Preparation of tert-Amyl Alcohol in a Reactive Distillation Column. 2

Aug 1, 1997 - (tert-amyl alcohol or TAA) in a 5.3-cm column using a bale-type catalytic packing. Acetone ..... This test was carried out in order to d...
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Ind. Eng. Chem. Res. 1997, 36, 3845-3853

3845

Preparation of tert-Amyl Alcohol in a Reactive Distillation Column. 2. Experimental Demonstration and Simulation of Column Characteristics J. Castor Gonza´ lez,† Hoshang Subawalla, and James R. Fair* Separations Research Program, The University of Texas at Austin, 10100 Burnet Road, Austin, Texas 78758

Reactive distillation has been proposed as a means of enhancing the conversion for reactions whose progress is limited by chemical equilibrium. In the present work, reactive distillation tests were carried out for the hydration of isoamylene (2-methyl-2-butene) to 2-methyl-2-butanol (tert-amyl alcohol or TAA) in a 5.3-cm column using a bale-type catalytic packing. Acetone was used as a solvent to avoid formation of a second liquid phase and also to enhance the rate of reaction. The experimental results showed TAA yields approaching 100% when the distillate stream from the top of the column was essentially suppressed. This yield is more than double that possible thermodynamically in a separate reactor. It was also found that an excess of catalytic packing can generate operational problems. The column was simulated using an equilibrium stage model together with the kinetic expression developed for this reaction and reported in Part 1 of this paper. The simulation and the experimental results matched reasonably well. Model results indicate that the TAA yield can be greater than 95% (at a pressure of 240 kPa) when the weight hour space velocity is lower than 30 kg/(h‚kg). A parametric study elucidated the effect of varying the feed composition and key operating variables. Introduction The hydration reaction of 2-methyl-2-butene (isoamylene) to produce tert-amyl alcohol (TAA) was studied using a batch reactor and is reported in part 1 of this work (Gonza´lez and Fair, 1997) A kinetic expression that accounts for the reverse reaction was developed in terms of the activities of reactants and products. Equilibrium isoamylene conversion was limited to 30% at a 1.5 M ratio of water to isoamylene and a temperature of 60 °C. These results confirm the strong thermodynamic limitations that affect this reaction. The reactive distillation concept has been proposed as an alternative to obtain high product yields for reactions that are thermodynamically inhibited. It is especially useful for exothermic reactions, since the heat of reaction is efficiently used to supply the heat of vaporization for column boilup (Stadig, 1987). It is also useful when there is a significant difference in relative volatility between the product and unconverted key reactant. It is essential that the temperature at which the reaction occurs matches the distillation temperature at a practical operating pressure (Westerterp, 1992). The liquid-phase synthesis of TAA seems to comply with these desirable characteristics of a reactive distillation application; the reaction is moderately exothermic (26.5 kJ/mol), and the relative volatility between isoamylene and TAA is very high (20-30). The temperature range at which the reaction takes place (60-85 °C) can be obtained in a distillation column that operates at pressures between 100 and 400 kPa. Acetone was used as a solvent to prevent the formation of aqueous and organic liquid phases; it was found that the acetone also enhanced the basic reaction rate significantly (about 15fold). The general goal of the work was to find a reactive distillation strategy that maximizes the yield of TAA. * To whom correspondence should be addressed. E-mail: [email protected]. † Present address: Intevep, S.A., P.O. Box 76343, Caracas 1070A, Venezuela. S0888-5885(96)00808-1 CCC: $14.00

