Multistage Cocurrent Contacting Va por-Liquid E quilibrium Unit Charles C. Peiffer, Robert H. McCormick’, and Merrell R. Fenske Department of Chemical Engineering, Pennsylvania State University, University Park, Pa. 16802 A unit was designed and tested to investigate cocurrent contacting of vapor and liquid streams as a method for determining accurate vapor-liquid equilibrium data. It consisted of four individual vertically stacked stages of six-in. i.d. borosilicate glass pipe. Each stage contained a contactor and a disengaging section for the vapor and liquid phases. The flow of the phases on each stage i s cocurrent, while the overall flow in the unit i s countercurrent. Two six-in. long annular contactors, %and %-in., were tested for efficiency, with and without Metex screen packing. The contactors were formed by inserting a %-in. 0.d. glass tube inside one- and twoin. i.d. glass tubes, respectively. The stage efficiency proved to be largely dependent upon the throughput and can be maintained at 100% for 1000 grams per hr per cm2, corresponding to 6 and 28 liters per hr for the f/4- and %-in. annular contactors, respectively. At a throughput of 5000 grams per hr per cm2, or 26 and 120 liters per hr, this efficiency decreases to 85%.
M a n y situations arise when the needed engineering accuracy for relative volatility data exceeds the available analytical accuracy. Today, particularly, when compounds of high purity are required for many industrial products, it is necessary to know the factors that affect the efficient separation of the compounds in multicomponent systems. Until more accurate and less time-consuming correlations for predicting the needed information are available, the usual experimental methods for obtaining the necessary vapor-liquid equilibria must be resorted to. A multistage vapor-liquid equilibrium unit would serve as a “calculating machine” for obtaining such data. A publication for the design of a versatile six-stage countercurrent contacting unit (McCormick et al., 1962) illustrates the minimization of errors resulting from analytical limitations in obtaining vapor-liquid equilibrium data. In this unit, each stage consists of a still equipped with a condenser, a heater, a mechanical mixer, and sample points for the vapor and liquid phases. The advantage of using a unit with more than one stage for vapor-liquid equilibria is illustrated in the following equation (Fenske, 1932):
This equation permits the calculation of relative volatility from the vapor and liquid compositions resulting from a separation over n stages. Thus, the relative volatility is then a function of the n t h root of the enrichment factor. Although a conventional (packed) fractionating column may be used for obtaining relative volatility data, certain shortcomings are inherent in such a unit. I n this technique, calibration of the number of theoretical plates with a known and similar system is necessary. Also, such a unit ’ T o whom correspondence should be addressed.
380 Ind. Eng. Chem. Process Des. Develop., Vol. 10, No. 3, 1971
is limited to use with ideal binary systems, whose relative volatility does not vary appreciably with concentration, since most packed columns do not provide for obtaining intermediate samples between the reboiler and condenser. Small diameter sieve or bubble cap plate columns could also be used, but problems of weeping and entrainment limit the operating range for such tests. Also, most plate columns cannot be operated a t 100% efficiency. Continuous fractionation, utilizing stages or plates, includes countercurrent and cocurrent contacting of vapor and liquid streams depending upon the particular flow regime on the plate. A plate or stage designed to give true cocurrent contact between phases would be an ideal design for a vapor-liquid equilibrium unit, since it would be impossible to obtain more than one theoretical stage of separation from one actual plate. True equilibrium was attained (Herring, 1948) by employing cocurrent vapor-liquid contacting in a packed contactor. Based on Herring’s results, an investigation was made to determine the usefulness of cocurrent contacting of gas and liquid phases for the design of a multistage vaporliquid equilibrium unit. The main difference between the present unit and the one described by Herring is the direction of flow of the two phases during contact. Herring’s unit employed downward flow of the two phases through the contactor, whereas in the present unit it is upward. Experimental
Flow Pattern. To carry out this investigation of cocurrent contacting to obtain vapor-liquid equilibrium data, a unit consisting of four vertically stacked stages was designed and constructed. The design permits cocurrent contacting of the vapor and liquid streams traversing each stage, although the overall phase flow is countercurrent. T h a t is, the vapor ascends from the reboiler through the individual stages to the condenser, while the flow of the liquid is descending through the stages from the condenser to the reboiler.
