic transport reactor. The mass balance for the solid particle in the reactor yields c b = 1 -
4W8 Pp7@Up
U p is obtained by (Yang 1977, 1978)
f ,can be evaluated by the following empirical correlation equations (Yang, 1978)
f,
=
f,
=
Ut can be obtained by (Yang, 1973)
ut = 0.153d,1.14g0.71(p, - P,)O.~’ (for 2.0
< (Re), < 1000)
~0.43p,0.29
(19) and
ut =
dp2(Pp - P& 18/1
(for (Re), 5 0.1)
(20)
where
Equations 15 through 18 are solved simultaneously to obtain the hydrodynamic properties including cb, f p , and Up. The values of q, and Upare required for the model equations.
\
\
EXTERNAL i E A T EXCHANGER
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 3, 1984 100
100
60
60
541
w 60
60
!i
x'
X
40
40
20 PARAMETERtype O f
0
10-8
10-6
c O , g-mole/cm3
g/sec
14,.
Sorbent
0
Figure 2. Effecta of the sorbent flow rate on the SO2conversion and
Figure 5. Effects of the inlet SO2 concentration on SO2 conversion.
the sorbent utilization for the four sorbents. 100
100
I.
I
20 PARAMETER: t y p e o f s o r b e n t
- so2
EO
60
-60-
ap
ae 6 0
60
2
.
40
20
20
__
0 0
eorn.nt
40 -
2o
t
w 15
I
x'
X
40
DD"".P.lo" .O?D."t ut111zation
. X
E001 ' e _
.'
10
//--
/--I-
---/!-.-----r 9202
9701
"tl~ir.tl0"
0
EO0
600
400
200
1000 Ca/S, -
Ugs, c m / s e c
Figure 6. Effects of Ca/S on SO2conversion and sorbent utilization
Figure 3. Effects of the superficial gas velocity on SO2conversion
with no recycle of calcined sorbents.
and sorbent utilization for the four sorbents.
50
I30
-
40
ae 6 0 -
30
60
80
w
2
60
x'
. X
40 -
20
40 10
20 0 '
PARAMETER t y o e o f 5 C r b e n t
,
3
0
40
80
123
160
,
I
,
2
3
4
200 Ca/S,
0
-
Figure 7. Effects of Ca/S on SO2conversion and sorbent utilization Figure 4. Effects of the sorbent diameter on SO2conversion for the four sorbents.
Results and Discussion The numerical simulation and comparison of model prediction with the experimental data obtained in this study are given in the following. Numerical Simulation. The sulfation behavior of limestone/dolomite sorbents in a vertical pneumatic transport reactor with or without the spent sorbent recycle
for recirculation ratio of 5.
is numerically simulated. Four limestone/dolomite sorbents indexed by 5101,8001,9202,and 9701 which were previously characterized by Fee et al., (1980)are used for the simulation. Note that these sorbents represent a wide range of sorbent reactivity properties reported by them. The properties which characterize these sorbents include the volumetric rate constant, k,, the variation constant of solid porosity, y,the solid initial effective diffusivity, DeAo, and the initial solid porosity, eo. The values of these
542
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 3, 1984
Table I. Specification and Operating Condition of the Experimental System combustor internal diameter combustor height height of dense bed (settled) height of dense bed (expanded) coal feed rate sulfur percentage in coal sand circulation rate total solids circulation rate CaCO, percentage in limestone material of dense bed average bed temperature average diameter of limestone particle residence time of limestone particles in dense bed Ca/Sa R,glsb a
1.95 63
2.0 65
Calcium-sulfur molar ratio.
2.14 69
2.26 13
2.55 83
25.4 cm (10 in.) 838.2 cm (27.5 f t ) 30.5 cm (12 in.) 152.4 cm (60 in.) 11.4 g/s ( 9 0 lb/h) 4% 433.3 g/s (3432 lb/h) 502-555 g/s (3975-4392 lb/h) 94.0% iron ore (crushed) 843 "C (1550 O F ) 250-350 p m -8 s 2.68 87
2.86 93
3.03 98
3.2 104
3.4 110
3.5 114
Sorbent circulation rate.