The reaction kinetics and chemical equilibrium aspects were described in part 1. The first objective of this second part was to determine the column operating conditions such that complete isoamylene conversion could be approached, thus minimizing the costs of separation and recycle of unreacted isoamylene. The second objective was to confirm the simulated results experimentally, using a continuous reactive column of small diameter (5 cm) equipped with suitable controls for demonstration. A complete description of the entire work may be found in the dissertation by Gonza´lez (1997). Previous Work Several alternatives have been proposed for locating catalyst particles inside a distillation column. Three different approaches have been used: (1) Place beds of catalyst in a tray column, either in downcomers (Haunschild, 1972) or above the trays (Paret, 1984; Quang et al., 1989). (2) Use random packing elements containing catalyst; the packings can be prepared by extrusion (Smith, 1980), impregnation of porous packings, or molding resins as random packings (Flato and Hoffmann, 1992). (3) Use ordered packing elements which contain catalyst particles between layers of supporting material. Elements of this type include the bale packing proposed by Smith (1980), as well as other arrangements (Adams, 1991; Hearn, 1993; Gelbein and Buchholz, 1991). The alternative of ordered packing has been used for industrial applications, specifically the packings described by Gelbein and Buchholz (1991) and by Smith (1980). Gelbein and Buchholz used a type of structured packing element in which the catalyst is sandwiched between two layers of metal gauze (“catalytic structured packing”). Smith proposed an arrangement in which the catalyst beads are placed in the pockets of a Fiberglas cloth belt, which is rolled up using a spacer of mesh knitted steel wire (“bale-type catalytic packing”). A sketch of the bale-type packing is shown in © 1997 American Chemical Society

3846 Ind. Eng. Chem. Res., Vol. 36, No. 9, 1997 Table 1. Binary Parameters for the UNIQUAC Activity Coefficient Modela

Figure 1. Sketches of the bale-type packing (Smith, 1980; used with permission).

Figure 1. DeGarmo et al. (1992) reported that the capacity and mass-transfer efficiency of the catalytic structured packing (Gelbein and Buchholz, 1991) was almost equivalent to that of a conventional structured packing of the same dimensions. They also stated that the reaction rate was about 75% of the value obtained in a continuous stirred tank reactor with the same amount of catalyst. The mass-transfer characteristics of the bale-type packing were evaluated by Zheng and Xu (1992). Three correlations were developed for calculating the masstransfer coefficients for the gas and liquid films at the gas-liquid interface and for the liquid film at the solidliquid interface. The HETP (height equivalent to a theoretical plate) values predicted with this correlation do not match the experimental HETP values measured in a 5-cm column under total reflux conditions (Gonza´lez, 1997). Subawalla et al. (1996) evaluated the mass-transfer efficiency and vapor flow capacity of a 5-cm prototype of the bale packing in total reflux tests with cyclohexane/ n-heptane and acetone/methyl ethyl ketone. They reported vapor and liquid flood loadings equivalent to a vapor F-factor value of about 1.8 (m/s)(kg/m3)1/2, at pressures of 138 and 241 kPa and under total reflux conditions. The experimental values of the HETP ranged between 45 and 58 cm for the system acetoneMEK in the preloading regime. The patent literature shows two configurations of a reactive distillation column used for hydrating isoamylene. Schleppinghoff et al. (1990) disclosed a process to prepare TAA over ion-exchange resins at low temperatures (about 40 °C). The laboratory equipment used was a Soxhlet device with resin inside; isoamylene conversions between 30% and 50% were reported. Smith and Arganbright (1991) described a process for the production of tertiary alcohols over acid resins using a catalytic distillation reactor. According to this patent, the catalytic packing is placed between the stripping and rectifying sections of the column. Olefins are fed at the bottom of the packed section, and stoichiometric water is supplied at the top. Some results of the tests performed in a 2.5-cm stainless steel column were given (temperatures between 60 and 77 °C), but conversions and/or yields were not reported (only TAA content in the bottoms stream). Hence, no clear conclusions could be drawn about the effectiveness of the scheme. Reactive Distillation Simulation Model The reactive distillation column used in the present work was simulated using the RADFRAC block of ASPEN PLUS. This module utilizes the equilibrium stage approach and allows for the inclusion of reactions.

binary pair

source of the data

Ai,j

Bi,j (K)

water-2m2b 2m2b-water water-TAA TAA-water water-acetone acetone-water 2m2b-acetone acetone-2m2b 2m2b-TAA TAA-2m2b TAA-acetone acetone-TAA

liq-liq equilibrium liq-liq equilibrium vapor-liq equilibrium vapor-liq equilibrium vapor-liq equilibrium vapor-liq equilibrium vapor-liq equilibrium vapor-liq equilibrium UNIFAC predictive model UNIFAC predictive model UNIFAC predictive model UNIFAC predictive model