To provide cocurrent contacting of the vapor and liquid phases on each stage, i t is necessary to drop the liquid reflux a height of two stages instead of one, as is done in conventional bubble-cap or sieve plate distillation columns. Referring to Figure 1, liquid reflux from the condenser flows into a downcomer which transports the liquid reflux to the bottom of the contacting section located in stage n - 1. Vapors emanating from the top of the contacting section located in stage n - 1 enter the bottom of the contactor in stage n - 1 by means of four %-in. diam vapor ports. The pressure in the disengaging section causes the vapor and liquid phases to rise together through the annular contacting section, thus providing intimate contact between the two phases. The effluent streams from the contactor in stage n - 1 enter stage n where the vapor and liquid are disengaged. The vapor ascends to the condenser while the liquid enters a downcomer and flows to the bottom of the contacting section in stage n 2. The above flow pattern applies to each stage in the unit . Construction. Four cocurrent contacting sections, five liquid conduits, a reboiler, and a condenser are arranged in the form of a vertical column. The cocurrent contacting sections and the downcomers can be varied in size from one to two in. i.d. and % to % in. o.d., respectively. The stages are separated by a %- x 10%-in. diam cold rolled-steel plate. The vapor and liquid disengaging section of each stage consists of a six-in. length of six-in. i.d. borosilicate glass pipe, which also houses a contacting section and a downcomer. The height of each stage is thus set a t six in. The contacting section consists of an annulus formed by inserting the downcomer into the center of a borosilicate glass tube. The bottom of this contacting section is positioned in a cup fabricated from 12-gage tin-cladded steel which during operation contains liquid from the downcomer. Vapors enter the bottom of the contactor through % in. diam vapor ports drilled through the borosilicate glass tube wall. The overall length of the annular contacting section is 6% in., but only six in. of it is utilized for the contacting of the streams since the vapor ports are located %-in. from the bottom. The downcomer is constructed of borosilicate glass tubing, 12% in. in length. The top of it is flush with the upper side of the steel plate, thus practically limiting the liquid holdup to that in the contactor and downcomer. The bottom of each downcomer extends % in. below the bottom edge of the vapor ports, thus creating a liquid seal which prohibits vapors from entering the downcomer. Liquid samples are removed from each stage through a %-in. diam hole drilled horizontally into each steel plate. The sample collects in a small cup, xe-in. in diam and %-in. deep, which was machined into the upper side of each steel plate. Similarly, vapor samples were obtained through holes drilled horizontally into the plate and connected to a %-in. diam x %-in. deep hole drilled from the bottom. The four stages are enclosed in an insulated box, heated by circulated air which enters at the bottom and leaves at the top of the unit. Operation. The total volume of liquid required to operate the unit satisfactorily ranged from 12 to 20 liters,
V A P O R TO COLDENSER
,112" S T E t L PLATt
6" I.D. PYRhX P I P E 6" HIGB
S I X E R TU
DISEYGAGING CHAMBER VAPOR PORTS
CUP FOR BOTIOM OF MIXER T U B E
LEGS HOLD c
STILLPOT
Figure 1. Four-plate vapor-liquid equilibrium unit
depending on the boilup rate, which varied from 8 to 30 liters per hr and from 13 to 80 liters per hr for the %- and %-in. annular contactors, respectively. Liquid holdup per stage ranged from 220 to 280 ml and from 340 to 400 ml for the %- and %-in. contactors, respectively, with the holdup decreasing as the boilup rate was increased. At minimum boilup rates of 8 and 13 liters per hr for the x - and %-in. contactors, respectively, the total volume of liquid on any one stage is replaced by liquid from the next higher stage every two to three min. I t is important that the temperature of the air circulated around the four stages approximates that of the liquid boiling on each stage. For a higher air temperature, liquid splashing on the walls of the disengaging sections could be vaporized, resulting in a somewhat lower stage efficiency. For a lower temperature, partial condensation of vapors can occur, leading to a higher stage efficiency. I n actual operation, the air temperature entering the bottom of the box was maintained within =i=40Fof the temperature on the first stage and decreased slightly as it traveled through the box. Sufficient air was circulated to maintain a =t4OF difference between the heated air and the temperature of the boiling liquid on any given stage in the unit. Preliminary studies revealed that equilibrium conditions were established within 30 min after the air temperature was properly adjusted, but two hr of operation was allowed. In testing the unit, samples were removed very rapidly from each stage, beginning a t the bottom. The liquid sample was withdrawn first, followed by that of the vapor, into cooled glass vials. Ind. Eng. Chem. Process Des. Develop., Vol. 10, No. 3, 1971 381
I
~
Y
~ ~~~
~
~~~
~
Both contactors were first tested without packing, using the system 2,2,4-trimethylpentane-toluene.Typical test data are presented in Table I. The maximum boilup rate for the %-in. annular contactor was ahout 5000 grams per hr per cm', or 35 liters of liquid per hr. At this rate an overall efficiency of 7590 over the four stages was attained. While the maximum boilup rate for the %-in. annular contactor was
~
~
~
Table I. Efficiency-Throughput Data Using Unpacked Contactors
Test system: 2,2,4-trimethylpentane-toluene Compositions, mol % toluene Plate 1, liquid
Run no.