Table 11. Properties of the Various Calcined Sorbents Used in the Model Simulation Obtained a t 1 1 2 3 K (Fee et al., 1980; Fan et al., 1981) sorbent ident no. 5101 8001 9202 9701
C,
type of sorbent dolomite limestone limestone limestone
9
cm2/s 0.402 0.095 0.0024 0.0008 ~
~~~
Table 111. Nominal Values of the Parameters Employed for Figures 2 through 8 pp =
k" cm3/g-mol s 3411.9 70 002.8 28 684.1 7799.1
DC?AO>
I
g-mol/cm3 0.0138 0.0221 0.0245 0.0226
€0
Y
0.6 0.53 0.55 0.64
0.64 0.75 0.50 0.85
100
2.5 g/cm3 for Figures 2-8
80
p g = 0.004 g/cm3 for Figures 2-8
D = 7.62 cm for Figures 2-8 W, = 0-300 g/s for Figure 2 = 167 g/s for Figures 3-5 = 4.5-100 g/s for Figures 6-8 Ugs= 500 cm/s for Figures 2, 4-8 = 100-1000 cm/s for Figure 3 d, = 50 p m for Figures 2, 3, 5-8 = 20-200 pm for Figure 4 C, = 3.25 x l o + g-mol/cm3 for Figures 2, 3 = 3.25 x l o - * g-mol/cm3 for Figure 4 g-mol/cm3 for Figures 6-8 = 0.1 X g-mol/cm3 for Figure 5 = 3 x lo-* to L = 800 cm for Figures 2-8
properties for calcined sorbents used in the model simulation are listed in Table 11. It is noted in the table that sorbent 5101 has the highest initial effective diffusivity which is followed in order by sorbents 8001,9202, and 9701. Sorbent 8001 has the highest reaction rate constant, which is followed in order by sorbents 9202,9701, and 5101. The performance of these sorbents in the reactor, however, is described by a nonlinear relationship among various parameters of the sorbent along with the transport properties of the reactor. The results of the numerical simulation for various sorbents are shown in Figures 2 through 8. The nominal values of the operating parameters employed in these figures are given in Table 111. Parametric effects of the sorbent flow rate, superficial gas velocity, sorbent diameter, and the initial sulfur dioxide concentration on the sulfur dioxide conversion, and the sorbent utilization without the spent sorbent recycle are shown in Figures 2 through 5. In Figures 6 through 8, effects of the recycle rate of spent sorbents and calcium-sulfur molar ratio (Ca/S) on the reactor performance are described. It should be noted in the table that the sorbent density, pp of 2.5 g/cm3 is used, which represents the density of uncalcined limestone particles. The density of the particle would reduce to 1.5 g/cm3 upon complete calcination in the reactor. It would then increase with the sulfation reaction, with the maximum value of 3.4 g/cm3 upon complete sulfation. The
60
i 40
20
0
1
1
I
2
3
4
'C
Ca/S. Figure 8. Effech of Ca/S on SO2 conversion and sorbent utilization for recirculation ratio of 10.
analysis conducted in this study indicates that under the dilute flow conditions, the change in the particle density has practically negligible effect on the hydrodynamic properties of the reactor. Thus, the volumetric flow rate of the solid particles practically remains constant throughout the reactor. Since W, in eq 15 is taken as the uncalcined limestone sorbent feed rate, the corresponding density of 2.5 g/cm3 would only yield an appropriate solid particle volumetric flow rate in the reactor. The effect of the solid flow rate on the sorbent utilization and the SOz conversion for these sorbents is shown in Figure 2. For sorbents 8001 and 9202, the complete conversion of the gas phase is achieved at the sorbent flow rate of 170 and 220 g/s, respectively, while for sorbents 9701 and 5101 the gas conversion monotonically increases with the sorbent flow rate. The sorbent utilization monotonically decreases with the sorbent flow rate for all the four sorbents. For a given solid flow rate, sorbent 8001 yields the highest SO2 conversion and the sorbent utilization and is followed in order by sorbents 9202,9701, and 5101. It is noted that sorbent 8001 has the highest volumetric reaction rate constant, k,, and is also followed in order by sorbents 9202, 9701 and 5101.
