pressure swing distillation process

346

Miller, K. J.; Noddings, C. R.; Nattkemper, R. C. Preventing Bed Fires in Carbon Adsorption Systems. &'roc.-APCA Annu. Meet. 1987, 80th, Vol. 3, 87/50.7, 1-11.

Naujokas, A. A. Preventing Carbon Bed Combustion Problems. Loss Prev., CEP Man. 1979,12,128-135.

Naujokas, A. A. Spontaneous Combustion of Carbon Beds. Plant/ Oper. h o g . 1985,4 (2), 120-126.

Takeuchi, Y.; Mizutani, M.; Ikeda, H. Prevention of Activated Car-

Ind. Eng. Chem. Res. 1992,31,346-357

bon Bed Ignition and Degradation During the Recovery of C y clohexanone. J. Chem. Eng. Jpn. 1990,23 (lo), 68-74.

Wildman, J. Practical Problems in Solvent Recovery Using Activated Carbon. Proc. Znt. Conf. Carbon 1988, 185-187.

Received for review February 4, 1991 Revised manuscript received August 2, 1991

Accepted August 15, 1991

A New Pressure-Swing-Distillation Process for Separating Homogeneous Azeotropic Mixtures

Jeffrey P. Knapp and Michael F. Doherty* Department of Chemical Engineering, Goessmann Laboratory, University of Massachusetts, Amherst, Massachusetts 01003

Mixtures in which the desired products lie in different distillation regions can be separated by a new pressure-swing process when one end of the distillation boundary is a pressure-sensitive azeotrope. Thus pressure-insensitive binary azeotropes can be separated using novel entrainers that form pressure-sensitive distillation boundaries. Because the columns already operate at different preasures, such sequences are readily thermally integrated. The separations of ethanol from water and acetone from methanol are used to demonstrate the new pressure-swing technique. These examples exhibit some interesting behavior such as (1) a region of multiplicity in the number of trays required to achieve the same separation at fixed reflux ratio, (2) a maximum reflux above which no feasible column exists, (3) a separation where the unexpected component is the distillate due to a reversal of the relative volatility as the pressure changes, and (4) a nonazeotropic separation that becomes easier as the pressure is increased.

Introduction Azeotropic mixtures are typically separated by either

homogeneous azeotropic distillation (includes extractive distillation), heterogeneous azeotropic distillation, distillation using salt effects, or pressure-swing distillation. Of these four methods, pressureswing distillation is the least studied. However, with a recent National Research Council report (King et al., 1987) declaring that energy conservation in nonideal mixture separations is a vital research area, pressure-swing distillation deserves a second look, since these sequences are readily thermally integrated.

Although the pressure sensitivity of azeotropes has been known since the 18609 (Roscoe and Dittmar, 1859; Roscoe, 1860, 1862), Lewis (1928) appears to be the first one to exploit this property to distill azeotropic mixtures. Since then pressure-swing distillation has been proposed to separate the ethanol-water azeotrope (Lewis, 1928; Black, 1980), a number of alcohol-ketone azeotropes including methanol-acetone, methanol-methyl ethyl ketone (MEK), ethanol-MEK, ethanol-methyl n-propyl ketone, 1- propanol-methyl n-propyl ketone, and 2-propanol-MEK (Britton et al., 1943), and the tetrahydrofuran (THF)- water azeotrope (Abu-Eishah and Luyben, 1985; Chang and Shih, 1989). In fact, pressureswing distillation is the method often used by industry to separate THF and water, and Chang and Shih (1989) showed that there is no cost advantage in switching to heterogeneous azeotropic distillation with n-pentane. Table I lists some other azeotropes that may also be amenable to this technique.

Since the conventional pressure-swing-distillation process is restricted to binary mixtures with pressure-sensitive azeotropes, it is of limited usefulness. In this paper, we will show how pressure-swing distillation can be extended to separate the much broader class of pressure-insensitive azeotropes by using an entirely new class of entrainers, and how it can be used to separate some multicomponent mixtures containing distillation boundaries that lie be-

0808-5885192 f 2631-0346$03.00/0

Table I. Selected Examples of Pressure-Sensitive Binary Azeotropes

components rep carbon dioxideethylene hydrochloric acid-water water-formic acid water-acetonitrile water-acrylic acid water-acetone water-propylene oxide water-methyl acetate water-propionic acid water-2-methoxyethanol water-2-butanone (methyl ethyl ketone [MEK]) water-tetrahydrofuran (THF) carbon tetrachloride-ethanol carbon tetrachloride-ethyl acetate carbon tetrachloride-benzene methanol-acetone methanol-2-butanone (MEK) methanol-methyl propyl ketone methanol-methyl acetate methanol-ethyl acetate methanol-benzene methylamine-trimethylamine ethanol-dioxane ethanol-benzene ethanol-heptane dimethylamine-trimethylamine 2-propanol-benzene propanol-benzene propanol-cyclohexane 2-butanone (MEK)-benzene 2-butanone (MEK)-cyclohexane isobutyl alcohol-benzene benzene-cyclohexane benzene-hexane phenol-butyl acetate aniline-octane

' (1) Horsley, 1973. (2) Britton et al., 1947.

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1, 2 1, 2 2 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

tween the desired products without the addition of an entrainer. A second intent of this paper is to point out

0 1992 American Chemical Society

Ind. Eng. Chem. Res., Vol. 31, No. 1, 1992 347

Entrainer Material Balance Line

. .. . . . . . . Recycle Balance Line Distillation Boundary at cited pressure

--- pressure

I Azeotrope

Pure B F F1 Pure A Comporltlon

(a)

I , Pressure , P :, , 1 Pressure pP , D2

I

Pure A Pure B

(b) Figure 1. Conventional pressure-swing distillation. (a) Presswe- sensitive, minimum-boiling azeotrope. (b) Column sequence.

some of the typical behavior that can occur during the distillation of azeotropic mixtures. These effects are not limited to pressure-swing distillation nor are they patho- logical artifacta of the particular mixtures discussed below.