0.00 0.00 0.00 0.00 -2.61 3.05 0.00 0.00 0.00 0.00 0.00 0.00

-747.26 -332.79 35.95 -386.44 924.15 -1380.05 -275.89 71.72 -287.22 110.37 -123.21 32.48

2m2b water TAA acetone

area parameter qi

vol parameter ri

3.22 1.40 4.28 2.34

3.55 0.92 4.60 2.57

a 2m2b: 2-methyl-2-butene. Aspen calculates parameter τ as i,j τi,j ) exp(Ai,j + Bi,j/T). Vapor-liquid and liquid-liquid equilibrium data were taken from Gmehling et al. (1977).

The kinetic expression is supplied in a separate subroutine. The reaction rate on each stage is calculated at the conditions of the liquid phase (temperature and activities). As in conventional distillation columns, each stage is described by equations representing the material balance, equilibrium, and heat balance. The material balance for the i component in the j stage is

Lj-1 xi, j-1 - Lj xi, j + Gj+1 yi, j+1 - Gj yi, j ) νirjwj (1) The rate of reaction in each stage is calculated using the following kinetic expression:

rj ) 4.9{[1.01 × 1014 exp(-8359/Tj)aIC5, jaW, j3.64 7.6 × 1015 exp(-10 448/Tj)aTAA] (1 + 26.2aW, j3.64)2} (2)

/

This expression is equivalent to that given in the first part of this work (Gonza´lez and Fair, 1997), but the activation energy of the reverse reaction is slightly different since the heat of reaction was calculated here using the gas-phase heats of formation of the pure compounds and then corrected using the heats of vaporization (Gonza´lez, 1997). The vapor-liquid equilibrium relations for each component are

Ki, j ) yi, j/xi, j )

γi, j P i,satj φi,satj φi, j Pj

(3)

The activity coefficients for the highly nonideal liquid phase were calculated using the UNIQUAC model. The interaction parameters for this model are shown in Table 1. They were regressed mainly from vapor-liquid and liquid-liquid equilibrium data. For binary pairs where data were not available, use was made of the UNIFAC predictive model. This model was found to predict equilibria reliably for those pairs for which data were available. Gas-phase fugacity coefficients were calculated using the Redlich-Kwong equation. Reid et al. (1987) provide details on the UNIQUAC model and the Redlich-Kwong equation of state. Two binary azeotropes are present in the column reaction mixture, which leads to the following order of volatility of species inside the column: acetone (25 mol

Ind. Eng. Chem. Res., Vol. 36, No. 9, 1997 3847 Table 2. Characteristics of the Bale Packing Used in the Reactive Distillation Tests diameter of bale (cloth only) diameter of bale (cloth + wire mesh) bale (cloth) length bale (cloth) width bale element length dry catalyst density type of catalyst specific surface area void fraction wall wipers

Figure 2. Effect of the distillate/feed flow ratio on the TAA yield.