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 3, 1984 543
Figure 3 shows the effect of the superficial gas velocity on the sorbent utilization and the SO2conversion for these sorbents. For sorbents 8001 and 9202 the complete SO2 conversion is reached a t the gas velocity of 400 cm/s or below and it decreases slightly as the gas velocity increases up to 1000 cm/s. For sorbents 9701 and 5101, the SO2 conversion dramatically decreases with the increase of the gas velocity from 100 to lo00 cm/s. At a given gas velocity, the SO2conversion decreases in the same sorbent order as does the reaction rate constant, It,. The variation of sorbent utilization for sorbents 9701,9202, and 5101 with the superficial gas velocity shows a maximum in the velocity range of 200-800 cm/s. High gas superficial velocity gives rise to a relatively high bulk concentration of sulfur dioxide in the reactor which promotes the overall rate of reaction. However, high gas superficial velocity also gives rise to a relatively low sorbent holdup which demotes the overall rate of reaction. The consequence of these two counteracting effects on the sorbent utilization yields a maximum in the relationship between sorbent utilization and superficial gas velocity. Figure 4 shows the effect of the sorbent diameter on the SO2conversion for the four sorbents. The SO2conversion for sorbents 8001,9202, and 9701 decreases with the increase in the sorbent diameter while it increases with the increase in the sorbent diameter for sorbent 5101. This is due to the combined effect in that for a small value of DeAO,the increase in the sorbent diameter results in substantial decrease of the overall rate of reaction and moderate increase of the residence time of the sorbent and hence, the decrease of the SOz conversion. On the other hand, for a large value of DeAO,the increase in the sorbent diameter results in the essentially unvaried overall rate of reaction and the moderate increase of the residence time of the sorbent and hence, the increase of the SO2 conversion. I t is noted that sorbents 9701 and 9202 have relatively low initial gas diffusivity while sorbent 5101 has high initial gas diffusivity. The effect of the inlet SO2 concentration on the SO2 conversion is described in Figure 5. The increase of the inlet SO2 concentration from 3 x lo-* g-mol/cm3 to lo4 g-mol/cm3,decreases the SO2conversion for sorbents 9202 and 8001 by a considerable amount, while for sorbents 5101 and 9701, the decrease in the SOz conversion is not substantial. In industrial operation, only part of the spent sorbent is discharged from the reactor with the remaining spent sorbent recycled back to the reactor (Nack et al., 1977). The amount of the sorbent discharged is then made up by the fresh sorbent. Calcium-sulfur molar ratio (Ca/S) is defined as the ratio of moles of Ca in fresh feed to the moles of S in flue gas entering the reactor. Effect of different operational variables on the gas conversion in an isothermal pneumatic transport reactor with spent sorbent recycle is similar to that without spent sorbent recycle exhibited in Figures 2 through 5. To account for the inlet sorbent concentration in the reactor with the spent sorbent recycle, it is assumed that the concentration of the solid reactant in the particle at the reactor inlet is uniform and can be evaluated from a simple material balance which considers the reactant concentration in the fresh sorbent and the recycle spent sorbent. Note that considerable nonuniformity of the solid reactant concentration and porosity in the particle would develop under the high extent of the solid conversion in the reactor and thus, the assumption employed here is valid only for the low extent of the solid conversion in the reactor. Figures 6 through 8 show the effect of Ca/S on the gas conversion and the
Table IV. Numerical Values of the Parameters Employed for Figures 9 and 10 ~~
= 2.5 g/cm3 = 0.004 g/cm3 D = 25.4 cm DeAo = 0.1 cm'/s k , = 0.295 X l o 5 cm3/(g-mols) Ugs= 609.6 cm/s for Figure 9 = 797.2 cmls for Figure 10 d, = 350 Mm for Figure 9 = 250 Mm for Figure 1 0 C, = 0.021 g-mol/cm3 C, = 4.6 X gmol/cm3 for Figure 9 = 3.5 X g-mol/cm3 for Figure 10 p = 4.5 x g/(cm s) L = 838.2 cm E , = 0.6 y = 0.7 pp pg
sorbent utilization for the four sorbents in the pneumatic transport reactor operated with spent sorbent recycle. The recirculation ratios, defined as the ratio of the amount of the recycle spent limestane sorbent to the amount of the make-up fresh feed, of 0, 5, and 10 are used for Figures 6 through 8, respectively. It is noted that the increase of the recirculation ratio at the same Ca/S increases the total Ca flow rate to the reactor. The SO2conversion and the sorbent utilization increase with increase in Ca/S for the same recirculation ratio as exhibited in Figures 6 through 8. The increase in recirculation ratio at the same Ca/S molar ratio increases the solids holdup in the reactor and thus decreases the gas residence time in the reactor. The combined effect of these two phenomena results in the increase of the gas conversion and the sorbent utilization up to the recirculation ratio of ten as shown in Figures 6 through 8. Experimental Data and Model Verification. The present model is used to analyze the experimental data on limestone sulfation in the vertical pneumatic transport reactor, or specifically, Battelle's Multisolid Fluidized Bed combustor. The values of the parameters used in this simulation are listed in Table IV. As described earlier, there are two sections in the reactor, namely, dilute section and dense section. The dense section is substantially smaller than the dilute section. However, the dense section provides higher residence time for the fine sorbent than does the dilute section. The residence time of the fine sorbent in the dense section, t,, for the operating condition considered in this study was experimentally measured as 8 s, as given in Table I. The volume fraction of the sorbent in the dense bed can be calculated by Fit,
=Vl
(23)
where Fl and VIare the volumetric feed rate of the sorbent and the volume of the dense section, respectively. The particle linear velocity in the dense section, Up, can be calculated by
u,,
UPS
=€81
where Up, is the superficial velocity of the sorbent. In the dilute section, it is assumed that the entrained particles including limestone, coal ash, coal carbon, and inert material (sands) act independently. The hydrodynamic and mass transfer properties of each type of particle, in the dilute section, can be evaluated by eq 15 through 22.