Conventional Pressure-Swing Distillation Before explaining the new pressure-swing-distillation

process, it is helpful to recall how the traditional pressure-swing-distillation process works. The conventional procedure requires two columns that operate at different pressures. For a binary mixture forming a homogeneous minimum-boiling azeotrope (Figure l), the fresh feed (F) is mixed with the recycled stream from the second column to form the feed stream (F,) to the first column, which operates at pressure P1. Since F1 lies to the right of the azeotrope at pressure P,, pure A is removed as the bottom product (B,) and a mixture near the azeotropic composition at pressure P1 is the distillate (D,). Stream D, is changed to pressure P2 and fed to the next column (as stream F2). Since F2 now lies to the left of the azeotropic composition at pressure P2 (Figure la), the other pure component, 8, can be recovered in the bottom stream (B2) and a near azeotropic mixture becomes the distillate (D2) for recycling to the fmt column. An analogous procedure is used for binary homogeneous maximum-boiling mixhues (see Van Winkle, 1967).

Clearly this process cannot be used when the azeotropic composition does not change with pressure.

New Pressure-Swing-Distillation Process If the binary azeotrope to be separated is pressure-in-

sensitive, then an entrainer must be used. If the selected

Pure A Pure B

I Pressure P1 I Pressurepp

I

Entrainer Make-up I D2

I I

Pure A

Figure 2. Pressure-swing distillation with an entrainer. (a) Material balance lines. (b) Column sequence.

entrainer forms no boundaries dividing the pure components into different distillation regions, then the methods of homogeneous azeotropic distillation can be used (Foucher et al., 1991; Doherty and Caldarola, 1985). Naturally, the columns in such sequences may be operated at different pressures in order to exploit thermal integration opportunities (see Knapp and Doherty (1990)). If the entrainer does form a boundary that divides the pure components into different distillation regions (henceforth called a feasibility boundary to distinguish it from other boundariea which may be present in the mixture but which do not constrain the separation), then it is not normally possible to find a feasible sequence in which all columns operate at the same pressure (Caldarola, 1983). [An ex- ception to this rule occurs when the boundary is very curved (see Figures 10-12 in Levy et al. (1985)). Levy (1985) and Van Dongen (1983) discuss boundary crossing in detail and they conclude that it is rarely cost compe- tative and normally not practical.] However, if the feasibility boundary moves with pressure (even though it does not move at the pressure-insensitive binary azeotrope), then ternary pressure-swing sequences can be devised. Thus, it is sometimes possible to overcome the restrictions imposed by the presence of distillation boundaries.

For illustration, consider the simplest case where the entrainer forms only one pressure-sensitive azeotrope and one distillation boundary between the desired products (Figure 2). As the pressure is changed from PI to P2 the boundary moves, pivoting around the pressure-insensitive azeotropic composition (Figure 2a). By operating two columns at different pressures, with the streams connecting

348 Ind. Eng. Chem. Res., Vol. 31, No. 1, 1992

Entrainer Material Balance Line ---

I Azeotrope

Pure A Pure 0 (a)

Entrainer

Pure A Pure 0 (b)

Figure 3. Estimating optimal stream compositions. (a) Preferred stream compositions. (b) Less desirable stream compositions.

the columns lying in the region between the positions of the boundary at the two pressures, it is possible to separate the original pressure-insensitive azeotrope into its pure components A and B. This is shown by the material balance lines in Figure 2a for the sequence in Figure 2b.

The distillate from column 2 (D2) is recycled and mixed with the fresh feed (F) to form the feed to the first column (FJ. In the fiirst column (operating at pressure Pl), pure A is removed as the bottom product (B,) and a stream whose composition lies near the boundary at pressure P1 is the distillate (D1). Stream D1 then becomes the feed (F,) to the second column (operating at pressure Pa), where it is separated into pure B and a distillate (D2) near the boundary at pressure P2 for recycle. Notice that no material balance line crwes the boundary at its own pressure and that the streams D1 = F2 and D2 interconnecting the columns lie between the two positions of the boundary, as stated. The resulting column sequence is the same as that used in the conventional process to separate binary pressure-sensitive azeotropes (Figure lb).

Optimal stream compositions tend to minimize the entrainer recycle rate and maximize the product flow rates. A quick estimate of the optimal compositions can be made by applying the lever-arm principle to the material balance lines superimposed on a skeleton residuecurve map, which shows only the positions of the distillation boundaries (Figure 3). The ratio of the recycle stream (DJ to the original feed (F) is the ratio of the lengths of the line segments F-F1 to F1-D2. Comparing Figure 3a and Figure 3b, it is clear that the preferred case (Figure 3a) has a smaller ratio than the less desirable case (Figure 3b) and

--

hence has a smaller recycle ratio and total cost (smaller flow rates mean smaller columns and lower utilities). The ratio of the product to nonproduct flow rates is given by the ratio of Fl-D1 to Fl-Bl in column 1 and Dl-D2 to D1-B2 in column 2 (recall D1 = F2). Clearly, these ratios are larger in the more desirable case, Figure 3a, than in Figure 3b. Thus, the preferred stream compositions in pressure-swing distillation minimize the recycle-to-fresh- feed ratio by locating the recycle stream composition in the residue-curve-map region where the boundaries shift most with changing pressure, and simultaneously maximize the product flow rates by positioning the nonproduct streams close to their respective boundaries.

Many systems contain multiple distillation boundaries. As long as only one of these boundaries separates the desired products (i.e., there is only one feasibility boundary), the column configuration remains essentially the same as shown in Figure 2b. However, when multiple boundaries lie between the pure components, pressure-swing distillation becomes rather complex, requires more than two columns, and will only work if all the feasibility boundaries shift with varying pressure. Therefore, it is very unlikely that it will ever be advantageous to use pressureswing distillation when more than one boundary lies between the desired products.

Suitable pressure-swing entrainers cause distillation boundaries to move in one of three ways.

(i) The entrainer forms no new azeotropes at atmospheric pressure, but when the pressure is increased (decreased), new azeotrope(s) appear which move rapidly with changing pressure, e.g., the dehydration of ethanol with acetone.

(ii) The entrainer forms one or more new azeotropes whose composition(s) change rapidly with pressure.

(iii) The entrainer forms one or more new azeotropes at atmospheric pressure, but they disappear as the pressure is increased (decreased), e.g., the separation of acetone and methanol using methyl ethyl ketone.

