Design and Analysis of Internally Heat-Integrated Reactive Distillation

Article pubs.acs.org/IECR

Design and Analysis of Internally Heat-Integrated Reactive Distillation Processes Yang Jiao,† San-Jang Wang,*,‡ Kejin Huang,† Haisheng Chen,† and Wei Liu† †

College of Information Science and Technology, Beijing University of Chemical Technology, Beijing 100029, People’s Republic of China ‡ Department of Chemical and Materials Engineering, Ta Hwa Institute of Technology, Chiunglin, Hsinchu 307, Taiwan ABSTRACT: Process intensification is aimed at integrating different processes in design to reduce utility consumption and capital investment, as well as at achieving environmental and safety benefits. Internally heat-integrated distillation and reactive distillation are representative examples of such design technology. In this study, the performance of internally heat-integrated reactive distillation, a novel technology combining both internally heat-integrated distillation and reactive distillation, is investigated for three ideal reactive distillation processes with a reaction zone located at top, middle, and bottom, respectively, of a reactive distillation column. The influences of reaction thermal effect (i.e., exothermic or endothermic reaction), relative volatilities between components, and chemical equilibrium constants on total annualized cost (TAC) are examined. Simulation results demonstrate that the internally heat-integrated reactive distillation can provide better economic benefit than conventional reactive distillation if the reaction zone can be properly placed in the high-pressure section or in both high- and low-pressure sections. The most TAC reduction from the implementation of internal heat integration can be obtained for the reactive distillation column with the reaction zone located at column top. Sundmacher and Kienle,14 and Luyben and Yu15 presented updated summaries. They showed the rapid progress of this technology sector in recent years. Most of these papers focused on real chemical systems, and every system had its own set of complexities in vapor−liquid equilibrium nonideality (azeotropes), reaction kinetics, and physical properties, etc. The discrete nature of chemical species and specific complexities in the vapor−liquid equilibrium seems to cloud the picture in understanding reactive distillation processes. On the other hand, the ideal reactive distillation processes16,17 offer a continuous spectrum in studying the process behavior by stripping away all the nonideal phase equilibrium and specific reaction rates. Design of ideal reactive distillation was explored,18−20 and comparison to conventional multiunit recycle processes was made for the processes with ‘‘stoichiometric’’ design where exact stoichiometric feeds were introduced.16,21 For reacting systems with unfavorable reaction kinetics (e.g., small chemical equilibrium constant) and vapor−liquid equilibrium, the design with one excess reactant may be a viable choice.22 Tung and Yu23 studied effects of relative volatility rankings on the stoichiometric design of reactive distillation processes. For quaternary reaction (A + B ⇔ C + D) systems, all possible relative volatility rankings can be classified into six processes based on the distribution of reactants and products in the relative volatility sequence. This provides a framework to study a range of reactive distillation processes (from easy to difficult). Following the approach of Tung and Yu,23 Lin24 explored the question of when and how the “excess-reactant” design should

1. INTRODUCTION Distillation column has long been known as an energy-intensive yet energy-inefficient process. It absorbs heat from a hightemperature heat source at the bottom and simply discharges it to a low-temperature heat sink at the top. To improve energy efficiency, the heat pump principle is often adopted as an effective means of reusing the rejected heat.1−3 In this approach, the vapor from the overhead is compressed as the heat pump fluid to supply energy required from the reboiler at the bottom. In this way, utility is saved by reusing the heat removed. The internally heat-integrated distillation column (HIDiC) used this heat pump concept to integrate the rectifying and the stripping zones. The rectifying zone is operated at a higher pressure than the stripping zone, and a compressor and a throttling valve are installed to adjust the pressures. The advantage of the HIDiC over the conventional column in terms of energy efficiency was shown by Mah et al.4 The exergy loss analysis of the HIDiC was studied and showed that the HIDiC had high energy efficiency.5,6 The HIDiC was also shown to give a substantial energy savings of around 30−50% in the separation of various mixtures when compared with the conventional column.7 In comparison with a direct vapor-recompression scheme, the HIDiC scheme can achieve up to a 20% reduction of the total annualized cost (TAC) due to its lower compression ratio, leading to a smaller sized compressor.8,9 Reactive distillation, the combination of reaction and separation operations into a single vessel, has demonstrated the potential for capital productivity improvements, selectivity enhancements, energy requirement abatement, and solvent reduction and even elimination in the process.10 The literature about reactive distillation up to 1992 was reviewed by Doherty and Buzad,11 and simulation and designs were reviewed by Taylor and Krishna.12 Books of Doherty and Malone,13 © 2012 American Chemical Society

Received: Revised: Accepted: Published: 4002

September 24, 2011 January 6, 2012 February 15, 2012 February 15, 2012 dx.doi.org/10.1021/ie202182h | Ind. Eng. Chem. Res. 2012, 51, 4002−4016

Industrial & Engineering Chemistry Research

Article

be used in the design of reactive distillation processes. An ideal quaternary system with a second-order reaction was used to illustrate the design. Results suggested that excess-reactant design was favored for systems with low chemical equilibrium constant and systems with the reactant being the lightest, types IR, IIR, and IIIR, tended to prefer “excess” design. Process intensification is a chemical engineering development that leads to a substantially smaller, cleaner, safer, and more energy-efficient process.25 It represents an important trend in chemical engineering and process technology and attracts more and more attention in industry and research communities. Internally heat-integrated distillation and reactive distillation are two promising technologies with capabilities of achieving substantial economical benefits from process intensification. In the study, the performance of internally heatintegrated reactive distillation, a technology combining internally heat-integrated distillation and reactive distillation, is investigated for ideal quaternary reaction systems with different types of reactions. To our best knowledge, there is no systematical study to examine the economical potential for internally heat-integrated reactive distillation. The ideal internally heat-integrated reactive distillation is designed and analyzed in the study through comparison with a reactive distillation column (RDC) involving a reaction zone located at top, middle, and bottom, respectively. The influences of reaction thermal effect (i.e., exothermic or endothermic reaction), relative volatilities between components, and chemical equilibrium constants on TAC are also examined. We will demonstrate that the internally heat-integrated reactive distillation can provide better economic benefit than conventional reactive distillation provided that the reaction zone can be properly placed in process synthesis and design.

