A Novel Reactive Distillation Process for the Indirect Hydration of

Ind. Eng. Chem. Res. 2008, 47, 9581–9587

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A Novel Reactive Distillation Process for the Indirect Hydration of Cyclohexene to Cyclohexanol Using a Reactive Entrainer Frank Steyer,† Hannsjo¨rg Freund,† and Kai Sundmacher*,†,‡ Max Planck Institute for Dynamics of Complex Technical Systems, Sandtorstrasse 1, 39106 Magdeburg, Germany, and Process Systems Engineering, Otto-Von-Guericke UniVersity Magdeburg, UniVersita¨tsplatz 1, 39106 Magdeburg, Germany

In the conventional process for cyclohexanol production, large amounts of energy are consumed and a considerable quantity of side products is formed. In addition, the process is inherently unsafe. The alternative process of cyclohexene direct hydration requires large amounts of catalyst to overcome kinetic limitations. This publication shows the feasibility of a new route from cyclohexene to cyclohexanol by means of reactive distillation using formic acid as a reactive entrainer. This allows overcoming kinetic limitations with moderate amounts of catalyst and makes large-scale cyclohexanol production by reactive distillation an interesting alternative. The suggested coupled reactive distillation process allows producing cyclohexanol in an inherently safe and energetically advantageous way without incurring significant amounts of side products. Introduction Cyclohexanol is an intermediate for nylon production and is thus produced on the megaton per year scale worldwide.1 The conventional production process starts with benzene, which is first hydrogenated to cyclohexane and then partially oxidized to a mixture containing cyclohexanol and cyclohexanone. The partial oxidation step is operated at low conversions of 7-12% to keep up the selectivity at values of 70-90%.2 Nevertheless, a considerable amount of side products is formed. The low conversion and the presence of side products necessitate very large recycle streams that have to be distilled at substantial energy costs to separate the desired products, the side products, and the unreacted cyclohexane. What makes the partial oxidation step even worse is that cyclohexane and air can form explosive mixtures during the mixing process. This has led to the complete destruction of at least one facility.3 Finally, the benzene hydrogenation step has the drawback of being very energy consuming due to the high hydrogen demand. In an alternative route, phenol is directly hydrogenated to cyclohexanol. This route suffers from high phenol prices and again has the drawback of needing three hydrogen molecules for hydrogenation which is energy intensive. Another alternative process developed by Asahi Chemical4 in Japan again starts with benzene, which is only partially hydrogenated to cyclohexene, saving one-third of the hydrogen otherwise needed. The cyclohexene is then hydrated directly with water using a very fine-grained custom zeolite catalyst of the HZSM5 type. The hydration reaction is highly selective, avoiding byproduct formation. Figure 1 gives an overview of the current routes from benzene or phenol to cyclohexanol and its uses in industry today. Even though the Asahi process is a big improvement because it avoids explosion hazards, saves energy, and avoids side products, it also has its drawbacks. The hydration step is being carried out in a slurry reactor, and the reaction mixture is then separated via distillation. This has two undesired effects. As there is a need for substantial amounts of the very fine-grained * To whom correspondence should be addressed. E-mail: [email protected]. † Max Planck Institute for Dynamics of Complex Technical Systems. ‡ Otto-von-Guericke University Magdeburg.

catalyst, the aqueous phase is laden with catalyst and has to be recycled, necessitating the handling of the resulting slurry. More significantly, as equilibrium conversion is limited to around 14%,1,2 the external separation step is energy intensive due to the large size of the recycle stream. As the reaction is equilibrium limited and (mildly) exothermic, it would seem ideally suited for a reactive distillation approach which would internalize the recycles and would allow for complete conversion within the column due to the countercurrent flows of the reactants. This approach was suggested in a previous paper by our group.5 The suggested process can be seen in Figure 2, and from the theoretical side it can be considered to be the ideal process being straightforward, safe, energy efficient, and without side products. However, integrating the direct hydration reaction into a reactive distillation column has two challenges: the phase splitting behavior and the reverse reaction which limits the product purity in case a fully reactive column were to be used. Catalyst Considerations for Reactive Column Design When considering integrating a reaction/separation process into a reactive distillation column, one of the first steps is to decide which catalyst should be used. As the reaction under consideration is catalyzed by strong acids, the use of sulfuric acid would be one option. However, if one is to produce pure

Figure 1. Reaction network for cyclohexanol production, use, and intermediates.1

10.1021/ie800303k CCC: $40.75  2008 American Chemical Society Published on Web 11/05/2008

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Figure 4. Flow sheet for the process idea of a two-step reactive distillation process for the indirect hydration of cyclohexene to cyclohexanol. Figure 2. Hypothetical ideal process for cyclohexanol production from cyclohexene by direct hydration.