544
Ind. Eng. Chem. Process Des. Dev., Vol. 23, No. 3, 1984
Ugg
t
95
x
60
73
-
:.c
Exoe-:menta:
, 5.5
2.0
2.5
Data
, 3.3
I
1
75
dp
7 9 7 . 2 ."reo 250
/
vm
/ -
-
1.0
3.5
-
1.5
2.0
2.5
3.0
3.5
Ca/s. -
Ca/'S, -
Figure 9. Comparison of the model prediction with the experi-
Figure 10. Comparison of the model prediction with the experi-
mental data obtained from the 25.4 cm i.d. pilot unit.
mental data obtained from the 25.4 cm i.d. pilot unit.
The volume fraction of the fine sorbent in the dilute section, eS2, can be calculated by the hydrodynamic properties and the solid flow rate of all types of particles. Equations 2 through 8 are solved for both the dense and dilute sections. The term (1- cb) in eq 2 and Up in eq 4 and 5 are replaced by eel and Up, for the dense section and by cs2 and Up, for the dilute section. Note that the following relationship holds u g l e b l = Ug2eb2 = u gs (25)
desired SO2 conversion is achieved. Increase in the spent sorbent recirculation ratio at a given Ca/S also results in an increase of the SO2 conversion and sorbent utilization for the range of recirculation ratios considered in this study. Experiments were conducted with a pilot scale reactor in which limestone is reacted with SO2 generated from coal combustion. The sulfur retention by limestone particles is found to increase with the increase of Ca/S. The model was shown to predict the variation of SO2 conversion with Ca/S reasonably well.
Most of the parameters listed in Table IV are operating variables or properties of the sorbents. Operating variables include D , U,,, p , d,, Co, w, and L and the properties of the sorbent incluAe pp, Dd0, Cd,and e@ The values of the remaining parameters listed in Table IV, i.e., y and k,, are obtained by fitting the model with one set of experimental data. The experimental conditions for this selected set of data are superficial gas velocity = 609.6 cm/s, calciumsulfur molar ratio (Ca/S) = 1.95, and sulfur retention = 84.9%. The same values of y and k, are then used in the model to simulate the sulfur retention for remaining data. Note that the values y (0.7) and k, (29 500 cm3/g-mol s) obtained in this manner for the sorbent used in this study are in the same range as that obtained by Fan et al. (1981) for similar limestone sorbents under similar reaction conditions. The comparisons between experimental data and model prediction for sulfur retention as a function of calciumsulfur molar ratio are presented in Figures 9 and 10. As expected, it is seen that sulfur retention increases with the calcium-sulfur molar ratio. The comparison is shown to be satisfactory. Concluding Remarks The limestone/dolomite sorbent utilization in a vertical pneumatic transport reactor as a scrubber for flue gas desulfurization is extensively analyzed by a mathematical model of Fan et al. (1984). Effects of various operating parameters on the sulfur retention and sorbent utilization are simulated for four representative sorbents previously characterized by Fee et al. (1980) and evaluated by Fan et al. (1981). The results indicate that effects of the type of sorbents on the sorbent utilization and the SO2 conversion are significant. Furthermore, the sensitivity of different parameters including the sorbent flow rate, superficial gas velocity, and particle diameter on the SO2 conversion is affected by the properties of the sorbents. The sorbent utilization can be maximized by operating the reactor at the minimum sorbent flow rate at which the
Nomenclature