%)-isoamylene azeotrope (bp 32.5 °C); 2-methyl-2butene (isoamylene) (bp 37.5 °C); acetone (bp 56.5 °C); water (25 mol %)-TAA azeotrope (bp 87.4 °C); water (bp 100 °C); tert-amyl alcohol (TAA) (bp 102 °C). This listing indicates that it should be possible to separate unreacted isoamylene and acetone at the top of the column, with TAA and excess water withdrawn at the bottom. Details about the algorithm used to solve the column equations have been given by Venkataraman et al. (1990). During the simulation runs, it was assumed that there was no pressure drop along the column and all information related to the feed stream was available, including flow, composition, thermal conditions, and injection point. A partial condenser was used in all simulations. The column operation was set by fixing the pressure, boil-up ratio, and ratio of distillate to feed. The following assumptions and considerations were also taken in the simulations of the TAA reactive column: (1) There is good contact between the liquid and the catalyst particles, which are fully wetted. The validity of this assumption depends on the conditions at which the column is operated; as the column approaches the flood point, the contact between the liquid and catalyst particles should improve. (2) Catalyst particles are isothermal at the temperature of the liquid phase in equilibrium with the gas. It has been calculated that the temperature difference between the center and the surface of the particle is less than 0.2 °C (Gonza´lez and Fair, 1997). (3) There is no resistance to mass transport between the bulk liquid and the catalyst particles. (4) There is no resistance to mass transport inside the pores or gel phase of the catalyst. The intraparticle mass-transfer evaluations described in part 1 of this work indicate that below 70 °C, this is a reasonable assumption (but at higher temperatures, the resistance in the pores increases). Calculations show that increasing the temperature from 70 to 80 °C decreases the catalyst effectiveness factor from ∼100% to 70-80%. (5) The only reaction taking place is the hydration of isoamylene. Reactive Column Configuration A preliminary set of simulation runs was performed to identify the appropriate column configuration for obtaining high yields of TAA. It was found that the most critical parameter to achieve this goal was the flow ratio of distillate to feed (see Figure 2). As this ratio increases, the nonreacted isoamylene escapes from the top of the column and the TAA yield is reduced. The

3.6 cm 4.6 cm 29.0 cm 16.6 cm 31.3 cm 107 g/L Amberlyst 15 170 m2/m3 0.76 m3/m3 3 per element

most convenient configuration is to operate the column under total reflux but at a finite boil-up ratio. The isoamylene cannot leave the column at the top and is forced to react to TAA or to leave unconverted at the bottom of the column. The elimination of the distillate stream is also justified since no nonreactive lighter compounds (diluents) were included in the feed; if light diluents were present, it would be necessary to maintain a distillate flow from the top of the column. Additional simulation runs were performed to ascertain the effect of nonreactive stages above and below the reactive packing. It was observed that two to three nonreactive equilibrium stages are needed below the catalytic packing (stripping section) in order to recycle the unreacted isoamylene back to the catalyst section. On the other hand, the addition of equilibrium stages above the reactive section does not improve the column performance. Water and solvent (acetone) should be fed above the catalytic packing to ensure their flow through the packing. On the other hand, the most convenient place to feed isoamylene is between the reactive and stripping sections of the column. However, the simulation showed that under conditions approaching total reflux, the point at which isoamylene is fed makes little difference as long as it is above the stripping section. Taking these findings into consideration, the experimental equipment was set up to operate under total reflux with a stripping section of about three equilibrium stages and without any additional packing above the catalyst section. For practical reasons, it was decided to have a single feed point, above the catalytic packing, recognizing that additional feed points would not have a significant effect. Experimental Work Catalytic Packing. A small-scale prototype (5-cm diameter and 31-cm length) of the bale-type catalytic packing (see Figure 1) loaded with Amberlyst 15 catalyst was used in this work. The properties of this prototype of the packing are given in Table 2. As previously mentioned, the efficiency and capacity of this packing were evaluated by Subawalla et al. (1996), at the conditions of interest for this work. Equipment. Figure 3 is a schematic representation of the equipment. The column has an internal diameter of 5.1 cm and a total length of 225 cm. The condenser is located at the top of the column and is serviced with chilled water at 10 °C. In operation, the feed components isoamylene, water, and acetone are blended to a target concentration in tank T1, which is kept at less than 15 °C by cooler E1. The membrane pump P1 sends the feed to the top of the column at controlled flow rates between 0.1 and 12 L/h. Preheater E-2 raises the feed temperature to about 55 °C. Between the bottom of the column and the reboiler is an enlarged section of 15-cm diameter loaded with 1 m

3848 Ind. Eng. Chem. Res., Vol. 36, No. 9, 1997

Figure 3. Schematic representation of the reactive distillation column.