This suggests a modification of the entrainer selection guidelines given by Doherty and Caldarola (1985) which can lead to the use of novel entrainers. Potential entrainers can be screened by the following four-step procedure:

(i) First check if the composition of the azeotrope in the original mixture changes appreciably with pressure. If so, one option is to separate the mixture without adding an entrainer by using conventional pressureswing distillation.

(ii) From the b o i i points of the two pure components, the potential entrainer, and all azeotropes, make a sketch of the residue-curve map (see Foucher et al. (1991) for details).

(iii) If the resulting map is one of the seven most attractive for homogeneous azeotropic distillation (see Figure 7 of Foucher et al. (1991)), then the candidate entrainer deserves further study. This screening process has been automated by Foucher et al. (1991). Doherty and Calda- rola (1985) list 28 other maps that are potentially useful for homogeneous azeotropic distillation.

(iv) If the r e s iduewe map has one or more boundaries between the required products, check whether these boundaries can be shifted by varying the pressure. If they move sufficiently, then this is a viable entrainer for pressure-swing distillation.

Once a group of feasible entrainers has been found, an economic comparison of the resulting sequences can be used to discriminate between them.

Because the columns necessarily operate at different pressures, thermal integration of preasure-swing-distillation sequences is straightforward. However, choosing the best

- -


Pure B Material Balance Line

. . . . . . . . . Recycle Balance Line )-I Distillation Boundary

at cited pressure Azeotrope

---

8 2

Pure A Pressure Pressure Pl p2

(a)

Pure C

I I I I

I I I I I I I I

0 2 I D3 Dl I

-

column pressure is not trivial for mixtures with distillation boundaries. Because the boundaries move with pressure, the feasible distillate and/or bottom compositions also change with pressure. Therefore, selecting the pressure for the high-pressure column of the pressure-swing sequence requires calculating the position of the boundaries at each possible pressure, determining the likely near-optimum, nonproduct stream compositions, performing bubble-point and dew-point calculation on the bottoms and distillate to be matched, checking whether the resulting condenser temperature is high enough to supply heat to the reboiler of the low-pressure column, and as- suring that the boundary movement is sufficient for the pressure-swing process to work. In practice, the column pressures that optimize the nonintegrated pressure-swing sequence are usually sufficient for thermal integration. Since the pressure level of each column is fixed relative to the other columns in the sequence in order to make the separation work, only one thermal integration alternative exists and multiple effects will not be useful.

Pressure-Swing Distillation of Ternary Mixtures In the previous section, the objective was to separate a

binary azeotropic mixture into ita two pure components. Another use for this generalized pressure-swing technique is the separation of ternary mixtures containing distillation boundaries into the three constituent pure components without the addition of a fourth component (entrainer).

The presence of a boundary in a ternary mixture means that the three pure components cannot be separated without the addition of an entrainer. However, if the boundary (boundaries) can be shifted by changing the pressure, a pressure-swing sequence like that shown in Figure 4 can be used. The precise sequence required depends on the position of the feed as well as the number and location of the boundaries. As the number of boundaries increases, this pressure-swing process becomes increasingly more complex, requires more columns, typically has higher recycle ratios (by the lever-arm rule), and thus becomes less attractive.

The minimum number of columns required for a given separation (but not necessarily the optimum, especially for dilute feeds-see Knight (1986)) can be calculated from

Nco, = Np + N* - 1 (1)

where N , , is the minimum number of columns required; Np is the number of pure component products leaving the distillation sequence; NB is equal to unity for extractive distillation and equal to the number of boundaries crossed for pressure-swing distillation. This does not include boundaries that disappear as the pressure changes.

For nonazeotropic systems, this reduces to the familiar Nm1 = Nc - 1, where Nc is the number of pure components.

Equation 1 tells us that the minimum number of columns required to separate ternary mixtures with boundaries by pressure-swing distillation is greater than or equal to the minimum number required for extractive distillation, i.e., three columns (assuming that an entrainer can be found which forms no boundaries with the original ternary mixture). The equality holds when only one boundary needs to be crossed. Consequently, it is unlikely that the pressure-swing process will be advantageous when more than one boundary is present.

Design of Pressure-Swing Sequences To illustrate the design principles for the new pres-

sure-swing process, we have selected two commercially important mixtures: (1) ethanol-water and (2) acetone-

F t h * , I F3 I I I I I I I I I I

- 3

(b)

F m 4. Pressureswing distillation for separating ternary mixtures into the three constituent pure components. (a) Material balance lines. (b) Corresponding column sequence.

methanol. Although only pressures greater than atmospheric are used in these examples, subatmospheric pressures may also be of interest.

Example 1. First consider the ethanol-water system. With today’s emphasis on waste minimization and envi- ronmental concerns, there is a growing trend to seek suitable entrainers among the chemicals already present in the plant. Taking this approach, we have chosen acetone, a common industrial solvent, as the entrainer. The first step is to determine the number and position of all the azeotropes and distillation boundaries in the system over a moderate pressure range by calculating residue- curve maps at several pressures using the methods proposed by Doherty and Perkins (1978). Unfortunately, there is an extreme shortage of experimental vapor-liquid equilibrium (VLE) data above ambient pressure. In the absence of such data, we use vapor-liquid equilibrium models to extrapolate atmospheric VLE data up to 10-atm pressure and compare the predicted azeotropic compositions with those available in the literature. In performing this extrapolation, we assume that the vapor phase remains a mixture of perfect gases.