Consider the following liquid-phase reversible reaction with second order:

The reaction rate of component i on tray j can be expressed as R j , i = νiMj(kFjxj ,Axj ,B − kBjxj ,Cxj ,D)

type III

LLK + HHK ⇔ LK + HK HK + HHK ⇔ LLK + LK

IIR

LLK + LK ⇔ HK + HHK

IIIP

LK + HHK ⇔ LLK + HK

IIIR

LLK + HK ⇔ LK + HHK

(4)

forward (EF)

12 000

backward (EB)

17 000

specific rate constant at 366 K (kmol s−1 kmol−1) forward (kF) backward (kB) vapor pressure constants (ln Psi = Avp,i − Bvp,i/T)

0.008 0.004

Avp Bvp

LK + HK ⇔ LLK + HHK

IIP

kBj = aBe−EB / RTj

activation energy (cal/mol)

A+B⇔C+D

type II

(3)

Table 2. Kinetic and Physical Properties for the Ideal System Studied

reaction:

IR

kFj = aFe−EF / RTj

where aF and aB are the preexponential factors, EF and EB are the activation energies, and Tj is the absolute temperature on tray j. Tray temperature is computed from the ideal vapor− liquid equilibrium, provided with Antoine vapor pressure expression. Kinetic holdup comes from the tray sizing by considering a weir height of 10 cm, and the column diameter is determined from the maximum vapor rate by assuming an Ffactor of 1. For the case when the reaction occurs in the reboiler and/or the condenser, the maximum kinetic holdup is taken to be 20 times the tray holdup. Kinetic and physical property data are given in Table 2. Ideal vapor- and liquid-

Table 1. Classification for Process Types

IP

(2)

where Rj,i is the reaction rate of component i on the jth tray (mol/s), νi is the stoichiometric coefficient which takes a negative value for the reactants and a positive value for the products, Mj is the kinetic holdup on reactive tray j (mol), and xj,i is the mole fraction of component i on tray j. The forward and backward specific rate constants described by the Arrhenius law of the rate constants on tray j are

2. IDEAL REACTING SYSTEMS For a quaternary system (A, B, C, and D) with the reversible reaction (A + B ⇔ C + D), there are 24 (4!) possible boilingpoint rankings for these four components. Because the two reactants (A and B) are interchangeable and the same applies to the two products (C and D), this leaves 6 (4!/2!/2!) possibilities. Tung and Yu23 classified these six possibilities into three types with two cases in each type. Table 1 gives the

type I

(1)

A+B⇔C+D

heat of reaction (cal/mol) heat of vaporization (cal/mol) relative volatilities (LLK/LK/HK/HHK)

−5000 6944 8/4/2/1

LLK

LK

HK

HHK

13.04 3862

12.34 3862

11.65 3862

10.96 3862

phase behavior is assumed for the reaction system, and the vapor−liquid equilibrium relationship can be expressed as

classification for process types according to the distribution of reactants and products in boiling-point ranking. LLK (lighter than light key), LK (light key), HK (heavy key), and HHK (heavier than heavy key) are used to denote the components ranging from the lightest boiler to the heaviest boiler. These six possibilities represent all possible combinations of relative volatility rankings for ideal reaction systems.

Pj = xj ,APA s + xj ,BP Bs + xj ,CPCs + xj ,DP Ds

(5)

yj , i = xj , iPi s/Pj

(6)

The vapor saturation pressure is calculated as ln Pi s = A vp, i − B vp, i /Tj

(7)

Equimolar overflow is assumed, so the liquid and vapor flow rates are constant in the nonreactive rectifying and stripping 4003

dx.doi.org/10.1021/ie202182h | Ind. Eng. Chem. Res. 2012, 51, 4002−4016


Article

sections. The reaction introduces no net changes in the molecular number, and therefore

and IIIP give one column configuration. The processes of types IR and IIR consist of a RDC and two separation columns. In the six configurations, the reaction zone is located at the upper section of the RDC for the processes of types IR and IIR while at the lower section of the RDC for the process of type IIP. The processes of the other three types have the reactive zone placed in the middle of the RDC. In this study, the effect of internal heat integration on the TAC of a RDC is investigated for the ideal reactive distillation processes. The location of reaction zone may have significant influence on the TAC for a HIRDC. Among the six cases mentioned, processes of types IP and IIP with stoichiometric design and type IR with excess design are selected as candidates for the implementation of internal heat integration due to their reaction zones located in the middle, lower, and upper sections of the RDC, respectively. In the following three sections, the three optimal nominal processes with the minimum TAC are designed first for every type and will be compared with the optimal processes of the internally heat-integrated reactive distillation with the minimum TAC later.

c

∑ νi = 0 i=1

(8)

For an internally heat-integrated reactive distillation column (HIRDC), the internal heat integration between high- and lowpressure sections is estimated with a lumped heat-transfer model given in the following equation, Q HI = USΔT

(9)

where QHI is the duty of heat integration. U, S, and ΔT represent overall heat-transfer coefficient, heat-transfer area, and temperature difference between high- and low-pressure sections, respectively. Vapor and liquid flow rates from tray j (Vj and Lj) in the rectification and stripping zones are calculated by Vj = Vj + 1 − Q HI/ΔHv

(10)

Lj = Lj − 1 + Q HI/ΔHv

(11)

3. DESIGN OF THE INTERNALLY HEAT-INTEGRATED REACTIVE DISTILLATION PROCESS FOR TYPE IP In exploring the optimal design for the processes of different types, TAC is used as the objective function. TAC is defined as

Those in the reaction zone are calculated by Vj = Vj + 1 − (λ /ΔHv )R j , i − Q HI/ΔHv

(12)

Lj = Lj − 1 + (λ /ΔHv )R j , i + Q HI/ΔHv

(13)