of nonreactive random packing (IMTP-15), which provides the two to three equilibrium stages needed for the stripping section. A liquid distributor was placed at the top of this section to ensure good liquid distribution under the lower loadings in the larger diameter section. The reboiler is heated with saturated steam at 900 kPa. The flow of steam to the reboiler is automatically controlled and the temperature of the outgoing condensate is measured, permittimg manipulation of the reboiler duty. The pressure of the column is regulated using two automatic control valves: one supplies N2 when the pressure is lower than the set-point value, and the other releases vapor from the top when the pressure is higher. The reboiler level is controlled automatically by a valve located downstream of the product cooler. The internal temperature of the column is measured at the top (close to the feed point) and at the bottom (at the column diameter transition). The wall temperature of the 5-cm column is measured at the top and bottom locations. All the equipment is carefully insulated, so these wall temperatures should be very close to the corresponding internal temperature. All exposed parts are fabricated from Type 304 stainless steel. Analysis. Liquid samples from the feed, top of the column (top distributor tray), and bottom stream are taken periodically during the tests and are analyzed using a GOW-MAC 550 chromatograph equipped with

a thermal conductivity detector (TCD) and a 10 ft (1/8-in.-o.d.) stainless steel column filled with Carbopack-B (mesh size 80/120). Discussion of Experimental Findings Several tests were performed with the objective of demonstrating experimentally that it is possible to obtain high yields of TAA in a reactive distillation column. Table 3 shows the operating conditions and feed compositions for four of these tests. For tests RD-6 and RD-9, the column was loaded with 6 catalyst bales, containing a total of 454 g of dry Amberlyst-15 catalyst, packed to a height of 191 cm. For tests RD-10 and RD11, only 2 bales with 141 g of catalyst were used (64cm height) and were located at the top of the column. The remaining part of the smaller diameter column for tests RD-10 and RD-11 was filled with 137 cm of Sulzer BX gauze structured packing. The F factor reported for each test was calculated as

F ) uvFv0.5

(4)

The vapor superficial velocity was calculated based on the reboiler heat duty, and the vapor density was calculated at the bubble point of the reboiler product. The yield of TAA is defined as the moles of TAA

Ind. Eng. Chem. Res., Vol. 36, No. 9, 1997 3849 Table 3. Operating Conditions during the Reactive Distillation Tests test catalytic packing length, cm pressure, kPa feed flow rate, kg/h temp top catalyst packing, °C reboiler duty, kJ/h F factor, (m/s)(kg/m3)1/2 percent flood (based on F factor) boil-up ratio (mol vapor/mol bottom) WHSV, kg/(kg‚h) duration, h feed comp, mol % isoamylene water acetone

RD-6

RD-9

RD-10

RD-11

191 145 0.88 65 7574 0.98 54 14.4 1.9 10.4

191 241 2.20 82 8173 0.84 47 6.4 4.8 4.25

64 241 2.22 82 8445 0.90 50 6.9 15.8 5.25

64 241 2.23 81 8621 0.90 50 6.9 15.8 12.5

3.21 5.65 91.14

2.93 7.04 90.03

2.62 6.45 90.93

4.67 12.31 83.02

Figure 4. Yield of products during test RD-6.

Figure 5. Yield of products during tests RD-9 and RD-10.

recovered in the product per mole of isoamylene in the feed, assuming 100% recovery of products. Test RD-6. Figure 4 shows the TAA yield during test RD-6 performed at 145 kPa. The yield for this test was very low (10-15%) during the entire run period. Also, a byproduct was detected in the bottoms stream. This material was identified by mass spectrometry to be mesityl oxide (4-methyl-3-penten-2-one) and is the product of the aldol reaction between two molecules of acetone followed by condensation to release water. The yield of this byproduct is defined as the moles in the bottom product stream per mole of acetone fed to the column. In test RD-6, the yield of mesityl oxide was between 2.7% and 4%. The low TAA yield in this test may be explained by two factors: an excessive amount of catalyst in the column and high temperatures during start-up with acetone as the reboiled medium. Compositions and temperatures in the reactive section should be such that they preferentially aid the forward reaction. For RD6, the top portion of the reactive section produced TAA. However, the bottom portion of the reactive section showed a low concentration of isoamylene, at a temperature where equilibrium favors the dissociation of TAA. Hence, the excess catalyst in the bottom portion effectively lowered the yield since it destroyed the TAA produced in the top section. Simulation results confirmed that only the top third of the catalyst packing actually promoted TAA formation, whereas the rest of the catalyst bed decomposed most of the TAA formed. Thus, the combination of increased temperature together with a 6-fold decrease in isoamylene concentration caused a decrease in net yield of TAA. Sundmacher and Hoffmann (1996) observed a similar phenomenon in their rate-based column simulation of methyl tertbutyl ether synthesis.