Of the four liquid solution models tried (Wilson, Van Lam, and one- and two-parameter Margules), the Wilson model is the best. (See the Appendix for the parameter values used.) The predicted ethanol-water azeotropic compoeitions closely follow the experimental ones of otsuki and Williams (1953) and Kleinert (1933). Two additional binary azeotropes and one ternary azeotrope appear as the pressure is increased. The acetone-water azeotrope is


Acetone Acetone

Water

Acetone

Ethanol

Pressure = 5.0 ATM Azeotrope

Water X l Ethanol

(b)

Pmssum 10.0 ATM m Azeotrope

Water

Acetone

Ethanol

- 1 ATM - - - - - 5 ATY 0000 10ATM

Azeotrope

0.4-

0.2-

0.0 0.2 0.4 0.6 0.8 1 .o Water X l Ethanol

(d)

Figure 5. Residue-curve maps for the ethanol-water-acetone mixture. (a) 1.0 atm. (b) 5.0 atm. (c) 10.0 atm. (d) Summary of the distillation boundary positions at the three pressures.

predicted to appear at a lower pressure than that cited by Horsley (1973), but the agreement improves as the pressure increases. The ethanol-acetone azeotrope is also predicted to appear at a lower pressure, contains more ethanol, and is more pressure sensitive than found experimentally by Campbell et al. (1987). (See Knapp (1991) for a more detailed comparison.) All the models predict a ternary azeotrope at elevated pressure. While there are no experimental data to confirm this, it is in qualitative agreement with Horsley’s (1973) data for the system with the next higher ketone, ethanol-water-MEK, and is therefore most likely correct. The predicted azeotropic compositions and distillation boundaries at 1, 5, and 10 atm are shown in Figure 5. This information can be summarized conveniently by superimposing the boundaries at each pressure onto a single diagram (Figure 5d).

For comparison with other designs from the literature, approximately 40.1 million kg/year of 99.8 mol % ethanol will be produced from a 4.2 mol % ethanol fermentation broth. Sizing and costing correlations are taken from Doherty and Malone (1991) using fourth-quarter 1986 prices. As a starting point, we chose the high-pressure column to operate at the highest pressure considered (10 atm) where the boundary movement is the greatest. Figure 6 shows the boundary movement between 1 and 10 atm. Though the movement of the feasibility boundary is not very great, it is sufficient for pressure-swing distillation to work.

For dilute feeds it is often economical to add a precon- centrating column. However, since the entrainer, acetone, is the lightest component in the mixture, the heavy component, water, can be removed from the bottom of the first


Acetone

Water X I Ethanol Figure 6. Position of the distillation boundaries at 1 and 10 atm for the ethanol-water-acetone mixture. The shaded region is where the streams interconnecting the columns must lie. The optimal material balance lines are also shown.

column, eliminating the need for a preconcentrator. In contrast, an extractive distillation sequence requires a preconcentrator before the extractive column to remove most of the water, because extractive entrainers are the heaviest component in the extractive distillation process; otherwise all the water would have to be boiled up twice before leaving in the distillate of the entrainer recovery column. (See Knight (1986) for an economic comparison of extractive distillation sequences with and without a preconcentrator.)

Because the low-pressure column of the pressure-swing sequence separates two streams of different composition (the fresh feed and the recycle stream), it seems likely that a doublefeed column would be preferred over a single-feed column. However, because the fresh feed flow rate is much larger than the flow rate of the recycle stream (apply the lever-arm rule to the material balance line F-F,-D2 on Figure 6), the overall feed is nearly coincident with the fresh feed and there is virtually no difference between single- and double-feed columns. Thus, two single-feed columns will be used. (In this example either a single- or double-feed column can be used because the products, ethanol and water, are both nodes on the residue-curve map (Figure 5). In contrast, for the extractive distillation residue-curve map (see example 2 below or the mixtures discussed in Knapp and Doherty (1990)) the pure component products are saddles on the residue-curve map and the desired separations are only feasible in a double-feed column.)

The columns are designed and optimized using the method of Knight and Doherty (1989). The dominant optimization variables are the stream compositions connecting the columns (because they control the recycle-to- fresh-feed ratio and the product flow rates), and the en- thalpic state of the feed to the first column. Notice that the optimal location of the nonproduct streams (D1 and D2) is near their respective boundaries (see Figure 6) as discussed above. The optimized sequence (Figure 6, Figure 7, and Table 11) has a total annualized cost (TAC) of $4.03 million/year and an energy requirement of 15 850 kJ/kg of ethanol. Of course it is very unlikely that the sequence would ever be operated without taking advantage of the

Table 11. Optimal Column Pressure for Thermally Integrated Pressure-Swing Distillation of Ethanol and Water with Acetone

sp energy TAC, IO6 consumption,

$/year kJ/kg

high-pressure column at 10 atm 3.47 7500

high-pressure column at 8 atm 3.69 8870 high-pressure column at 9 atm 3.55 8070

2664.3 2492.6

FRACTION Water

. Acetone

Figure 7. Flowsheet for the optimized, nonthermally integrated, pressure-swing sequence for the ethanol-water-acetone mixture.

c. w

63.5%

'0°+

lim -!? 139.3%

n

Steam

Figure 8. Thermally integrated, pressure-swing sequence for ethanol-water-acetone at 10 atm.

energy integration opportunities that arise from having two adjacent columns at different pressures. Thermal integration (see Figure 8 and Table 11) reduces the TAC by 14% to $3.47 million/year and cuts the energy consumption by 53% to 7500 kJ/kg. The optimal pressure for the high-pressure column is between 9 and 10 atm (Table 11). Comparing our pressure-swing results with the extractive distillation designs of Knight and Doherty (1989) and Knapp and Doherty (1990), and the heterogeneous azeo-


120 1

0.0

90 - u)

b- L o - .c 0

70 - s z 60 -

50 -

40 -

,

mm 0 5 10 15 20

Reflux Ratio Figure 9. Multiple solutions and maximum reflux. For fixed product purities and feed composition, two solutions exist for reflux ratios between 11.467 and 18.977 (r-).

tropic distillation design of Ryan and Doherty (1989), we see (Table 111) that thermally integrated pressure-swing distillation of ethanol and water with acetone consumes less energy than heterogeneous azeotropic distillation with benzene and consumes about the same amount of energy as conventional extractive distillation, but costs more to build and operate on an annualized basis. (Black (1980) concluded that conventional pressure-swing distillation is not practical for this mixture.) However, on the basis of energy alone, none of these alternatives can compete with thermally integrated extractive distillation. Thus, despite its higher cost, we have shown that even the small amount of boundary movement that occurs between 1 and 10 atm in the ethanol-water-acetone system is sufficient for pressure-swing distillation to work, though the movement is not enough to make it economical.