TAC = operating cost + capital cost/(payback period) (14)

where λ and ΔHv are the heat of reaction and the heat of vaporization, respectively. The last term in eqs 10−13 will be deleted for a RDC without heat integration. By considering column pressure as a design variable, Lin24 followed the approach of Tang and Yu23 and designed all of the six processes for the types in Table 1 with exact stoichiometric feeds, referred to as “stoichiometric” design later on, by using TAC as an objective function. Design results showed that type IP had the most favorable boiling point arrangement for a reactive distillation and thus gives the lowest TAC among all six cases. Type IR, the worst case scenario, gave the highest TAC, which was almost 10 times of that of type IP. Two column configurations were necessary for types IIR, IIP, and IR, where reaction zones were placed respectively at top, bottom, and both top and bottom sections of the RDC. However, poor reactant composition distribution in the reaction zone (e.g., high concentration of one reactant while a low one of the other) could be observed for all of the three most expensive types (i.e., IR, IIR, and IIIR). Lin24 found that RDC cost increased as the averaged reaction rate decreased due to poor performance in the reaction zone in types IR, IIR, and IIIR. To improve the reactant distribution in the RDC, making one of the reactant feeds to the RDC in excess, is a straightforward approach. By using TAC as the objective function, Lin24 found that the TAC of the excess design was less than half (41%) of that of stoichiometric design in type IR. The excess-reactant design in type IIR reduced the TAC by 30% from the stoichiometric design. However, design with excess reactant in type IIIR could have only 3% reduction of TAC. Simulation results showed that excess design was preferred for processes with reactant being the lightest (types IR, IIR, and IIIR), while the stoichiometric design was favored for processes with the product being the lightest (types IP, IIP, and IIIP). Figure 1 shows the process configurations with stoichiometric design or excess design for these six cases. Only the processes of types IP

Here, a payback period of 3 years is used. The formula for the TAC computation is given in the Appendix. A similar method to calculate TAC can also be found in Chen et al.26 and Zhang et al.27 We also limit the tray temperature in the reaction zone to be less than or equal to 403 K. This is a typical requirement for solid-catalyzed (e.g., Amberlyst 39) reactions. Thus, the design problem can be formulated as

minimize TAC X

(15)

subject to: x product C ≥ 0.95

(16)

x product D ≥ 0.95

(17)

Treaction,max ≤ 403 K

(18)

where X is the vector of design variables and the design variables vary as process configuration changes, xproduct C and xproduct D are product compositions of C and D, respectively, and Treaction,max denotes the maximum temperature in the reaction zone. During iterative search, the temperature is checked and operating pressure is adjusted if the temperature exceeds the limit. Notice that the design and simulation of the ideal reactive distillation use the rigorous distillation model programmed with Mathematica software. The steady-state model is solved using a modified Newton−Raphson method, and the satisfaction of constraints given in eqs 16−18 is taken as the convergence criterion. A multidimensional gridsearch approach is used in the study to minimize TAC. In this section, the effect of internal heat integration implemented in a RDC on the TAC is investigated for the process of type IP with exothermic reaction, endothermic reaction, different relative volatilities between components, and 4004



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Figure 1. Process configurations with stoichiometric design for types (a) IP, (c) IIP, and (e) IIIP and with excess design for types (b) IR, (d) IIR, and (f) IIIR.

column to three columns. Design variables also vary as process configuration changes. Here, type IP is used as an example to illustrate the design procedure for the optimal condition of the RDC. (1) Give the pressure of the RDC. (2) Guess the number of reactive trays. (3) Guess the number of trays in the rectifying zone. (4) Guess the number of trays in the stripping zone. (5) Change the reflux ratio and boilup ratio until the product specifications are met, and check if the constraint of maximum temperature in the reaction zone is satisfied. Go to step 1, and adjust the pressure of the RDC if the temperature exceeds the limit. (6) Go back to step 4 and repeat step 5 until the TAC is minimized. (7) Go back to step 3 and repeat steps 4 and 5 until the TAC is minimized. (8) Go back to step 2 and repeat steps 3−5 until the TAC is minimized.

different chemical equilibrium constants. TAC comparison is made among different configurations. 3.1. Design of the HIRDC for Exothermic Reaction. In type IP, two products are the lightest and heaviest components while two reactants are LK and HK. The reaction zone is located in the middle of a RDC from a simple rule that the reaction zone is placed at where the reactants are most abundant.23 Huang et al.28 and Kumar and Kaistha29 showed that changes in the feed tray location could lead to substantial energy saving. However, to independently identify the effect of internal heat integration on the TACs of a RDC and a HIRDC, in the study, two reactants are fed to the column by the same feed policy. In the process of type IP, the design variables include the pressure of the RDC, the number of rectifying trays, the number of reactive trays, the number of stripping trays, reflux ratio, and boilup ratio. The reflux and boilup ratios are adjusted to satisfy purity specifications of C and D from two outlet streams, respectively. Two reactants are fed onto the top and bottom of the reaction zone, respectively. In the processes of different types, process configurations may vary from one 4005



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Figure 2. Configurations of (a) nominal column, (b) H-HIRDC, and (c) M-HIRDC with exothermic reaction for type IP.

off as vapor for the HIRDC because the reflux flow rate at top is less than the vapor boilup flow rate at bottom in the nominal RDC of process type IP shown in Figure 2a. Not only capital investment but also utility consumption can be reduced. For the process of type IP, three configurations of HIRDC are designed, depending on the reaction zone located in the highpressure section (H-HIRDC), in both the high- and lowpressure sections (M-HIRDC), or in the low-pressure section (L-HIRDC). The internally heat-integrated design proposed by Olujic et al.8,30 is used in the study with the high-pressure section in the center and the low-pressure section in the annularity. Therefore, the number of trays for the high-pressure section and that for the low-pressure section are assumed to be the same in the study so that this annular structure can easily be constructed. Figure 2b shows the configuration of the H-HIRDC where rectification and reaction zones are located in the high-pressure section while the stripping zone is in the low-pressure section. Design variables in the H-HIRDC include the pressure of the high-pressure section, the numbers of trays of rectification and reaction zones, and boilup ratio. The top and bottom product

(9) Go back to step 1 and repeat steps 2−5 until the TAC is minimized. Figure 2a shows the optimal process flowsheet of type IP with stoichiometric design. Reactants A (LK) and B (HK) with a flow rate of 12.6 mol/s, respectively, are fed into a RDC operated at 9 bar. The RDC has a four-tray rectifying zone, followed by 15 reactive trays and four stripping trays. Products C (LLK) and D (HHK) with the same purity of 95 mol % are obtained from the distillate and bottoms, respectively. Table 3 gives the simulation data of the conventional RDC under the optimal condition. In this study, we number the top tray in the column to be the first one and the bottom tray in the column the last one. The TAC is calculated and given at the table bottom. In implementing internal heat integration for a HIRDC, the column is divided into a low-pressure section and a highpressure section. The vapor from the low-pressure section is pressurized by a compressor and entered into the high-pressure section. With heat transfer from the high-pressure section into the low-pressure section, heat input from the reboiler and heat removal by the condenser can be reduced. In addition, the condenser can be eliminated and the distillate product is drawn 4006