The second important factor that limited TAA yield in RD-6 was the start-up procedure. Initially the column was loaded with acetone and heated under total reflux conditions without fresh feed. At these conditions, the column temperature reached higher values than during normal operation at steady state. When the feed stream contacted the hot metal wall, much of the isoamylene was vaporized quickly, forcing it to leave at the top of the column through the pressure control valve. Tests RD-9 and RD-10. The yields for these tests are shown in Figure 5. Tests RD-9 and RD-10 were conducted at the same operating conditions except for the amount of catalyst used. The TAA yield during test RD-10 (only 2 bales) reached values of about 50%, whereas in RD-9 (6 bales) this yield was always lower than 5%. Further, the yield of byproduct mesityl oxide was reduced from 3-4% to 1%, about as expected considering the 69% reduction in the amount of catalyst used. Figure 5 clearly demonstrates the importance of using the correct amount of catalyst in a reactive distillation column. A larger amount than necessary (overdesign) may result in an unacceptably low yield of desired product and a high yield of byproducts. It was obvious that steady state was not achieved during test RD-10 since the TAA yield increased with time, but the yield trends during the unsteady-state period clearly indicate the importance of the amount of catalyst in the column. Test RD-11. This test was carried out in order to determine the steady-state yields of TAA and mesityl oxide byproduct. The operating conditions were similar to those for RD-10, but the concentrations of water and isoamylene in the feed were higher. Figure 6 shows that the TAA yield reached steady-state values of around 100%, implying close-to-full conversion of isoamylene to

3850 Ind. Eng. Chem. Res., Vol. 36, No. 9, 1997

Figure 6. Yield of product during test RD-11.

Figure 8. Simulation model profiles at the conditions of RD-11.

Column Simulation

Figure 7. Isoamylene content at the column top in tests RD-9, RD-10, and RD-11.

alcohol. A deviation of about (20% in the steady-state yield of TAA is noticed in Figure 6, mainly resulting from the error associated with measuring the flow rate of bottom product. On the other hand, the yield of byproduct was reduced to 0.3%. This reduction in the byproduct yield, in relation to RD-10, is thought to result from a lower liquid-phase concentration of acetone in the catalytic zone (higher steady-state concentrations of water and isoamylenes). This test clearly demonstrates that very high conversion of isoamylene to TAA is possible in a reactive distillation column operating at total reflux and with the appropriate amount of catalyst. Comparison of Tests. Figure 7 shows the liquid mole fraction of isoamylene in the distributor tray (just above the catalytic zone) during tests RD-9, RD-10, and RD-11. Note that in RD-11, a steady concentration of isoamylene in the top of the column was reached after about 6 h. The high concentration of isoamylene at the top during test RD-9, as compared to that of RD-10, also supports the hypothesis that the bottom section of the column decomposes the TAA formed in the upper part, leading to the discharge of isoamylene at the top of the column. RD-11 shows a high concentration of isoamylene at the top since the isoamylene feed concentration in this test was higher than in previous tests. The bench-scale data generated in this work may be affected by the characteristics of small-diameter columns, mainly wall flow. Subawalla et al. (1996) show the important effects that the wall exerts on the vaporliquid mass-transfer efficiency and capacity of the packing in the 5-cm column. Reactive distillation tests at the pilot scale (20-40-cm column diameter) are needed, especially for low surface area packings such as the catalytic bales.