The high-pressure column in the ethanol-wateracetone pressure-swing process exhibits some interesting behavior. As the reflux ratio is increased at fixed feed and product compositions, the column-section profiles move relative to each other in the composition triangle, and the number of column profile intersections (solutions) changes from zero, to one, to two, to one, and then back to zero, all for finite reflux (Figures 9 and 10). Thus, there is a range of reflux ratios where multiple solutions to the design problem exist (Figure 9). In this range, it is possible to design two columns operating at the same reflux and yielding the same product compositions, but with different numbers of trays. (Once a column is built, this multiplicity is no longer an issue because the number of trays has been fixed). These multiple solutions are different from those found by other researchers (e.g., Rovaglio and Doherty (1990), Venkataraman and Lucia (1988), Chavez et al. (19861, or Prokopakis and Seider (1983a,b)), who formulate the problem as a column simulation and not a column design. They find multiple sets of distillate and bottom

-- -..- A Distillate + Feed o Bottoms

Tray Composition

I ' \ Reflux - Ratio = 11.467

Acetone

i

-

-

1 .o

0.8

0.6

*2

0.4

0.2

First solution

I 0.0 0.2 0.4 0.6 0.8 1 .o

MoleFnc F n d Dlrtlllatr Bottom

Ethanol 0.5500 0.3231 0.9980 Water 0.0600 0.0894 0.0020 Acetone 0.3900 0.5875 1.0 lo'* \

First solution

Acetone 1.0

Laand

0.8

0.6

XZ

0.4

0.2

0.0

First solution

I I 0.2 0.4 0.6 0.8 1 .o

Water

Acetone 1 .o

Ethanol Xl

(b)

0.8

0.6

x2

0.4

0.2

0.0

Water Xl Ethanol (C)

nl. I 1 .. - .* n I - L.., 1 . r igure iu. cIoiumn prorues snowing me mumpie soiuuons. \a) Minimum reflux for the first solution. (b) Minimum reflux for the second solution. (c) Maximum reflux. No feasible column exists above this reflux ratio.


Acetone 56.1 % 1 .O^

Table 111. Comparison of Pressure-Swing Distillation and Extiactive Dietillation for Distilling Ethanol and Water

sp energy TAC, 108 consumption,

$/year k J / k EtOH pressure-swing distillation 3.47 7500

high pressure = 10 atm thermally integrated

extractive distillation with ethylene glycol

feed preheatera added

optimized, nonthermally integrated extractive distillation with ethylene glycol

distillation with ethylene glycol

azeotropic distillation with benzene

Knight and Doherty (1989)o 2.51 8920

Knapp and Doherty (1990) 2.01 7670

Knapp and Doherty (1990) 2.36 2700 thermally integrated extractive

Ryan and Doherty (1989)O 2.05 10380 optimized heterogeneous

feed preheater added

"Corrected cost. (1990) for more details.

See the appendix of Knapp and Doherty

compositions for fixed feed composition, reflux ratio, and number of trays, while we find multiple number of trays for fixed feed, reflux ratio, and product compition. This new type of multiplicity was also observed by Van Dongen (1983) and is not limited to pressure-swing distillations.

It is commonly believed that all reflux ratios above the minimum yield a feasible column, with the number of trays decreasing as the reflux ratio increases. However, as il- lustrated in Figures 9 and 10, this is not always true for nonideal mixtures. For fixed feed and product compositions (i.e., the design problem) a finite maximum reflux ratio (r-) can exist beyond which no feasible column is possible. In fact, the existence of a maximum reflux ratio in the design problem is not uncommon. As described in a forthcoming article (Knapp and Doherty, 1991), every extractive distillation exhibits a maximum reflux ratio. For the ethanol-water-acetone example presented here, r-

These two phenomena can have important implications for the design and control of distillation columns. If a column is operated close to the maximum reflux, either because the true minimum reflux was not found at the design stage (a difficult task with the commonly used process simulators) or because the reflux was increased to achieve a product specification, then it is possible to exceed rmm, making the desired separation impossible. In the region of multiple solutions, choosing the wrong solution needlessly increases the cost of the column. Example 2. Next consider the separation of acetone

and methanol. The azeotrope in this binary mixture is sufficiently pressure sensitive that the conventional pressure-swing technique is one option. Methyl ethyl ketone (MEK) appears to be an excellent pressure-swing entrainer for this system. The VLE data for this system are modeled using the Van Laar equation with the parameters given in the Appendix. These parameters accu- rately represent the available experimental VLE and azeotropic data over the pressure range 1-10 atm. (See Knapp (1991) for more details.) At 1 atm MEK and methanol form an azeotrope containing 85 mol % methanol and the ternary mixture has a distillation boundary running from the methanol-acetone azeotrope to the methanokMEK azeotrope, putting the desired acetone and methanol products into different distillation regions

= 18.977.

Pressure L 1 .O atm

79.0% 64.2OC 64.5OC

(a)

Acetone 103.4%

0.0 0.2 0.4 0.6 0.8 1 .o

130.3'C 704.4OC MEK X1 Methanol

(b) Figure 11. Residue-curve maps for the acetone-methanol-MEK mixture a t (a) 1.0 and (b) 4.0 atm. Notice the disappearance of the methanol-MEK azeotrope and the distillation boundary.

(Figure lla). Therefore, this mixture cannot be separated by homogeneous azeotropic (extractive) distillation at atmospheric pressure. However, by about 4 atm the methanol-MEK azeotrope has moved to the methanol vertex, ceasing to exist and causing the distillation boundary to disappear. The resulting residue-curve map (Figure l lb) contains only the acetone-methanol azeotrope and is the map for extractive distillation. Thus one option is to distill this mixture by extractive distillation with all of the columns operating above 4 atm. A better choice is to use pressure-swing distillation to purify acetone in an atmospheric column and methanol in a column operating above 4 atm.

For comparison with existing extractive distillation and traditional pressure-swing-distillation designs, approximately 204.4 million kg/year of a saturated-liquid, equi- molar, binary, acetonemethanol mixture will be separated into acetone and methanol products of 99.5 mol % purity


MEK

I

Figure 12. Schematic diagram of the methanol-acetone-MEK pressure-swing-distillation sequence.