Article

Lower pressure in the high-pressure section is designed for M-HIRDC than H-HIRDC. Utility consumption is substantially reduced by implementing the internal heat integration between the high-pressure and low-pressure sections. A 32.8% reduction of TAC is obtained for M-HIRDC in comparison with RDC. When reaction zone is all placed in the low-pressure section, reaction rate and conversion are reduced, and convergent solution cannot be obtained to satisfy the design constraints in eqs 7−9. Between the two HIDRC configurations, M-HIRDC provides more economical benefit than H-HIRDC. 3.2. Design of the HIRDC for Endothermic Reaction. For a RDC involving a kinetically controlled reversible endothermic reaction, increasing operating pressure accelerates reaction rates. The enhancement of operating pressure facilitates reactant conversion and reaction heat load, reinforcing process intensification between the reaction operation and the separation operation involved. In the design of a RDC involving a kinetically controlled reversible endothermic reaction, the same design variables are used as those in the study for a kinetically controlled reversible exothermic reaction to minimize the TAC. The forward activation energy remains unchanged (12 000 cal/mol) while the backward activation energy is reduced as 7000 cal/mol. The reaction heat absorbed is 5000 cal/mol. Figure 3a shows the optimal flowsheet of process type IP for endothermic reaction, and Table 4 gives the simulation data. The TAC of the RDC proceeding endothermic reaction is higher than that proceeding exothermic one. The same design variables and design procedures as those for the exothermic reaction are also used to design three configurations of the HIRDC with endothermic reaction. Parts b and c of Figure 3 show the designed H-HIRDC and M-HIRDC, and their simulation data are also given in Table 4. More reactive trays are distributed in the low-pressure section than in the high-pressure section for the M-HIRDC. 5.1 and 20.4% reductions of TAC are obtained for H-HIRDC and M-HIRDC, respectively, in comparison with RDC. There is also no convergent solution for the L-HIRDC to satisfy the design constraints given in eqs 7−9 due to the low reaction rate and conversion. Again, the M-HIDRC provides the favorable design for type IP with the endothermic reaction. 3.3. Design of the HIRDC for Different Relative Volatilities. The analysis presented here indicates that M-HIRDC and H-HIRDC can provide more economic benefit than the RDC for exothermic or endothermic reaction. Moreover, M-HIRDC gives lower TAC than H-HIRDC. However, these results are based on the relative volatility equal to 2. To illustrate the TAC sensitivity to relative volatility, design of the HIRDC for a wide range of relative volatilities is investigated here. The same design variables and design procedures as those in sections 3.1 and 3.2 are also repeated in the design of RDC and HIRDCs with different configurations. There is still no convergent solution for L-HIRDC under the constraints of eqs 7−9. Figure 4 shows the effect of relative volatility on capital cost, operating cost, and TAC for the optimal RDC, H-HIRDC, and M-HIRDC configurations. Two HIRDCs need more capital cost than RDC. Especially, capital cost of H-HIRDC increases sharply as the relative volatility decreases. However, operating cost can be substantially reduced for HIRDCs by internal heat integration. Higher sensitivity of operating cost to relative volatility is observed for the RDC with low relative volatility. Saving of operating cost increases with decreasing relative volatility for two HIRDCs. There is only little

Table 3. Simulation Data for the Process of Type IP with Exothermic Reaction column configuration pressure (bar) total no. of trays no. of trays in rectifying zone (NR) no. of trays in reaction zone (NRX) no. of trays in stripping zone (NS) reactive trays heavy reactant feed tray light reactant feed tray heavy reactant feed flow rate (mol/s) light reactant feed flow rate (mol/s) top product flow rate (mol/s) bottom product flow rate (mol/s) reflux flow rate (mol/s) vapor boilup rate (mol/s) column diameter (m) Weir height (m) condenser heat-transfer area (m2) reboiler heat-transfer area (m2) total capital cost ($1000) column shell column trays heat exchangers compressor total operating cost ($1000/year) catalyst electricity utility TAC ($1000/year) a

RDC

H-HIRDC

M-HIRDC

9 23 4

12.79/2 36 a 7

10.66/2 34 a 8

15

11

18

4

18

8

5−19 5 19 12.6

8−18 8 18 12.6

9−26 9 26 12.6

12.6

12.6

12.6

12.6 12.6

12.6 12.6

12.6 12.6

31.15 35.14 0.782 0.1 107.34

3.98 0.905/1.402 0.1 -

3.98 0.810/1.205 0.1 -

51.64 381.3 130.5 3.6 247.2 363.6 8.6 355.0 490.7

5.85 772.0 325.2 10.5 396.8 39.5 120.7 8.4 78.0 34.3 378.1

5.85 683.5 268.8 8.0 376.1 30.6 102.1 17.7 50.1 34.3 329.9

a

Excluding a condenser and a reboiler.

purities are maintained by adjusting the pressure of the highpressure section and boilup ratio, respectively. The pressure of the low-pressure section is set at 2 bar, the same value as that used by Lin24 in the conventional distillation columns for the processes of different types. The number of trays in the lowpressure section is equal to that in the high-pressure section. Simulation data can also be found in Table 3 for the optimal H-HIRDC configuration. Operating cost mainly from utility consumption can be substantially reduced by making use of internal heat integration between the high-pressure and low-pressure sections. A 22.9% reduction of TAC is obtained for H-HIRDC in comparison with RDC. Figure 2c shows the configuration of M-HIRDC where reaction zones are located in both the high-pressure section and the low-pressure section. Design variables in the M-HIRDC include the pressure of the high-pressure section, the numbers of trays of rectification and stripping zones, the number of trays of the reaction zone in the high-pressure section, and the boilup ratio. The top and bottom product purities are also maintained by adjusting the pressure of the high-pressure section and the boilup ratio, respectively. Simulation data can also be found in Table 3 for the optimal M-HIRDC configuration. Reaction zones are equally distributed in high- and low-pressure sections. 4007



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Figure 3. Configurations of (a) nominal column, (b) H-HIRDC, and (c) M-HIRDC with endothermic reaction for type IP.