Plate 4, vopor
Throughput'
G per 1. per hr
hr per Cm
Ovecoll Relative volotilNtp
column efficiency,
R.M.S.
Yo
1.339 1.334 1.349 1.354 1.360
84.5 81.5 77.0 76.2 77.5
1.400 1.404 1.412 1.415 1.422 1.419 1.426 1.427 1.431
94.2 91.0 87.3 86.0 81.7 83.7 79.2 79.0 76.3
%-In. annular contactor 15 18 14 13 20
57.4 56.0 58.0 58.6 59.5
33.4 33.2 35.5 36.0 36.2
7.8 9.0 22.1 27.6 22.6
1370 1590 3900 4850 3980
%-In. annular contactor 1
2 3 4 5 6 7
8 9
-..-
68.0 68.2 68.5 69.0 69.0 69.0 69.2 69.1 69.1
37.5 37.5 39.4 40.3 41.3 40.8 42.4 42.1 42.9
".
15.1 21.6 27.5 36.7 47.8 43.0 53.5 60.5 78.0
535 770 980 1710
1550 1920 2180 2820
', across the condenser. ' A root mean square (R.M.S.) relative volatility was calculated from the individual values corresponding t o the liquid compositions on each plate. yI'"y6.1~y'
' 6 7 8 9 Figure 2. Metex screen packing
382 Ind. Eng. Chem. Process Der. Develop.,Vot. 10, No. 3, 1971
~~~. ...........
relationship, with efficiencies approaching 95% a t the lower boilup rates and leveling off a t ahout 75% as the maximum is reached. It was noted visually in both contactors that a t boilup
Results and Discussion
6 7 8 9
~
~
It is a loosely woven mesh of 0.0136-in. diam wire and was rolled into the shape of a cylinder to fill the annular contacting section. The test systems used were: 2,2,4-trimethylpentanetoluene, methylcyclohexane-toluene, and cyclohexane-nheptane. For the first system, the equilibrium data of Gelus et al. (1949) and Prengle and Palm (1957) were utilized by drawing a smooth curve through the points. This correlation resulted in a change of the relative volatility from 1.11 to 1.77 a t respective compositions of 10 and 90 mol % toluene in the liquid. The data of Kirk (1946) were utilized for the system of methylcyclohexanetoluene, in which the relative volatility varies from 1.10 to 1.59 as the composition of methylcyclohexane in the liquid changes from 91.6 t o 5.8 mol %. According to Myers (1957), the system cyclohexane-n-heptane exhibits a constant relative volatility of 1.65 over the concentration range from 5 to 95 mol % cyclohexane in the liquid. Two annular contactors were investigated for the throughput-efficiency relationships: a %-in. annulus, formed by a downcomer of %-in. 0.d. glass tuhe inside a one-in. i d . borosilicate glass tube; and a %-in. annulus, formed by a downcomer of K-in. o.d. glass tube inside a two-in. i.d. borosilicate glass tube.
y.ll*
l"y"
ba*Q.Iya"Iy
llylll
rates exceeding approximately 2000 grams per hr per cm' some channeling of the vapor stream occurs, which, in effect, reduces the intimate contact area between the vapor and liquid phases. This led to the decision to utilize Metex screen packing in the contacting space. Typical data for both the %- and 3/1-in. annular packed contactors are presented in Table I1 and plotted in Figure 4 for the three test mixtures. I t is quite evident from the data in Figure 4 that the maximum throughput of the smaller contactor was not markedly decreased by the use of the Metex screen packing, since throughputs approaching 5000 grams per hr per cm-! were realized. A similar throughput would be expected for the larger contactor. Efficiencies of 100% were attained on both contactors at boilup rates of approximately 1000 grams per hr per cm2. At higher rates, the efficiencies decreased to approximately 85% a t a maximum throughput of 5000 grams per hr per cm' for the %-in. contactor. Even though the maximum boilup for the %-in. contactor was not attained, it is believed that 5000 grams per hr per cm', or 120 liters per hr, can be reached and an efficiency of 85% realized.