Acetone

Legend Azeotrope

A Distillate + Feed o Bottoms # Tray Composition

Pressure 10.00 atrn 0.60

'..k 0.20 \ 0.00 0.20 0.40 0.60 0.80 1 .oo

MEK x, Methanol

Figure 13. Column profile for the first column of the methanol- acetone-MEK pressure-swing-distillation sequence. Notice the tangent pinch in the rectifying section.

with fractional recoveries of 99.9%. The columns are restricted to operate between 1 and 10 atm. The sequence is designed, costed, and optimized as described in the first example.

This example also contains interesting effects. From looking at the residue-curve maps (Figure ll), we would expect to remove the lowest boiling pure component, acetone, in the distillate of the first column operating at 1 atm and obtain essentially a binary methanol-MEK stream as the bottom product. Then, in the second column operating above 4 atm, this mixture would be separated into the methanol product and the MEK recycle stream. However, it is impossible to bring acetone out of the top of the first column at any pressure between 1 and 10 atm. Instead, the binary methanol-MEK azeotrope comes off as the distillate in the pressure range 1-4 atm. At higher pressures, pure methanol comes off the top of the first column. Methanol changes from being the intermediate- boiling pure component'to the low-boiling pure component between 4 and 5 atm due to the different slopes of the vapor pressure curves for acetone and methanol. (See Knapp and Doherty (1991) for a method to a priori predict which component will be the distillate and for examples.) Consequently, the first column of the pressure-swing sequence (Figures 12 and Figure 13) operates above 4 atm and produces methanol as the distillate. The bottom

1 .o

0.5

0.0 0.0 0.5 1 .o

XI

Mole Fraction of Methanol Figure 14. Binary vapor-liquid equilibrium curves for methanol- MEK at 1, 4, and 10 atm showing how the azeotrope becomes a tangent pinch that is still present at 10 atm.

stream from this first column is essentially a binary acetone-MEK mixture, which is separated into pure acetone and pure MEK for recycle in the second column which operates at atmospheric pressure.

Now that we know where each product wil l be removed, the sequence can be designed and optimized. The separation of methanol from acetone and MEK turns out to be a very difficult separation. The optimization procedure drives the pressure of the high-pressure column above the maximum pressure considered, so we set the column pressure at the upper bound, 10 atm. The minimum feed ratio (upper feed flow rate divided by lower feed flow rate) for the separation is 0.79. That is over five and a half times higher than for extractive distillation of the same acetonemethanol mixture with water and indicates that the double-feed column will be considerably larger, more expensive, and consume more energy than its counterpart in the extractive distillation process. The minimum reflux ratio (rmh) is on the order of 4.2, and the separation requires about 110 theoretical trays at 1.2rmin (95 trays at l.5rmin) with a feed ratio of 1.7.

The reason for the very difficult separation is a severe tangen pinch in the rectifying section. The extremely large number of trays in the rectifying section near the methanol distillate in Figure 13 indicates that there is a tangent pinch at 10 atm. Since the rectifying section essentially lies on the methanol-MEK edge of the composition triangle, we can visualize its behavior by examining the binary y-x diagram for methanol and MEK. Figure 14 shows VLE data for methanol and MEK between 1 and 10 atm. At atmospheric pressure there is an azeotrope containing about 85 mol 9% methanol. As the pressure increases to 4 atm, the amount of methanol in the azeotrope increases until it reaches 100% methanol (i.e., the azeotrope disappears), and the azeotrope is replaced by a severe tangent pinch. (The relationship between the appearance or disappearance of an azeotrope and tangent pinches is explained in Knapp (1991)J As the pressure is further increased from 4 to 10 atm, the tangent pinch becomes leas severe and the VLE curve moves farther away from the 45O line, indicating that the separation becomes easier (i.e., lower rmin and fewer trays required). This is


Table IV. Comparison of Conventional and New Pressure-Swing Distillation with Extractive Distillation for Separating Acetone and Methanol

sp energy TAC, lo6 consumption, $/year kJ/kg EtOH

Extractive Distillation with Water feed ratio = 0.55

sequence optimized, nonthermally integrated 2.70 3190

thermally integrated sequence 2.72 1950

optimized, nonthermally integrated 2.79 3120


feed ratio = 1.0

sequence

Pressure-Swing Distillation with MEK optimized, nonthermally integrated 6.97 7800

sequence thermally integrated sequence 5.83 3810

optimized, nonthermally integrated 3.54 5140


Conventional Pressure-Swing Distillation

sequence

Table V. Antoine Equation Constants component A B C

methanol 23.4832 -3634.01 -33.768 ethanol 23.5807 -3673.81 -46.681 acetone 21.3099 -2801.53 -42.875

water 23.2256 -3835.18 -45.343 MEK (2-butanone) 21.1480 -2899.47 -51.392

Table VI. Wilson Constants (AiJ for the Ethanol-Water-Acetone Mixture

component j component i ethanol water acetone

ethanol 1 0.1782 0.692 water 0.8966 1 0.492 acetone 0.726 0.066 1

in direct contrast to conventional wisdom which says distillation becomes more difficult as the column pressure increases. This rule of thumb is correct in most instances, but whenever an azeotrope disappears with increasing pressure, the azeotrope will be replaced by a tangent pinch, which becomes less pronounced as the pressure continues to be increased, resulting in an easier separation. (Simi- larly, distillation becomes continually more difficult and tangent pinches become more restrictive as the pressure is increased when azeotropes appear with increasing pressure (see example 2 of Knapp and Doherty (1990))J The fact that the separation becomes easier as the pressure increases accounts for the optimization procedure driving the pressure of the high-pressure column above 10 atm. Unfortunately, for the specified methanol product purity, the tangent pinch remains severe at 10 atm, resulting in a difficult and expensive separation. The final version of the sequence costa $6.97 million/year and consumes 7800 kJfkg of acetone. Thermal integration reduces the cost by 16% to $5.83 millionfyear and cuts the energy consumption by 51% to 3810 kJfkg. This is more than twice the cost and energy consumption of extractive distillation with water (see Table IV). The new pressure-swing-distillation process using MEK was also compared with traditional pressure-swing distillation (Figure 1) of methanol and acetone. Because of the tangent pinch problem discussed above, pressureswing distillation with MEK is also considerably more expensive than the conventional pressure-swing process (Table IV), though the energy con- sumptions of the thermally integrated versions are com-

Table VII. Van Laar Constants (Aij) for the Methanol-AcetoneMEK Mixture

component j component i methanol acetone MEK

methanol 0 196.0 242.20 acetone 182.0 0 15.20 MEK (2-butanone) 237.60 101.20 0

parable. This example demonstrates the need to consider more than just the amount of boundary movement for potential pressure-swing-distillation systems. One must also be aware of the presence of tangent pinches and the influence of pressure on their severity. Therefore, movement of distillation boundaries is a necessary, but not sufficient, condition for pressure-swing distillation to be practical.