A = dimensionless group defined as ro2U,/LDeAo Co = inlet SOz concentration, ~ o ~ J L ~ C1 = SO2 concentration in the gas phase, mol/L3 C2 = SO2 concentration in the solid phase, mol/L3 CBo= initial solid reactant concentration, mol/L3 C, = solid reactant concentration in the solid phase, mol/L3 D = inside diameter of the reactor, L DeA = effective diffusivity of the gas in the sorbent, L2/8 DeAO = initial effective diffusivity of the gas in the sorbent, L2/8 DA =' DeA/DeAO DM= molecular diffusivity in the gas phase, L2/8 d = sorbent diameter, L #=dimensionless group defined as ( 3 0 - eb)kdL)/(UgebrO) F, = volumetric feed rate of the sorbent, L3/8 f l = CdCO
fi = C2/CO f2 fp
= average dimensionless SO2 concentration in the solid phase = solid friction factor
g1 = c,/c,o G = dimensionless group defined as (k,C&)/Up g = gravitational acceleration, L/82 g, = conversion factor h = u,,/ug, k , = volumetric reaction rate constant, L3/8(mol) kd = gas film mass transfer coefficient, L / 0 L = length of the reactor, L m = order of reaction with respect to the solid reactant n = order of reaction with respect to the fluid reactant Nshe= Sherwood number, defined as rOkd/DeAO R = limestone sorbent recirculation rate, MI6 r = radial position within the sorbent, L
ro = radius of the sorbent, L (Re), = Reynolds number, defined as d,(Ug - U , ) p g / ~ (Re), = Reynolds number, defined as d Qpg/,u S c = Schmidt number, defined as f i / p $ M t , = residence time of the sorbent in the dense section, 0 U g = linear gas velocity, L / 0
Ind. Eng. Chem. Process Des. Dev. 1904, 23, 545-552
= linear gas velocity in the dense section, L/O
545
Fan, L-S.; Mlyanaml, K.; Fan, L. T. Chem. Eng. J . 1977, 13(1), 13. Fan, L-S. Chem. Eng. J . 1981, 21, 179. Fan, L-S.; Satlja, S.; Wllson, I.; Fee, D.; Myles, K. M.; Johnson, 1. “Sulfation Klnetlcs of Calcined Llmestones/Dolomltes In a Thermogravlmetrlc Analyzer: Experlment. Modeling and Slmulation”, presented at the AIChE 74th Annual Meetlng, New Orleans, LA, Nov 8-12, 1981; Chem. Eng. J . In D r e s s . Fan,-L-S.; Satija, S.; Kim, B. C.; Nack, H. AIChEJ. 1084, 30(1),21. Fee, D. C.; Wllson, W. 1.; Myles, K. M.; Johnson, I.; Fan, L A . Chem. Eng. Scl. 1083. 38(11).1917. Fee, D. C.; Wilson, W. I.; Shearer, I. A.; Lenz, J.; Fan, L-S.; Myles. K. M.; Johnson, I. “Sorbent UtWlratlon Prediction Methodology-Sulfur Control In FluMIzed Bed Combustors”, ANL/CEN/FE-80-10, Argonne Natlonal Laboratory, 1980; 484 pages. Gear, C. W. “Numerlcal Initial Value Problems In Ordinary Dlfferentlal Equatlons”; Prentlce-Hall: Englewocd, NJ. 1971; pp 102-221. Hartman. M.; Coughlln, R. W. AICh€ J . 1978, 22(3), 490. Marroquln, 0.; Fan, L-S.; Fee, D. C.; Myles, K. M. “An Analytical Model for Freeboard and In-Bed Limestone Sulfation In Fluidized Bed Coal Combustors”, presented at the 13th Annual Meeting of the Fine Partlcle Soclety, Chicago, IL, Aprll 12-14, 1982. Nack, H.; Felton, G. W.; Llu, K. T. “Battelle’s Multlsolid FluMizebBed Combustion Process”, paper presented at the 5th Internatlonal Conference on Fluldlzed Bed Combustion, Washington, DC, Dec 1977. Nack, H.; Kiang, K. D.; Lln, K. J.; Murphy, K. S.; Smlthson, 0. R., Jr.; Oxley, J. H. “Fluldlzatlon Technology”; Kearlns, D. L.. Ed.; Hemisphere: Washington, DC, 1976; Vol. 2, p 739. Peterson, V.; Daradlnos. 0.;Serbent, H.; Schmldt, H-W. “Combustion in the Circulating FluM Bed: An Alternetlve Approach In Energy Supply and Environmental Protection”, Proceedings of the Sixth International Conference on Fluldlzed Bed Combustlon, 1980 Vol. 11, p 212. Setterfield, C. N. “Mass Transfer In Heterogeneous Catalysis”; MIT Press: Cambrldge, MA, 1969. Slncovec, R. F.; Madson, N. K. ACM Trans. Math. Software 1975, l(3) 232. Vlchnevetsky, R. Simuktion, Aprll 1971, 168. Wen, C. Y. Ind. Eng. Chem. 1968 60(9), 34. Yang, W. C. Ind. Eng. Chem. Fundam. 1973, 12, 349. Yang, W. C. J . Powder Bulk Solids Techno/. 1977, 1 , 89. Yang, W. C. AIChE J . 1978, 24(3), 548.