Table I I . Efficiency-Throughput Data Using Packed Cpntactors Compositions, mol
Plate 1 ,
liquid
Yo
Plate 4, vopor
Throughput"
G per hr per cm'
1. per hr
Relotive volatility
Overall column
R.M.s.~ efficiency, %
Contactor: %-in. annular Packing: 18 in.' Metex screen Test system: 2,2,4-trimethylpentane-toluene Toluene 56.0 57.3 57.3 58.i 57.9
31.0 32.i 32.8 34.0 33.8
7.0 15.7 19.5 25.6 28.3
1240 2790 3460 4580 5020
1.305 1.320 1.320 1.333 1.330
97.5 91.5 91.0 88.5 86.5
Test system: methylcyclohexane-toluene 61.6 62.1 61.8 61.1 62.1
36.0 36.9 37.2 37.4 38.3
10.8 15.0 19.0 22.4 27.5
2090 2880 3600 4320 5320
1.310 1.321 1.320 1.317 1.323
95.0 92.7 90.7 87.7 86.7
Test system: cyclohexane-n-heptane Cyclohexane
\LEGEND
6a 90
27.2 33.4 31.8 32.8 34.1
I/4-INCH ANNULAR
0-
3 I 4 - I N C H ANNULAR
6.8 12.0 14.2 18.0 19.6
1.65 1.65 1.65 1.65 1.65
100.0 95.5 94.0 92.0 91.5
Test system: 2,2,4-trimethylpentane-toluene Toluene
T E S T SYSTEM
: TOLUENE
2,2,4-TRIMETHYLPENTANE
70 0
1200 2270 2610 3320 3600
Contactor: Vi -in. annular Packing: 90 in.' Metex screen
CONTACTOR
0-
73.4 77.2 75.4 75.4 76.3
1000 2000 3000 THROUGHPUT- GMSIHOURI
71.3 69.7 71.6 70.0 70.7 70.7 70.7 70.7 71.5
5000
4000
37.9 37.3 38.0 37.9 38.4 39.0 39.2 39.4 40.1
13.0 23.1 18.8 36.5 44.5 52.4 60.7 66.7 83.3
465 820 670 1288 1590 1880 2170 2380 2980
1.419 1.404 1.422 1.410 1.413 1.414 1.417 1.418 1.429
100.0 99.7 100.0 98.2 97.7 96.0 95 .O 93.7 94.2
Test system: methylcyclohexane-toluene
C M ~
Toluene Figure 3. Efficiency-throughput relationship for unpacked contactors
l
l
o
o
~
I
I
e-b
I i
50.7 50.7 50.7 59.1 59.1 59.2 59.5 51.4
27.4 29.4 29.4 34.5 34.8 35.2 36.0 31.3
12.5 19.0 25.2 30.4 41.7 51.7 60.0 70.0
516 775 1030 1240 1700 2110 2450 2760
1.250 1.250 1.250 1.287 1.288 1.289 1.288 1.254
101.0 101.0 101.0 99.5 98.5 97.0 94.5 93.0
Test system: cyclohexane-n-heptane Cyclohexane 27.2 27.6 30.9 30.2 30.2 30.2
W
CONTACTOR
0-
V
-
CYCLOHEXANE: NORMAL HEPTANE CYCLOHEXANE: NORMAL HEPTANE
3/4-INCH ANNULAR I14-INCH ANNULAR 3/4-INCH ANNULAR 114-INCH ANNULAR
lo00
2000
THROUGHPUT
3000
4000
- GMS I HOUR I CM2
so00
73.7 73.7 76.7 75.4 75.0 74.7
20.0 25.3 35.9 44.9 56.5 64.0
760 960 1385 1740 2190 2480
1.65 1.65 1.65 1.65 1.65 1.65
100.5 100.2 99.5 97.5 96.8 96 .0
" T h e throughput of the unit was calculated from a heat balance across the condenser. * A root mean square (R.M.S.) relative volatility was calculated from the individual values corresponding to the liquid compositions on each plate.