Figure 3. Effects of catalyst polarity on reactant mole fractions at catalytic sites.

cyclohexanol at the bottom of the column, the bottom of the column has to be nonreactive as the backward reaction would otherwise split the pure product back into its constituents. This rules out sulfuric acid as it would always be present especially at the bottom of the column as it has a sufficiently high boiling point. In the literature, Zhang et al.6 have reported reaction rates for different strongly acidic zeolites and for Amberlyst 15, a macroporous ion-exchange resin. The reaction rates reported for zeolites were significantly higher than that reported for the resin. We have carried out similar measurements and have also conducted measurements with heteropolyacids, another group of strong, solid acids which we deemed potential candidates for catalyzing the reaction. The results of these experiments have shown that under certain conditions the zeolites can be the best catalysts for the direct hydration. However, there are many zeolites which in contrast showed very low activity even though they seemed very similar to others with high activity. Both the resin and the heteropolyacids offer much lower catalytic activity. Our interpretation of these results is that the ideal catalyst for catalyzing reactions with two liquid phases which show such extreme phase splitting as the system cyclohexene/water does is a catalyst which resides between the two liquid phases, allowing good access by both reactants. In all other cases the amount of the reactant not in direct contact with the catalyst is very low, and significant mass transfer resistances limit the reaction rate. In Figure 3 these effects have been visualized for better understanding. As a consequence, a reactor with a membrane at the interface with a mutual solubility for both phases would seem to be ideally

suited. However, this does not necessarily mean that a typical membrane reactor has to be used as some zeolites are basically shell catalysts concentrating their active sites on the outside. The inside of the catalyst pellets is hydrophobic and cyclohexene is concentrated there, whereas the outside is hydrophilic and immersed in water. This is our interpretation of the experimental data with respect to the different catalyst types and seems to be the reason why some zeolites perform so much better than all other catalysts found. If this interpretation is correct, however, the resulting consequences are that there is a need for catalysts having a polarity gradient and a very large catalyst surface area. This leads to very small catalyst particles. The result of this in turn is the catalyst slurry encountered in the Asahi process. If the solid catalyst were to be handled as a slurry within the column, this has two drawbacks in that it cannot be limited to the upper part of the column and that handling the slurry in the column runs the risk of it adhering to the column internals and clogging the column. An alternative to the slurry approach of handling fine catalyst particles is growing them directly on the surface of the column internals. We have conducted some exploratory experiments in collaboration with the Schwieger group (Erlangen University) which has vast experience in zeolite coating of metal structures7 to test such an approach. The results of these experiments indicate that in principle growing zeolite particles on column packings is possible but also that the amount of catalyst that can be fixed in the column this way is very low. Other approaches like catalysis in thermomorphic solvent systems as proposed e.g. by the Behr group (Dortmund University8) and the use of microemulsions to enlarge the surface areas as suggested e.g. by the Schoma¨cker group (Berlin Technical University9) are interesting current topics of research. However, they have not yet been applied to the varying temperatures in distillation columns and are generally not easily available. For this reason, in our work a different approach was chosen which was directly available: the use of a reactive entrainer. From the literature10 it is known that formic acid reacts at a significant rate with cyclohexene forming formic acid cyclohexyl ester (FCE). In a second step this ester can then be split with water into cyclohexanol and formic acid, retrieving the acid which hence acts as a reactive entrainer. This reaction scheme was previously suggested by our group11 and can be seen in the center of Figure 4. In previous publications11,12 we have published the reaction kinetic and activity coefficient model parameters needed for simulating this reaction scheme. The reaction route offers the advantage of having no significant amounts of side products if operated at temperatures below 333 K using Amberlyst 15 as catalyst. At higher