for HIRDCs. M-HIRDC still gives the most economical design under different chemical equilibrium constants. A reduction of above 30% TAC can be obtained for M-HIRDC in comparison with RDC. The results of Figures 4c and 5c indicate that the TACs of RDC and HIRDCs are more sensitive to relative volatility than chemical equilibrium constant.

difference in the operating costs between H-HIRDC and M-HIRDC. In the TAC comparison, two HIRDCs appear to be more economical than RDC. TACs of RDC and two HIRDCs decrease with increasing relative volatility. M-HIRDC again shows the lowest TAC among these designs. However, little difference of TACs between H-HIRDC and M-HIRDC is observed for relative volatility equal to 2.6. These simulation results indicate that HIRDC provides more economical benefit than RDC for various relative volatilities. The TAC difference between these two HIRDCs decreases with increasing relative volatility. 3.4. Design of the HIRDC for Different Chemical Equilibrium Constants. The effect of chemical equilibrium constant on TAC is also explored for the RDC and HIRDCs. Every process is optimized in terms of the TAC for a wide range of chemical equilibrium constants. No convergent solution is obtained for L-HIRDC under the constraints of eqs 7−9. Figure 5 shows the effect of chemical equilibrium constant on capital cost, operating cost, and TAC for different configurations. Again, operation cost can be substantially reduced by using internal heat integration even though RDC has the least capital cost. Capital cost of H-HIRDC gives the most sensitivity to the chemical equilibrium constant. Little variation in operation cost for various chemical equilibrium constants is observed

4. DESIGN OF THE INTERNALLY HEAT-INTEGRATED REACTIVE DISTILLATION PROCESS FOR TYPE IIP In this section, the design of an internally heat-integrated reactive distillation process including a HIRDC and a separation column is explored for type IIP with exothermic and endothermic reactions, respectively. 4.1. Design of the Internally Heat-Integrated Reactive Distillation with Exothermic Reaction. In the process of type IIP, the reaction zone is placed in the lower section of the RDC (C1) because the two reactants are heavier than the products. No products are withdrawn from the column bottoms. A second column (C2) is needed to separate the two products (LLK and LK). Figure 6a shows the optimal process flowsheet of type IIP with stoichiometric design. The design variables of the reactive distillation process include the pressure of the RDC, the numbers of rectifying and reactive trays of the RDC, 4008



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Table 4. Simulation Data for the Process of Type IP with Endothermic Reaction column configuration pressure (bar) total no. of trays no. of trays in rectifying zone (NR) no. of trays in reaction zone (NRX) no. of trays in stripping zone (NS) reactive trays heavy reactant feed tray light reactant feed tray heavy reactant feed flow rate (mol/s) light reactant feed flow rate (mol/s) top product flow rate (mol/s) bottom product flow rate (mol/s) reflux flow rate (mol/s) vapor boilup rate (mol/s) column diameter (m) weir height (m) condenser heat-transfer area (m2) reboiler heat-transfer area (m2) total capital cost ($1000) column shell column trays heat exchangers compressor total operating cost ($1000/year) catalyst electricity utility TAC ($1000/year) a

RDC

H-HIRDC

M-HIRDC

10 25a 5

13.98/2 30a 5

11.96/2 26a 7

15

10

16

5

15

3

6−20 6 20 12.6

6−15 6 15 12.6

8−23 8 23 12.6

12.6

12.6

12.6

12.6 12.6

12.6 12.6

12.6 12.6

21.9 43.15 0.778 0.1 84.71

21.22 1.001/1.585 0.1 -

21.22 0.804/1.267 0.1 -

63.42 381.7 138.8 3.9 239.0 452.5 8.5 444.0 579.7

31.19 756.2 317.5 10.5 379.8 48.4 297.9 9.4 105.9 182.6 549.9

31.19 601.2 223.3 6.4 338.3 33.2 261.3 18.6 60.1 182.6 461.7

Excluding a condenser and a reboiler.

the numbers of rectifying and stripping trays of the second column, the feed location of the second column, the reflux ratios of both columns, and the boilup ratio of column C2. The reflux and boilup ratios of column C2 are adjusted to satisfy purity specifications of C and D from two outlet streams, respectively. The pressure of column C2 is set at 2 bar. The process is designed with the minimum total TAC. The RDC has a fivetray rectifying zone, followed by seven reactive trays (including a reactive reboiler). It is operated at 6 bar. Reactants A (HK) and B (HHK) with a flow rate of 12.6 mol/s, respectively, are fed at the bottom and top of the reaction zone. The distillate product of column C1 is then fed to a 36-tray conventional distillation column (C2) at the 20th tray. Products C (LLK) with purity 99 mol % from distillate and D (LK) with purity 95 mol % from bottoms are obtained. Figure 6a also shows the stream information and operating conditions of the process with the minimum total TAC when the reflux ratio of column C1 is 1.03. In the design of HIRDC, the RDC shown in Figure 6a is divided into high- and low-pressure sections. M-HIRDC and L-HIRDC are the possible configurations for the RDC to implement internal heat integration. Design variables in the HIRDC include the pressure of the high-pressure section and the numbers of rectifying and reactive trays. The pressure of the

Figure 4. Effect of relative volatility on (a) capital cost, (b) operating cost, and (c) TAC for different configurations of type IP.

low-pressure section is also set at 2 bar. In the separation column (C2), design variables include numbers of rectifying and stripping trays, the feed location, the reflux ratio, and the boilup ratio. The reflux and boilup ratios of the separation column are manipulated to maintain purity specifications of C and D in two outlet streams, respectively. The same feed policy as that in the RDC is also adopted in the design of HIRDC. Two internally heat-integrated distillation processes with M-HIRDC and L-HIRDC, respectively, are optimized in terms of the total TAC. Simulation results again indicate that there is no convergent solution for the process with L-HIRDC. Only the process with M-HIRDC can be designed for type IIP. Figure 6b shows the configuration of the optimal process with M-HIRDC. The condenser is eliminated and the top product is drawn off as vapor because the reflux flow rate at the top is less than 4009



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Figure 6. Configurations of (a) nominal process and (b) internally heat-integrated process with M-HIRDC for type IIP with exothermic reaction.