Conclusions A pressure-swing process for separating azeotropic

mixtures has been developed. This new technique uses varying pressure to move distillation boundaries that lie between the desired products which otherwise would make the separation impossible. Both pressure-sensitive and pressure-insensitive binary azeotropes can thus be separated in a two-column, pressureswing sequence using novel entrainers that form pressure-sensitive distillation boundaries between the two pure components to be iso- lated. New entrainer selection guidelines which include potential pressure-swing entrainers are presented. The new distillation method was applied to the separation of ethanol and water, and acetone and methanol.

The new pressure-swing process can also be used to separate ternary mixtures containing distillation boundaries into the three pure components, without the addition of an entrainer (fourth component). However, such schemes are not of practical interest when more than one distillation boundary must be crossed.

Although only ternary mixtures were discussed, the new pressure-swing-distillation technique can be extended to multicomponent mixtures in an obvious way. For ternary mixtures the distillation boundary that must move with pressure is a line. For quaternary mixtures the feasibility boundary which must be pressure-sensitive is a two-dimensional surface. In general, for an N-component mixture the feasibility boundary will be an (N - 1)-dimensional surface.

Several phenomena were observed in the examples studied. First, in the high-pressure column of the ethanol-water-acetone system, for fixed compositions and reflux ratio there were two feasible column designs, each with a different number of trays. Second, in the same column, there was a maximum reflux ratio above which no feasible column design was possible. Third, in the acetone-methanol-MEK system, the relative volatility of acetone and methanol reverses and methanol becomes the overhead product. Lastly, this separation becomes easier as the column pressure is increased. All of these effects have been observed for other mixtures and other distillation methods and are discussed more thoroughly in Knapp (1991) and Knapp and Doherty (1991).

Acknowledgment

We acknowledge G. A. Caldarola, who did some prelim- inary work on this problem in his Master’s Thesis at the University of Massachusetts (1983). We are also grateful for the research support provided by the Link Foundation in the form of an Energy Fellowship for J. P. Knapp.


Nomenclature B = bottom stream D = distillate stream F = feed stream NB = unity for extractive distillation and equal to the number

Nc = number of pure components to be separated Ncol = minimum number of columns required for a given

Np .= number of pure component products leaving the dis-

P = pressure r,, = maximum reflux ratio rmin = minimum reflux ratio

Appendix Vapor-liquid equilibrium calculations were carried out

using an in-house thermodynamic-property software package and database. This appendix contains the model equations and physical property parameters used for the mixtures discussed in the text.

Vapor pressures are calculated using Antoine's equation:

of boundaries crossed for pressure-swing distillation

separation

tillation sequence

B In et = A + - T + C (A-1)

with the vapor pressure in pascals and the temperature in kelvin. Antoine constants are given in Table V.

Vapor-liquid phase equilibrium was calculated assuming that the vapor phase is a mixture of perfect gases, Le., by solving

subject to Cyi = 1. Activity coefficients were calculated using the model which most closely matched the available experimental data. The ethanol-water-acetone mixture was modeled using a nonstandard form of the Wilson equation

yip = xiYi(x,T) PYT) (A-2)

n n n

j = l i = l j = 1 In Yk = 1 - In (XAkjXj) - C(XiAik/CXjA;j) (A-3)

In this model Aii = 1 and ideality is implied by Aij = 1. Wilson interaction parameters are given in Table VI. The methanol-acetone-MEK mixture was modeled by the Van Laar equation:

(A-4) where zi is the effective volume fraction

(A-5)

If Aji/Aij = O/O, set Aji/Aij = 1. In this model Aii = 0 and ideality is implied by A , = 0. The Van Laar interaction parameters are given in Table VII.

67-64-1; MeOH, 67-56-1.

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Column Azeotropic Distillation System. Znd. Eng. Chem. Process Des. Dev. 1985,24, 132-140.

Black, C. Distillation Modeling of Ethanol Recovery and Dehydra- tion Processes for Ethanol and Gasohol. Chem. Eng. Prog. 1980,

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Caldarola, G. A. Synthesis of Distillation Trains for Multicomponent Azeotropic Mixtures. M.S. Thesis, The University of Massachu- setts a t Amherst, 1983.

Campbell, S. W.; Wikak, R. A.; Thodos, G. Vapor-Liquid Equilib- rium Measurements for the Ethanol-Acetone System at 372.7, 397.7, and 422.6 K. J. Chem. Eng. Data 1987,32,357-362.

Chang, T.; Shih, T. T. Development of an Azeotropic Distillation Scheme for Purification of Tetrahydrofuran. Fluid Phase Equi- lib. 1989, 52, 161-168.

Chavez, R. C.; Seader, J. D.; Wayburn, T. L. Multiple Steady-State Solutions for Interlinked Separation Systems. Znd. Eng. Chem. Fundam. 1986,25,566-576.

Doherty, M. F.; Perkins, J. D. On the Dynamics of Distillation Processes-I. The Simple Distillation of Multicomponent Non- Reacting, Homogeneous Liquid Mixtures. Chem. Eng. Sci. 1978,

Doherty, M. F.; Caldarola, G. A. Design and Synthesis of Homoge- neous Azeotropic Distillations. 3. The Sequencing of Columns for Azeotropic and Extractive Distillations. Znd. Eng. Chem. Fundam. 1985,24,474-485.

Doherty, M. F.; Malone, M. F. Multicomponent Fractionation Sys- tems. Manuscript in preparation, 1991.

Foucher, E. R.; Doherty, M. F.; Malone, M. F. Automatic Screening of Entrainers in Homogeneous Azeotropic Distillation. Znd. Eng. Chem. Res. 1991,30, 760-772.

Horsley, L. H. Azeotropic Data ZZfi Gould, R. F., Ed.; Advances in Chemistry 116; American Chemical Society: Washington, DC, 1973.