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temperatures, the formic acid decomposes into water and carbon monoxide in the presence of Amberlyst 15 at a measurable rate. For this reason the reaction is being suggested to be carried out at 10 kPa to move the boiling point for most ternary mixtures to below 333 K. The proposed process configuration is also shown in Figure 4. The following sections will show that such a configuration is feasible and should lead to the desired cyclohexanol as final product. Reactive Residue Curve Maps Reactive residue curve maps have become the standard tool for deciding whether a given reaction/separation task is feasible. The implicit assumptions that are underlying the reactive residue curve equations first formulated by Ung and Doherty13 are that a homogeneous catalyst is used whose amount does not change during reaction and distillation. For certain reactions which are e.g. catalyzed by sulfuric acid, we can expect this assumption be valid to a good extent. However, the reaction system being studied here is not of this type as the formic acid is both acidic catalyst and reaction partner in these reactions. For this reason a more general framework is needed that does not rely on these assumptions. When setting up their equation set to describe the residue curves, Ung and Doherty13 considered a batch reactor with a vapor side draw which is in thermodynamic equilibrium with the liquid. The reaction in their model system takes place exclusively in the liquid phase. In our work, in contrast, the model idea is not based on the batch reactor analogy but on an analogy with a continuously operated stirred tank reactor (CSTR) with one inlet and one (vapor phase) outlet: dNT )N _ in - N _ out + dt

∑ ∑ν

i,rRr

r

(1)

i

dNi ) xi,inN _ in - xi,outN _ out + dt

∑ν

i,rRr

(2)

r

In these equations, N _ in and N _ out are the total molar in- and outflows of the system. Ni and NT are the component and the total molar amounts respectively, νi,r are the stoichiometric coefficients of component i in reaction r, and Rr is the reaction rate of reaction r. As can be easily seen, eq 1 is the overall molar mass balance, and eqs 2 are the component molar mass balances. The reaction rate equations are given by

(

het (T)mcat,r Rr ) kf,r

KSLE,AKSLE,B (1 +

∑aK

2

i SLE,i)

i

)

hom + kf,r (T)ncat,r ×

(∏

Ve ae,r -

e

1 Keq,r

∏a

νp p,r

p

)

(3)

This formulation was chosen as the model was set up for solidcatalyzed and autocatalytic reactions. The first term in the left bracket is of a Langmuir-Hinshelwood type that is used for kinetic modeling of a heterogeneous catalytic reaction according to A + B S C (+D)

(4)

where A, B, C, and the optional D are adsorbed species on the catalyst. The right term in the left bracket of eq 3 is used to compute the homogeneous reaction rate which is assumed to depend on the molar amount of the homogeneous catalyst. The term in

the right bracket represents the thermodynamic driving force of the reaction. Inserting the overall molar mass balance into the component mass balance, setting the feed composition equal to the composition within the reactor and introducing the dimensionless time τ leads to dxi ) (xi - yi) + dτ

(

(1 + Darhet

∑

∑ (ν

i,r r 2 aref i KSLE,i)

i

(1 +

∑aK

2 i SLE,i)

xiνT,r) × het kf,r

het,ref kf,r

hom + xcat,r ·Darhom

i

hom kf,r hom,ref kf,r

)

R (5)

with Darhet )

het,ref kf,r mcat,rKSLE,AKSLE,B

N _ out(1 +

∑a

ref 2 i KSLE,i)

(6)

i

Darhom )

hom,ref kf,r NT N _ out

(7)

These definitions of the Damko¨hler numbers were selected to allow setting them to constant values. As our choice of the inlet concentration mathematically removed the inlet term from the equation because the concentration difference became zero, we can set the inlet flow rate arbitrarily. It is chosen to be such that the overall molar amount in the reactor/separator stays constant. This was the reason why the CSTR analogy was introduced and leads to the same mathematical result as the heating policies introduced by Ung and Doherty13 (i.e., the mathematical result of our feeding strategy is the equivalent to Doherty’s heating policy) even though formally a different model reactor was used. By specifying a vapor flow rate, it is now possible to set the homogeneous Damko¨hler number as desired for a given reference temperature and reaction. The heterogeneous Damko¨hler number can then be set by adjusting the catalyst amount (in moles or equivalents as e.g. given for Amberlyst 15). As the kinetic constants are strongly temperature dependent, in the definition of the Damko¨hler numbers a “reference” kinetic constant is used. This is the kinetic constant at a reference temperature, which is usually chosen to be the lowest boiling temperature observed in the system at a given pressure value. The same approach and reference state were chosen for the squared sums of adsorption potential terms. This allows for temperature-independent Damko¨hler numbers while temperature and adsorption dependencies are introduced in eq 5. The mole fraction of the catalyst in the homogeneous autocatalytic case is also modeled to be variable and hence not directly included into the homogeneous Damko¨hler number. Design of Cyclohexene Esterification Column The design of the cyclohexene esterification column shown in this section is based on the assumption of a pure cyclohexene feed without the inert impurities such as cyclohexane. These impurities behave like cyclohexene from a distillation standpoint but do not react. This simplification makes the feasibility studies presented here much clearer. The Asahi process also works with pure cyclohexene as a feed and separates the inert components using an upstream distillation column.1