Table 5. Cost Comparison of Different Configurations for the Process of Type IIP with Exothermic Reaction RDC capital cost ($1000) operating cost ($1000/year) TAC ($1000/year) total TAC ($1000/year)

C2

363.74 486.92 387.43 311.35 508.67 473.66 983.33

M-HIRDC

C2

734.77 530.48 277.53 209.21 522.45 386.04 908.49

(12 000 cal/mol) while reducing the backward activation energy as 7000 cal/mol. The same design variables as those for the exothermic reaction are used to minimize the total TAC. Figure 7a shows the nominal process flowsheet for endothermic reaction under the optimal condition, and Table 6 gives some costs of RDC and the separation column (C2) and the total TAC. The RDC with endothermic reaction needs more TAC than that with exothermic reaction for the process of type IIP due to more energy supply needed for the endothermic reaction. In the design of HIRDC configuration for type IIP with endothermic reaction, simulation results also indicate that there is no convergent solution for L-HIRDC. However, when reactive trays are extended from the low-pressure section to the high-pressure section, M-HIRDC can be designed and its optimal configuration is shown in Figure 7b. The M-HIRDC has an eight-tray rectifying zone, followed by two reactive trays in the high-pressure section. Some costs of M-HIRDC and C2 and the total TAC are also given in Table 6. It indicates that not only the TAC of separation column but also that of M-HIRDC are decreased by making use of internal heat integration for type IIP with endothermic reaction. A 12.2% reduction of total TAC can be obtained for the process with M-HIRDC in comparison with the conventional reactive distillation process.

Figure 5. Effect of chemical equilibrium constant on (a) capital cost, (b) operating cost, and (c) TAC for different configurations of type IP.

the vapor boilup flow rate at the bottom in the RDC given in Figure 6a. The pressure of the high-pressure section is designed at 7.47 bar. The vapor with higher pressure than that (6 bar) in the RDC is then fed to the separation column, and less TAC of the separation column can then be expected. Table 5 shows the cost comparison of different configurations for type IIP with exothermic reaction. While M-HIRDC gives a little higher TAC than RDC, the total TAC of the internally heat-integrated process is reduced by 7.61% in comparison with the conventional reactive distillation process. 4.2. Design of the Internally Heat-Integrated Reactive Distillation with Endothermic Reaction. To investigate the effect of a chemical reaction that absorbs heat on the TACs of the RDC and HIRDC, the heat of reaction is assigned as 5000 cal/mol by keeping the forward activation energy unchanged 4010



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configuration with the least total TAC for type IR with exothermic reaction under excess design. The RDC (C1) shown in Figure 8a consists of two zones (i.e., reaction and stripping zones) and is operated under total reflux. Products C (LK) and D (HK) and the excess reactant B leave the bottom of the RDC to the succeeding columns for further separation, where light product C of 95 mol % leaves the first conventional distillation column (C2) as distillate. Heavy product D with purity 95 mol % is obtained from the distillate of the column C3, while the excess reactant B with purity 95 mol % is withdrawn from the bottom of the same column and recycled back to the RDC. The reflux ratio of column C2 and reflux and boilup ratios of column C3 are adjusted to secure the purity specifications on C, D, and B from three outlet streams, respectively. The remaining design variables in the configuration consists of the following: the pressure of the RDC, the numbers of reactive and stripping trays in the RDC, the boilup ratio of the RDC, the numbers of trays of the first and second separation columns, the feed tray locations of the first and second separation columns, the boilup ratio of the first separation column, and the recycle ratio. They are all used to minimize the total TAC. The pressures of two separation columns are still set at 2 bar. Figure 8a also shows the stream information and operating conditions of the process with the minimum total TAC. Table 7 gives some costs for the conventional reactive distillation process of type IR with exothermic reaction. In the design of a HIRDC for the process of type IR, the RDC shown in Figure 8a is also divided into high- and lowpressure sections. The pressure of the low-pressure section is set the same as that of the separation column. In addition to reflux ratio of column C2 and reflux and boilup ratios of column C3 used to maintain product purities in three outlet streams, the remaining design variables include the pressure of the highpressure section in the HIRDC, the numbers of reactive and stripping trays in the HIRDC, the numbers of trays of the first and second separation columns, the feed tray locations of the first and second separation columns, the boilup ratio of the first separation column, and the recycle ratio. Figure 8b shows the optimal design of H-HIRDC, with the reaction zone located in the high-pressure section, for the process of type IR with exothermic reaction. The reboiler can be eliminated in the HIRDC because the vapor boilup flow rate at bottom is less than the reflux flow rate at top in the conventional RDC. The utility consumption is substantially reduced by implementing internal heat integration between the high- and low-pressure sections. Some costs for the internally heat-integrated reactive distillation process can also be found in Table 7. A 45.2% reduction of TAC can be obtained for the H-HIRDC in comparison with the RDC. M-HIRDC is the other possible reactive distillation configuration with internal heat integration by extending the reaction zone from the high-pressure section to the low-pressure one. Simulation results indicate that M-HIRDC with one reactive tray in the low-pressure section has a little higher TAC than the H-HIRDC shown in Figure 8b. The TAC of M-HIRDC increases with the number of reactive trays in the low-pressure section. That is, H-HIRDC gives the most economical design among different configurations for the process of type IR with exothermic reaction. 5.2. Design of the Internally Heat-Integrated Reactive Distillation with Endothermic Reaction. The effect of an endothermic reaction on the TACs of the RDC and HIRDC is also investigated by keeping the forward activation energy

Figure 7. Configurations of (a) nominal process and (b) internally heat-integrated process with M-HIRDC for type IIP with endothermic reaction.