Experimental vs Modeling Results. Before evaluating the effects of the operating conditions in the reactive column, the equilibrium stage model was used to simulate the best experimental test, RD-11 (RD-9 and RD-10 did not reach steady-state conditions). Considering the available information on HETP values of the catalytic and noncatalytic packings, a total of 11 nonreactive stages were taken for modeling purposes, while only 1 reactive stage was considered (third stage from the top). Figure 8 shows the concentration and temperature profiles inside the column according to the simulation model. The available experimental data from RD-11 (temperature and concentrations) are also plotted in this figure. Note that the TAA yield is very high (94.5 mol %), as in the experimental case, while the values of temperature and concentration agree reasonably well with the model-predicted curves. The top of the column is rich in isoamylene, while the bottom is practically free of isoamylene. The water composition profile is flat along the column, while the TAA concentration tends to be higher in the bottom section. This figure also indicates that the stripping capacity is greater than required since there is practically no change in temperature and composition between stages 6 and 10. Hence, a simulation with fewer equilibrium stages would give essentially the same results. Sensitivity Analysis. An analysis was conducted to determine the effect of changes in the column operating conditions. The base case simulation study used conditions similar to those of test RD-11; however, the number of nonreactive stages was reduced from 11 to 6. From the base case, a sensitivity study was performed, each parameter being changed over a wide range of WHSV, which is calculated by dividing the total mass flow rate of feed to the column (kg/h) by the mass of dry catalyst inside the bale packing (kg). Table 4 shows the reactive column conditions for the base case as well as the range of values at which each parameter was evaluated. Number of Reactive Stages. The effect of the number of reactive stages and WHSV on yield of TAA is shown in Figure 9. For WHSV values smaller than 30 kg/(h‚kg), the TAA yield remains practically constant at 94-96%, independent of the number of reactive stages. For this range of WHSV, the amount of catalyst is large enough to hydrate most of the isoamylene that is fed to the column. The nonreactive section in the bottom of the column recycles the unconverted isoamylene back to the top, thus producing a bottom stream essentially free of isoamylene. For values of

Ind. Eng. Chem. Res., Vol. 36, No. 9, 1997 3851 Table 4. Operating Conditions Used for the Simulation of the TAA Reactive Column parameter pressure, kPa WHSV, kg/(kg‚h) percent of F factor at flooding no. of reactive stages isoamylene in feed, mol % mole feed ratio water to isoamylene boil-up ratio,a mol/mol no. of nonreactive stages feed point distillate-to-feed ratio

base case

range evaluated

241 15 50 1 5 2.5

138-345 5-90 50-95 1-3 5-7.5 1.5-2.5

7 1 partial condenser 4 stripping stages 1 reboiler stage 1 0.004

1.17-7 unchanged unchanged unchanged

a

Boil-up ratio adjusted to satisfy the desired values of WHSV, flooding, and distillate-to-feed ratio.

Figure 10. Effect of vapor flow in the column (flooding approach) on TAA yield.

Figure 11. Effect of the column pressure on TAA yield. Figure 9. Effect of WHSV and number of reactive stages on TAA yield.

WHSV higher than 30 kg/(h‚kg), the TAA yield drops very quickly for both cases (one and three reactive stages). In this region, isoamylene begins to appear in the bottom product in larger amounts as WHSV increases. The TAA generation rate (in kmol/h) is practically constant for WHSV values higher than 60 kg/ (h‚kg); hence, any additional increment in the feed rate will cause more isoamylene to appear in the bottom stream, further reducing the observed TAA yield. An important observation from Figure 9 is that for the same value of WHSV, the column with a single reactive stage gives a higher TAA yield than the column with three reactive stages, especially for WHSV > 30 kg/(h‚kg). This means that for the same feed flow rate and amount of catalyst inside the column, a less masstransfer efficient packing (high value of HETP) that accommodates more catalyst in one equilibrium stage will give a higher TAA yield than a more efficient packing with a lower HETP value. This can be explained by considering that for a single stage, all of the catalyst is in contact with an isoamylene-rich liquid phase that produces a very high rate of TAA, while for the case of three reactive stages, the last stage is in contact with an isoamylene-stripped liquid, which makes this equilibrium stage inefficient from the reaction point of view. For low values of WHSV, this effect is negligible, as there is an excess of catalyst in the column, while at high values of WHSV, where all the catalyst is needed, the lower stages are inefficient owing to the low isoamylene content in the liquid phase. This simulation result (better yield using a lower efficiency device) cannot be extrapolated to other reactions or configurations, and it would be advisable to confirm it experi-