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curve maps for 100 and 10 kPa are shown. As can be easily seen, the mixing gap is fairly limited, and at both pressure levels FCE is produced as the desired bottom product. The composition range in which the reaction can be safely carried out even at the lower pressure of 10 kPa without risking catalyst and formic acid decomposition is, however, still limited as the pure FCE reaches a boiling temperature exceeding 363 K.

Figure 5. Nonreactive residue curve maps for P ) 100 and P ) 10 kPa for the cyclohexene esterification reaction.

The esterification of cyclohexene with formic acid to produce FCE leads to ternary mixtures which in some cases show liquid-liquid phase splitting. In Figure 5 the nonreactive residue

Of course, the nonreactive residue curve map is a purely theoretical idea as formic acid is both a component and a catalyst for the reaction. If one were to carry out the reaction without the heterogeneous catalyst, the result can be seen in Figure 6 for different Damko¨hler numbers. As the Damko¨hler number increases, the pure FCE node is no longer stable as FCE can split back into its constituents. For a certain range of Damko¨hler numbers the separation and reaction effects counter each other, leading to a stable node close to the pure FCE corner. Above a certain critical Damko¨hler number the stable node disappears altogether, and pure formic acid is the bottom product to be expected. The change in location of the stable node as a function

Figure 6. Residue curve maps for different homogeneous Damko¨hler numbers at P ) 10 kPa for cyclohexene esterification.

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Figure 7. Trajectory of the stable and unstable nodes as a function of the Damko¨hler number for the cyclohexene esterification reaction (the Damko¨hler number at the critical point depends on h).

of Damko¨hler number was calculated by means of a oneparameter continuation calculation and can be seen in Figure 7. As the homogeneous reaction will always be present, the heterogeneous catalyst amount was introduced indirectly in the form of a ratio of Damko¨hler numbers abbreviated as h: hr )

Darhet Darhom + Darhet

Figure 8. Suggested column configuration and computed concentration profiles for the cyclohexene esterification column. The profiles are based on residue curve computations; the suggested countercurrent reactant feed points are not considered in the computations.

(8)

This ratio states to which extent the overall reaction rate can be attributed to the heterogeneous reaction. The purely homogeneously catalyzed residue curves shown in Figure 6 are obviously for an h ) 0. Interestingly, the trajectory of the stable node with increasing Damko¨hler number (shown in Figure 7) does not depend significantly on the value of h. The only effect h has is to change the position on the trajectory curve at a given overall Damko¨hler number. The critical Damko¨hler number at which the stable node close to the FCE corner disappears, and pure formic acid remains as the stable node also depends on the value of hsrising from 0.79 to 1.49 as h goes from 0 to 1. With respect to the reactive distillation process design, Figure 7 shows clearly that in order to attain pure FCE as the desired intermediary product one needs to have a nonreactive stripping section. However, considering the fact that formic acid selfcatalyzes the reaction, it is not possible to completely eliminate the reaction anywhere in the column. Nevertheless, it can be suppressed to its minimum limit by introducing the heterogeneous catalyst only into the top part of the column. It will significantly raise the overall Damko¨hler number there allowing setting a very low Damko¨hler number in the bottom part. Such a combination is shown in Figure 8, including a potential column concentration profile derived from the residue curve computations. To derive the column concentration profiles, two partial residue curves were computed that meet at the same concentration point at the lower end of the heterogeneously catalyzed zone. One curve was computed that starts at the composition where the heterogeneous catalytic zone starts and which ends in the stable node FCE. This partial residue curve was computed for a low Damko¨hler number of 0.001 as this part of the column is not heterogeneously catalyzed. The second partial residue curve was computed starting at the unstable cyclohexene/formic acid node and stops at the joining composition. This partial residue curve was computed at a high Damko¨hler number of 1 due to the fact that this is the region where the heterogeneous catalyst is placed. Since the two partial residue curves meet at the same point, they can be joined there. The combined curve

Figure 9. Nonreactive residue curve map for the ester splitting system at P ) 10 kPa.