King, C. J.; et al. Separation & Purification-Critical Needs and Opportunities; National Research Council, National Academy Press: Washington, DC, 1987; pp 63-64.

Kleinert, T. Zur Kenntnis der Dampf-Fluessigkeitsgleichgewichte von Aethylalkohol-Wassergemischen bei Temperaturen von 120 bis 180°. Angew. Chem. 1933,46, 18-19.

Knapp, J. P. Exploiting Pressure Effects in the Distillation of Ho- mogeneous Azeotropic Mixtures. Ph.D. Dissertation, The Univ- ersity of Massachusetts a t Amherst, 1991.

Knapp, J. P.; Doherty, M. F. Thermal Integration of Homogeneous Azeotropic Distillation Sequences. AZChE J. 1990,36,969-984.

Knapp, J. P.; Doherty, M. F. Minimum Entrainer Flows and Max- imum Reflux Ratios in Double-Feed Extractive Distillation Col- umns: A Bifurcation-Theoretic Approach. Submitted for publi- cation in AZChE J. 1991.

Knight, J. R. Synthesis and Design of Homogeneous Azeotropic Distillation Sequences. Ph.D. Dissertation, The University of Massachusetts at Amherst, 1986.

Knight, J. R.; Doherty, M. F. Optimal Design and Synthesis of Ho- mogeneous Azeotropic Distillation Sequences. Znd. Eng. Chem. Res. 1989,223, 564-572.

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Levy, S. G.; Van Dongen, D. B.; Doherty, M. F. Design and Synthesis of Homogeneous Azeotropic Distillations. 2. Minimum Reflux Calculations for Nonideal and Azeotropic Columns. Znd. Eng. Chem. Fundam. 1985.24.463-474.

33, 281-301.

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Prokopakis, G. J.; Seider, W. D. Dynamic Simulation of Azeotropic Distillation. AZChE J. 198313, 29, 1017.

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ton-Like Methods. Comput. Chem. Eng. 1988,12,5H9.

GENERAL RESEARCH

An Investigation on the Reaction of Phosphoric Acid with Mica at Elevated Temperatures

Chandrika Varadachari* Department of Agricultural Chemistry & Soil Science, University of Calcutta, 35 B.C. Road, Calcutta 700 019, India

Various aspects of the reaction of phosphoric acid with muscovite mica at 250,300, and 350 "C were studied with a view to understanding the nature of such reactions, particularly (i) the reaction kinetics, (ii) the relation between muscovite dissolution and polymerization of phosphoric acid, (iii) the probable mechanism of reaction, and (iv) the nature of the residue. Solubilization of the K+ ion from muscovite was observed to be linearly dependent on the degree of dehydration of the system as well as the average chain length of the poly(phosphoric acid) formed. It is suggested that the breakdown of the complex muscovite structure is due to attack by OH- ions, which are produced when phosphoric acid polymerizes; oxide bonds are cleaved forming M-OH and M-O-P bonds, and the elimination of water from other P-OH groups results in polyphosphates. The reaction product consists of soluble and insoluble amorphous polyphosphates that form a coating over the core of unreacted mineral.

1. Introduction Phosphoric acid at elevated temperatures shows a re-

markable reactivity that is unique among the inorganic acids. Thus, at high temperatures, phosphoric acids and phosphates have been observed to react with quartz (Mellor, 1925), with silicates and glasses (Kingery, 1950; Ray, 19701, and with a large number of metals and oxides (Bailar et al., 1973; Thilo, 1962), including gold and platinum (Van Wazer, 1966). Such reactions of phosphates find use in the manufacture of phosphate glasses (Ohashi, 1964), phosphate-bonded refractories (Kingery, 1950; Mamykin et al., 1973), and micronutrient fertilizer glasses (Roberts, 1975). In spite of the varied application of the phosphate-silicate/oxide reactions, however, the nature, mechanism, and products of such reactions have neither been investigated in detail nor been established with certainty. Thus, although the nature of polymerization of phosphoric acid and phosphate melts (Van Wazer, 1966) is now fairly well understood, its reactions with silicates and refractory oxides are subject to a great deal of spec- ulation. There are indeed very few studies pertaining to the mechanism of such reactions. Reactions of glasses with phosphoric acid have been studied by Ray (1970) and Walters (1983). According to Ray (1970), the reactivity of the glass is a function of acid dehydration, whereas Walters (1983) attributed the solubilization to attack by polyphosphates produced during heating. Ohashi (1964) proposed that the reaction occurs due to the ability of PO4

* Present address: Polymer Science Unit, Indian Association for the Cultivation of Science, Jadavpur, Calcutta 700 032, India.

0SSS-5SS5/92/2631-0357$03.00/0

units to form copolymers with silicates and other elements having X04 tetrahedra. Workers who have extensively inveatigated the products of the reaction of phosphoric acid with phyllosilicates concluded that various crystalline phosphates of silicon and aluminium are produced by the replacement of SO4 tetrahedra of the silicate, by the corresponding tetrahedra of the acid; their nature depends on the temperature and the relative proportions of the constituents (Kingery, 1950; Lyon et al., 1966; Mamykin et al., 1973; Zamyatin et al., 1972). The existing knowledge is, however, quite inadequate for understanding exactly how and why stable polymeric structures, such as the silicates, are broken down so readily to their component units by hot phosphoric acid. Neither can it explain the unique reactivity of this acid and the cause of the unusual change in its properties on heating.

In this paper, the results of a comprehensive study on the high-temperature reaction of phosphoric acid with a silicate are presented. It is hoped that this investigation will provide a better understanding of the reactions of hot phosphoric acid with silicates and oxides in general and perhaps also provide a clue to the cause of ita phenomenal solubilizing power. The study has accordingly been dealt with in four parts, which may be broadly classified as (1) the reaction kinetics of muscovite dissolution by phosphoric acid, (2) kinetics of polycondensation of phosphoric acid, (3) deduction of the reaction mechanism, and (4) identification of the products of reaction.

2. Methodology 2.1. Studies on the Dissolution of Muscovite by

Phosphoric Acid. Large flakes of muscovite (from Gir-

0 1992 American Chemical Society

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