= linear gas velocity in the dilute section, L/O superficial gas velocity, L / O linear sorbent velocity, L/O = linear sorbent velocity in the dense section, L/O = linear sorbent velocity in the dilute section, L/O = superficial sorbent velocity, L/O $= terminal velocity of the sorbent, LIB VI = volume of the dense section, L3 W,= sorbent flow rate, M/O X = SO2 conversion at the reactor outlet X , = solid reactant conversion x = rfro
8
y =z/L z = axial distance along the reactor, L
Greek Letters t b = void fraction in the reactor t b l = void fraction in the dense section
= void fraction in the dilute section = void fraction in the sorbent to = initial void fraction in the sorbent tal = volume fraction of the sorbent in the dense section ts2 = volume fraction of the sorbent in the dilute section 9, = local solid reaction rate, mol/L30 y = constant defined by eq 9 pg = gas density, M I L 3 pp = sorbent density, M / L 3 p = gas viscosity, MILO 4” = ~ o ~ ~ ~ ” ~ , o ~ / ~ ~ , , , ~ l ” 2 Registry No. Sulfur dioxide, 7446-09-5; dolomite, 16389-88-1. tb2 t
Literature Cited
Received for review May 10, 1982 Revised manuscript received September 29, 1983 Accepted October 31, 1983
Borgwardt, R. M. Environ. Sd. Techno/. 1970, 4(1), 59. Chen, L. H.; Wen. C. Y. “Reaction In Fluldlzed Bed Freeboard”. presented at the AIChE Annual Meeting, New Orleans, LA, Nov 8-12. 1981.
Polyurethane Waste Recycling. 1 Glycolysis and Hydroglycolysis of Water-Blown Foams John Gerlock, Jacob Brarlaw, and Mlklo Zlnbo Research Staff, Ford Motor Company, Dearborn, Mlchlgan 48 12 1
I n this paper, glycolysis of toluenedlisocyanate based water-blown polyurethane foam has been examined by high performance liquid chromatography and gel permeation chromatography to determine the product distribution. Glycolysis with diethylene glycol (DEG) yields toluenedlamlne (TDA), TDA mono- and di- DEG carbamates, a series of urea-linked mono- and dC DEG carbamate TDA oligomers, and polyether triol (polyol). The complexity of the product mixture suggests problems in applying simple glycolysis to the recovery of mixed and/or contaminated polyurethane wastes. A simpler product mixture results when water and a base catalyst are added to the glycolysis reaction (hydroglycolysis). Hydroglycolysis yields TDA and polyol as principal products. Data for the rate of the hydroglycolysis reaction are presented in the temperature range of 150 to 190 O C . These results suggest that hydroglycolysis could be used to recover polyols from mlxed and/or contaminated water-blown polyurethane wastes.
Introduction As a continuation of our interest in developing practical methods to recycle water-blown polyurethane wastes (Mahoney et al., 1974; Gerlock et al., 1980), we have examined recycling by glycolysis in some detail. Glycolytic recycling of polyurethane foam is currently the only method proven in large-scale commercial practice among the three most promising schemes proposed to date: (1) glycolysis (Ulrich, 1978; Hill, 1955; Bayer et al., 1950), (2) pyrolysis (Albert and Tacke, 1955, U.S. Patent 3 143 515), 0198-4305/84/1123-0545$01.50/0
(3) steam hydrolysis (Mahoney et al., 1974; Gerlock et al., 1980; Campbell and Meluch, 1976). Nippon Soflan of Japan currently operates a 1.3 X lo6 lb/year polyurethane recycling plant based on the Upjohn glycolysis process. Glycolylic recycling is performed by “dissolving” a polyurethane waste in a nearly equal weight of high boiling diol or mixture of diols heated to between 190 and 230 “C in an inert atmosphere. The fact that the immediate dissolution mixture can be sufficiently ”polyol-like”, and directly substituted for up to 10% by weight virgin poly01 0 1984 American Chemlcal Society