Table 6. Cost Comparison of Different Configurations for the Process of Type IIP with Endothermic Reaction RDC capital cost ($1000) operating cost ($1000/year) TAC ($1000/year) total TAC ($1000/year)

C2

358.01 486.92 505.01 311.35 624.34 473.66 1098.00

M-HIRDC

C2

531.88 530.48 400.26 209.21 577.55 386.04 963.59

5. DESIGN OF THE INTERNALLY HEAT-INTEGRATED REACTIVE DISTILLATION PROCESS FOR TYPE IR In this section, the nominal process including a RDC and two separation columns is designed first by minimizing the total TAC. The internally heat-integrated process is then designed and compared with the nominal process. 5.1. Design of the Internally Heat-Integrated Reactive Distillation with Exothermic Reaction. The principle for the selection of the excess reactant is to identify the limiting reactant in the reactive zone. For the process of type IR, where the two reactants have a maximum boiling-point difference, the analysis of the composition profile given in the work of Tung and Yu23 revealed that reactant B (HHK) is the limiting reactant in the upper part of the RDC, and reactant A (LLK) is the limiting reactant in the lower part of the RDC. Limiting reactant should be made in excess to improve poor distribution of this reactant in the reaction zone. If reactant B is chosen to be in excess, the reactive zone should be placed in the upper section of the RDC, where the concentration of reactant A is higher. If reactant A is made in excess, the reactive zone should be located in the lower section of the RDC, where the other reactant B has a high concentration instead. Lin24 designed these different processes and showed that heavy-reactant-excess configuration with two reactants fed to the condenser gives the least TAC. The feed policy is also adopted in the design of a HIRDC for the process of type IR. Figure 8a gives the nominal 4011



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Figure 8. Configurations of (a) nominal process and (b) internally heat-integrated process with H-HIRDC for type IR with exothermic reaction.

Table 7. Cost Comparison of Different Configurations for the Process of Type IR with Exothermic Reaction capital cost ($1000) operating cost ($1000/year) TAC ($1000/year) total TAC ($1000/year)

RDC

C2

C3

H-HIRDC

433.10 314.13 469.86

526.87 345.12 520.74 1438.15

424.53 306.04 447.55

558.97 71.33 257.66

C2 526.87 345.12 520.74 1225.95

C3 424.53 306.04 447.55

among different configurations. A reduction of 40.1% on TAC is obtained for the H-HIRDC in comparison with the RDC. In these three sections, the effect of internal heat integration on TAC is investigated for the processes of different types with reaction zones located at the top, middle, and bottom in RDCs, respectively. The most economical processes of internally heatintegrated reactive distillation are designed for types IP, IIP, and IR, respectively. Simulation results indicate that the most TAC reduction from the implementation of internal heat integration can be obtained for the RDC with the reaction zone located at column top.

unchanged (12 000 cal/mol) and reducing the backward activation energy as 7000 cal/mol. The same design variables as those for the exothermic reaction are used to minimize the total TAC. Figure 9a shows the optimal nominal flowsheet for endothermic reaction, and Table 8 gives some costs of the configuration. The TAC of the RDC with endothermic reaction is less than that with exothermic reaction for type IR because the enhancement of operating pressure facilitates reactant conversion and reaction heat load for the endothermic reaction, reinforcing process intensification between the reaction operation and the separation operation involved. Two configurations of HIRDC, H-HIRDC and M-HIRDC, are designed to explore the effect of internal heat integration on the TAC. The optimal condition is obtained by using the same design variables as those for exothermic reaction to minimize the total TAC. Figure 9b shows the optimal configuration with H-HIRDC, and the corresponding costs can also be found in Table 8. The TAC of M-HIRDC is still higher than H-HIRDC and increases with the number of reactive trays in the lowpressure section. H-HIRDC gives the most economical design

6. DISCUSSION For a RDC involving reactions with high thermal effect, Wang et al.31 demonstrated that operating pressure affects significantly reactant conversion and the reaction heat load, giving rise to strong influences to process intensification between the reaction operation and the separation operation involved. The pressure gives the direct effect on the boiling point of the liquid in a RDC. The temperature of the bubbling liquid in the RDC in which the reaction occurs increases with the 4012



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Figure 9. Configurations of (a) nominal process and (b) internally heat-integrated process with H-HIRDC for type IR with endothermic reaction.

Table 8. Cost Comparison of Different Configurations for the Process of Type IR with Endothermic Reaction capital cost ($1000) operating cost ($1000/year) TAC ($1000/year) total TAC ($1000/year)

RDC

C2

C3

H-HIRDC

297.89 233.91 333.20

526.87 345.12 520.74 1301.49

424.53 306.04 447.55

412.04 59.18 196.52

C2

C3

526.87 345.12 520.74 1164.81

424.53 306.04 447.55

Table 9. Comparison of Optimal Improvement in TAC Reduction by Internal Heat Integration for Different Types with Exothermic and Endothermic Reactions

column pressure. If the reaction is exothermic, the equilibrium conversion of the forward reaction is reduced but the reaction rates of both forward and backward reactions are increased. If the reaction is endothermic, the equilibrium conversion of the forward reaction is enhanced and the reaction rates of both forward and backward reactions are increased. Table 9 shows the comparison of optimal improvement in TAC reduction by internal heat integration for different types with exothermic and endothermic reactions when relative volatility and chemical equilibrium constant (at 366 K) are both equal to 2. For type IP, the reaction zone is located at the middle of the RDC. Reactive trays are located at the bottom of the high-pressure section and also the top of the low-pressure section in the HIRDC. As pressure is increased in the highpressure section and decreased in the low-pressure section, reaction rate is increased in the reactive trays in the highpressure section and decreased in the reactive trays in the lowpressure section. Therefore the overall number of trays in the

RDC

case type IP type IIP type IR

NRX exothermic endothermic exothermic endothermic exothermic endothermic

15 15 6 4 17 18

HIRDC

reaction zone location NRX,1 NRX,2 middle middle bottom bottom top top

9 6 2 2 13 9

9 10 16 10 0 0

TAC reduction (%) 32.8 20.4 −2.7 7.5 45.2 41.0

HIRDC is approximately the same as that in the RDC. The use of HIRDC is desirable. For type IIP, the reaction zone is located at the bottom of the RDC. Most of reactive trays are located in the low-pressure 4013



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section in the HIRDC. As pressure is decreased in the lowpressure section, the reaction rate is decreased due to reduced pressure and lower temperature. Therefore the overall number of reactive trays required in the HIRDC is larger than that in the RDC. The benefits of using HIRDC are minimal. For type IR, the reaction zone is located at the top of the RDC. Reactive trays are all located in the high-pressure section in the HIRDC. As pressure is increased in the reaction zone, the reaction rate is increased due to increased pressure and higher temperature. Therefore the overall number of reactive trays needed in the HIRDC is smaller than that in the RDC. The use of HIRDC is most desirable. In all of the cases, the location of the reaction zone for HIRDC and RDC is similar for the process of the same type. L-HIRDC configuration is not found because the reaction zone is at the bottom of the RDC and reducing the pressure of this zone would lead to slower reaction rate and require an increase in reactor volume. It should be pointed out that the suitability of using HIRDC is not affected by whether the system is endothermic or exothermic. This is probably because the reactions are kinetically controlled and that limitation of equilibrium conversion is minimized by simultaneous removal of products in a RDC.