shows the composition change such a partially heterogeneously catalyzed column has along its height coordinate. These are the column concentration profiles shown in Figure 8. One should keep in mind, however, that countercurrent feed locations are suggested in Figure 8 but were not included in the residue curve calculations the column profiles are based on. Furthermore, the profiles are theoretically limited to columns of infinite lengths at total reflux and are thus in the strict sense representing this limiting case only. Design of FCE Splitting Column The splitting of FCE with water leads to quaternary mixtures which in some cases exhibit liquid-liquid phase splitting. The system has four binary azeotropes of which the one between water and formic acid is high boiling and a ternary azeotrope between water, formic acid, and FCE. Also, there are two binary mixing gaps. The nonreactive residue curve map is shown as Figure 9. The system shown has two unstable nodes (formic acid and the azeotrope between water and FCE), two saddle points (the azeotrope between FCE and cyclohexanol and a ternary point containing FCE, water and formic acid), and two stable nodes (FCE and cyclohexanol). When considering Figure 9 closely, it can be seen that on top of the region in which residue curves are shown there exists another small region. The residue curves in this region were not shown for clarity reasons.

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Figure 10. Reactive residue curve map at chemical equilibrium conditions for the ester splitting reaction at P ) 10 kPa. Figure 12. Suggested column configuration and computed concentration profiles for the FCE splitting column. The profiles are based on residue curve computations; the suggested countercurrent reactant feed points are not considered in the computations.

Figure 11. Transformed reactive residue curve map with formic acid as reference component for the ester splitting reaction at P ) 10 kPa.

As can be seen, the nonreactive behavior of this system is such that by distillation preferentially the two light boiling components formic acid and the light boiling azeotrope between water and FCE are removed first. Depending on the location of the remaining mixture with respect to the cyclohexanol/FCE binary light boiling azeotrope, either cyclohexanol or FCE is obtained as the pure bottom product. From the perspective of column design for cyclohexanol production, the situation is attractive, as the composition at the azeotrope contains more FCE than cyclohexanol, allowing most starting compositions in composition space to converge to pure cyclohexanol as the bottom product. When forcing the system into reaction equilibrium, only one reactive distillation region remains as can be seen in Figure 10. All reactive residue curves are now forced onto the reaction equilibrium plane. The qualitative effect this has on the residue curve map, however, is fairly small as the behavior described for the nonreactive case still holds: the reactive residue curves still first approach the cyclohexanol/FCE binary edge where reaction stops playing a role as the respective second reaction partners are depleted. Then, depending on the location with respect to the azeotrope, either pure cyclohexanol or pure FCE is produced. Assuming reaction equilibrium allows for using the transformed composition variables Xi and Yi as suggested by Barbosa and Doherty.14 With these transformed variables, the threedimensional plot can be reduced by one dimension by looking only at the reaction equilibrium plane. The result is shown in

Figure 13. Suggested flow sheet for the new indirect cyclohexene hydration process for the production of cyclohexanol.

Figure 11. Here it was chosen to use formic acid as reference component. This diagram is especially suited to demonstrate the very large part of composition space that will ultimately lead to pure cyclohexanol as the bottom product of an RD column. The design implications that can be derived from Figures 9-11 are that both fully reactive and nonreactive columns will produce pure cyclohexanol as bottom product given the right starting composition. For economical reasons it might make sense to design the lower part of the column to be nonreactive to save catalyst where it is not needed. Figure 12 gives a suggested column design and a potential column concentration profile computed from the residue curves. Again, it should be noted that the suggested column configuration has a countercurrent feed strategy for the reactants which is suggested but not included in the concentration profile computations. Conclusions and Outlook In this paper, the feasibility of a novel reactive distillation process for the indirect hydration of cyclohexene to cyclohexanol using formic acid as a reactive entrainer is demonstrated (see Figure 13). The feasibility study presented is based on calculation and analysis of reactive and nonreactive residue curve maps. The reaction rates encountered in the two step process are acceptably fast while using Amberlyst 15 as a

Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 9587

catalyst, allowing for a good productivitysespecially when compared with the direct hydration route. As a next step, the residue curve based feasibility studies shown here need to be extended by means of detailed column simulations which also serve to identify suitable operating conditions of such columns. These simulations then have to be validated by experimental investigations especially of the reactive distillation in the presence of two liquid phases. This work is currently ongoing. Overall, the process suggested in this publication (see Figure 13) seems to be an attractive and safe new process for the production of cyclohexanol. Finally, from a scientific standpoint, it remains to be seen whether the two steps suggested above can be combined into a single reactive distillation column with one or multiple reaction zones which would in essence regain the ideal process initially suggested in Figure 2. It also remains to be seen whether such a strongly integrated system is economically and operationally desirable. Notation a ) activity Da ) Damko¨hler number h ) fraction of reaction rate attributable to heterogeneous reaction kf ) forward reaction rate constant Keq ) equilibrium constant KSLE ) Langmuir-Hinshelwood adsorption equilibrium parameters mcat ) mass of heterogeneous catalyst [kg] ncat ) mole number of homogeneous catalyst [mol] N ) mole number [mol] N _ ) molar flow [mol/s] R ) reaction rate [mol/s] R ) thermodynamic driving force of reaction t ) time [s] T ) temperature [K] x ) mole fraction, liquid phase y ) mole fraction, vapor phase ν ) stoichiometric coefficient τ ) dimensionless time (time divided by residence time) Subscripts and Superscripts e ) reactants i ) component index in ) inflow out ) outflow p ) products

r ) reaction index ref ) reference conditions as defined in text hom ) homogeneous het ) heterogeneous T ) total/overall

Literature Cited (1) Musser, T. M. Cyclohexanol and Cyclohexanone. In Industrial Organic Chemicals: Starting Materials and Intermediates; an Ullmann’s Encyclopedia; Wiley-VCH: Weinheim, 2003; Vol. 10, pp 279-290. (2) Behr, A.; et al. Cyclohexanol und Cyclohexanon. In Winnacker · Ku¨chler Chemische Technik; Wiley-VCH: Weinheim, 2005; Vol. 5, pp 7074. (3) Venart, J. E. S. Flixborough: The explosion and its aftermath. Process Saf. EnViron. Prot. 2004, 82, 105. (4) Mitsui, Osamu, Fukuoka & Yohei Process for Producing Cyclic Alcohol. United States Patent 4,588,846, 1986. (5) Steyer, F.; Qi, Z.; Sundmacher, K. Synthesis of cylohexanol by threephase reactive distillation: influence of kinetics on phase equilibria. Chem. Eng. Sci. 2002, 57, 1511. (6) Zhang, H.; Mahajani, S. M.; Sharma, M. M.; Sridhar, T. Hydration of cyclohexene with solid acid catalysts. Chem. Eng. Sci. 2002, 57, 315. (7) Mabande, G. T. P.; Pradhan, G.; Schwieger, W.; Hanebuth, M.; Dittmeyer, R.; Selvam, T.; Zampieri, A.; Baser, H.; Herrmann, R. A study of Silicalite-1 and Al-ZSM-5 membrane synthesis on stainless steel supports. Microporous Mesoporous Mater. 2004, 75, 209. (8) Behr, A.; Urschey, M.; Brehme, V. A. Aqueous biphasic catalysis as a powerful tool for catalyst recycling in telomerization and hydrogenation chemistry. Green Chem. 2003, 5, 198. ¨ nveren, H. H. Y.; Schoma¨cker, R. Hydroformylation with rhodium (9) U phosphine-modified catalyst in a microemulsion: comparison of organic and aqueous systems for styrene, cyclohexene and 1,4-diacetoxy-2-butene. Catal. Lett. 2005, 102, 83. (10) Saha, B.; Sharma, M. M. Esterification of formic acid, acrylic acid and methacrylic acid with cyclohexene in batch and distillation column reactors: Ion-exchange resins as catalysts. React. Funct. Polym. 1996, 28, 263. (11) Steyer, F.; Sundmacher, K. Cyclohexanol Production via Esterification of Cyclohexene with Formic Acid and Subsequent Hydration of the Ester - Reaction Kinetics. Ind. Eng. Chem. Res. 2007, 46, 1099. (12) Steyer, F.; Sundmacher, K. VLE and LLE data set for the system cyclohexane + cyclohexene + water + cyclohexanol + formic acid + formic acid cyclohexyl ester. J. Chem. Eng. Data 2005, 50, 1277. (13) Ung, S.; Doherty, M. F. Synthesis of reactive distillation systems with multiple equilibrium chemical reactions. Ind. Eng. Chem. Res. 1995, 34, 2555. (14) Barbosa, D.; Doherty, M. F. Design and minimum-reflux calculations for single-feed multicomponent reactive distillation columns. Chem. Eng. Sci. 1988, 43, 1523.

ReceiVed for reView February 21, 2008 ReVised manuscript receiVed June 20, 2008 Accepted October 6, 2008 IE800303K