mentally or with a more rigorous column model such as a rate-based reactive distillation model instead of the equilibrium stage model used here. Vapor Flow Rate. Figure 10 shows the effect of the vapor rate inside the column, expressed in terms of approach to flood and considering that the packing floods at an F-factor value of 1.8 (m/s)(kg/m3)1/2. It appears that in the high TAA yield zone, the higher the vapor flow, the higher the yield of TAA. High vapor flow increases the recycle of unconverted isoamylene from the bottom to the top of the column. At high values of WHSV, this additional recycle helps very little, since the catalyst is operating at maximum capacity and the accumulation of isoamylene actually reduces the temperature of the reactive stage. Column Pressure. The effect of pressure on TAA yield is shown in Figure 11. A pressure increase raises the column temperature; thus, a pressure of 345 kPa (50 psia) leads to high TAA yields over the whole range of WHSV values studied (WHSV < 90 kg/(h‚kg). Operation at a lower pressure (138 kPa or 20 psia) has the opposite effect. Feed Composition. The effect of feed composition on TAA yield is shown in Figures 12 and 13. Water is essentially an inhibitor of the reaction (see eq 2). For this reason, it was expected that the reduction of the molar ratio water:isoamylene (at constant isoamylene content) would increase the TAA yield, since the rate of reaction would be higher. Figure 12 shows that this is not the case. The reduction of the ratio from 2.5 to 1.5 actually reduces the TAA yield. At the lowest ratio, the column fills with isoamylene and the column temperature falls 7-10 °C. This counteracts the positive

3852 Ind. Eng. Chem. Res., Vol. 36, No. 9, 1997

Figure 12. Effect of the water/isoamylene ratio on TAA yield.

facilitated when carried out in an acetone medium that enhances the rate of reaction and avoids the formation of two liquid phases. The key issue to achieve conversions of isoamylene close to extinction is to minimize the distillate flow at the top (reduce removal of unconverted isoamylene) and also to load the column with moderate amounts of catalytic packing so as not to promote TAA decomposition. Excess catalyst in the column is ineffective because the bottom portion of the catalyst bed can actually reverse the hydration reaction. A byproduct from acetone (mesityl oxide) was observed in these tests, but its yield was only 0.3 mol % under conditions where the yield of TAA approached 100%. The simulation of the column using an equilibrium stage model showed the existence of two regions: one of very high TAA yields, obtained at low WHSV, and the other where there is not enough catalyst to react all the isoamylene fed to the column. The high TAA yield is achieved at low space velocities (WHSV< 30 kg/ (h‚kg)) for a feed composition of 5 mol % isoamylene, high column temperature in the reaction zone (more than 70 °C at a pressure of 240 kPa or higher), and moderate concentrations of isoamylene in the reactive section of the column (10-30 mol %). At very high concentrations of isoamylene, the column may actually have a lower internal temperature, leading to low TAA yields. Acknowledgment

Figure 13. Effect of the isoamylene concentration in the feed on TAA yield.

effects of the reduction of water mole fraction and increase of the isoamylene concentration in the liquid phase. Figure 13 shows that increasing the content of isoamylene in the feed, up to 7.5 mol %, increases the TAA yield almost 2% at low WHSV values. At WHSV ) 28 kg/(h‚kg) and 7.5 mol % of isoamylene, the model gives two different solutions depending on the initial estimate of the values of temperature and composition given to the program. One of the solutions produces high yields of TAA, and the column operates at high temperature with moderate concentrations of isoamylene. The other solution produces low TAA yields (