Figure 4. Efficiency-throughput relationship for packed contactors Ind. Eng. Chem. Process Des. Develop., Vol. 10, No. 3, 1971
383
Conclusions
From these studies, it can be concluded that cocurrent contacting of vapor and liquid streams is a very satisfactory method to utilize for the design of a multistage vapor-liquid equilibrium unit. A contactor, six in. in length, with an annular space of %- or Y4-in., packed with Metex screen, will be 100% efficient a t throughputs up to 1000 grams per hr per cm2, which corresponds, respectively, to 6 and 28 liters of liquid per hr. T o realize this efficiency a t higher throughputs, a longer packed contactor can be utilized. Nornenclatu re
a b E.F. n
= more volatile component = less volatile component = enrichment factor = (yo/yb)(xb/x,) = number of theoretical stages
y = concentration in vapor, mol % x = concentration in liquid, mol % O( = relative volatility Literature Cited
Fenske, M. R., Znd. Eng. Chem., 24, 482 (1932). Gelus, Edward, Marple, Stanley, Miller, M. E., ibid., 41, 1757 (1949). Herring, J. D., MS Thesis, Pennsylvania State University, University Park, Pa., 1948. Kirk, Norman, PhD Thesis, Pennsylvania State University, University Park, Pa., 1946. McCormick, R. H., Barton, P., Fenske, M. R., AZChE J . , 8, 365 (1962). Myers, H. S., Petrol. Refiner, 3, 175 (1957). Prengle, H. W., Palm, G. F., Znd. Eng. Chem., 49, 1769 (1957).
RECEIVED for review August 14, 1970 ACCEPTED March 15, 1971
Selection of Metal Oxides for Removing SO2 From Flue Gas Philip S. Lowell', Klaus Schwitzgebel', and Terry B. Parsons' Tracor, Znc., Austin, Tex. 78721 Karl J. Sladek' Department of Chemical Engineering, University of Texas at Austin, Austin, Tex. 78712 Oxides of 47 elements were evaluated for use as sorbents for removing SO2 from flue gas, in processes based upon thermal regeneration of sorbent. Thermodynamic analysis plus literature data were used to evaluate reaction paths for each oxide. The most important sorbent requirements involve sulfite and sulfate decomposition behavior. Sulfites were grouped according to the extent of disproportionation to sulfide and sulfate during decomposition. Sulfate groups correspond to simple decomposition, decomposition to oxysulfates, and decomposition with valence change. Examination of all the relevant data indicated that oxides of AI, Bi, Ce, Co, Cr, Cu, Fe, Hf, Ni, Sn,
Th, Ti, V, U, Zn, and Zr are most promising. I t is widely recognized that the emission of sulfur oxides from fossil fuel combustion causes a serious atmospheric pollution problem. A variety of methods have been proposed for removing sulfur compounds either from fuels before combustion or from flue gas afterward (7, 36). Processes for treating flue gas include wet scrubbing, gasphase reaction to make a removable solid or liquid product, sorption by a nonregenerable solid, and sorption by a regenerable solid. A process of this last type, described recently by Newel1 ( 3 0 ) ,employs sorption of sulfur oxides on an alkalized alumina, which is an active form of NaAIOr, and regeneration with reducing gas to produce H B . The cost of producing reducing gas is a major operating cost item in this process.
' Present address, Radian Corp., Austin, Tex. 78758 ' To whom correspondence should be addressed. 384
Ind. Eng. Chem. Process Des. Develop., Vol. 10, No. 3, 1971
This paper covers dry metal oxide processes which produce SO2, SO3, or H2SOaas by-product. I t attacks the problem of selecting from 47 pure metal oxides and an additional group of binary metal oxides those which are most promising for a thermally regenerable process. The use of thermal, rather than chemical, regeneration should result in operating cost advantages. A flow diagram illustrating the type of process and the sorbent requirements is given in Figure 1. Feed flue gas containing SOz, 02,N2, H 2 0 , and C 0 2 enters the sorber, which operates at temperature T s , and is contacted with regenerated sorbent. For a coal-burning power plant, typical flue gas SO2 content would be 0.05 to 0.3%, and typical O2 would be 2.0 to 3.5%. A lower limit of 100°C is chosen for T s to minimize corrosion and to prevent plume droop and resulting localized pollution. The critical sorption requirement is that the exit SO1 content must