The height of a distillation column is calculated from the following equation: H = NT × 2 × 1.2/3.281

(A2)

The heat-transfer areas of the condenser and reboiler are calculated using the following equations: SCON = Q CON/(UCONΔTCON)

(A3)

SREB = Q REB/(UREBΔTREB)

(A4)

In terms of the above size estimations, the capital and energy costs of a distillation column are estimated using the following equations. column shell cost = 17.640DC1.066H 0.802

(A5)

tray cost = 229DC1.55NT

(A6)

total heat exchanger cost HI

= 7296SCON 0.65 + 7296SREB0.65 +

∑ (7296Si0.65) i=1 (A7)

7. CONCLUSIONS Internally heat-integrated distillation and reactive distillation are two promising technologies with capabilities of achieving substantial economical benefits. In the study, internally heatintegrated reactive distillation, a technology combining internally heat-integrated distillation and reactive distillation, is employed to further enhance the advantages of both technologies. Reactive distillation processes with internal heat integration are designed to achieve possible TAC saving for three ideal quaternary systems of types IP and IIP with stoichiometric design and type IR with excess design. For type IP with the reaction zone located in the middle of the RDC, M-HIRDC provides the most economical benefit for the exothermic reaction, endothermic reaction, various relative volatilities, and chemical equilibrium constants. The reaction rate is increased in the reactive trays in the high-pressure section and decreased in the reactive trays in the low-pressure section. The use of HIRDC is desirable. For type IIP with the reaction zone located at the bottom of the RDC, the process with M-HIRDC gives better economical design than conventional reactive distillation process for exothermic and endothermic reactions. In the M-HIRDC, most of the reactive trays are located in the lowpressure section and the reaction rate is decreased due to reduced pressure and lower temperature. The benefits of using HIRDC are minimal. For type IR with the reaction zone located at the top of the RDC, the process with H-HIRDC gives the least total TAC for exothermic and endothermic reactions. In the H-HIRDC, reactive trays are all located in the high-pressure section and reaction rate is increased due to increased pressure and higher temperature. The use of HIRDC is most desirable. The most TAC reduction can be obtained for the RDC with the reaction zone located at the column top.

compressor cost = 0.345FLP0.82 capital cost = column shell cost + tray cost + total heat exchanger cost + compressor cost

(A9)

utility cost = Q CONCCON × 24 × 300 + Q REB CREB × 24 × 300

(A10)

catalyst cost = 4∑ [π(DC/2)2 LR × 0.5 × DOC]POC (A11)

Ep = [γ /(γ − 1)]LPFLP[(HP/LP)(γ − 1)/γ − 1]/η (A12)

electricity cost = EpCELC × 24 × 300

(A13)

operating cost = utility cost + catalyst cost + electricity cost

(A14)

Nomenclature

Avp,i = Antoine vapor pressure coefficient aB = backward preexponential factor aF = forward preexponential factor Bvp,i = Antoine vapor pressure coefficient CCON = energy cost for condensing ($/kJ) CELC = electricity cost for compressing ($/kW) CREB = energy cost for reboiling ($/kJ) DC = diameter (m) DOC = density of the catalyst (kg/m3) EB = backward activation energy (cal/mol) EF = forward activation energy (cal/mol) Ep = electric power of compressor (kW) FLP = vapor flow rate at the top of the low-pressure column (kmol/s) H = height of distillation column (m)

■

APPENDIX Assuming an F factor of 1 in engineering units, the diameter of a distillation column is calculated from the following equation: DC = 0.01735(MW T /P)0.25 VNT0.5

(A8)

(A1) 4014



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(4) Mah, R.; Nicholas, J. J.; Wodnik, R. B. Distillation with Secondary Reflux and Vaporization: A Comparative Evaluation. AIChE J. 1977, 23, 651. (5) Takamatsu, T.; Nakaiwa, M.; Huang, K.; Akiya, T.; Noda, H.; Nakanishi, T.; Aso, K. Simulation Oriented Development of a New Heat Integrated Distillation Column and Its Characteristics for Energy Saving. Comput. Chem. Eng. 1997, 21, S243. (6) Nakaiwa, M.; Huang, K.; Owa, M.; Akiya, T.; Nakane, T.; Sato, M.; Takamatsu, T.; Yoshitome, H. Potential Energy Savings in Ideal Heat-Integrated Distillation Column. Appl. Therm. Eng. 1998, 18, 1077. (7) Iwakabe, K.; Nakaiwa, M.; Huang, K.; Nakanishi, T.; Røsjorde, A.; Ohmori, T.; Endo, A.; Yamamoto, T. Energy Saving in Multicomponent Separation Using an Internally Heat-Integrated Distillation Column (HIDiC). Appl. Therm. Eng. 2006, 26, 1362. (8) Olujic, Z.; Sun, L.; de Rijke, A.; Jansens, P. J. Conceptual Design of an Internally Heat Integrated Propylene-Propane Splitter. Energy 2006, 31, 3083. (9) Schmal, J. P.; van der Kooi, H. J.; de Rijke, A.; Olujic, Z.; Jansens, P. J. Internal Versus External Heat Integration Operational and Economic Analysis. Chem. Eng. Res. Des. 2006, 84, 374. (10) Malone, M. F.; Doherty, M. F. Reactive Distillation. Ind. Eng. Chem. Res. 2000, 39, 3953. (11) Doherty, M. F.; Buzad, G. Reactive Distillation by Design. Trans. Inst. Chem. Eng. 1992, 70A, 448. (12) Taylor, R.; Krishna, R. Modelling Reactive Distillation. Chem. Eng. Sci. 2000, 55, 5183. (13) Doherty, M. F.; Malone, M. F. Conceptual Design of Distillation Systems; McGraw-Hill: New York, 2001. (14) Sundmacher, K.; Kienle, A. Reactive Distillation; Wiley-VCH: Weinheim, Germany, 2003. (15) Luyben, W. L.; Yu, C. C. Reactive Distillation Design and Control; Wiley-VCH: New York, 2008. (16) Kaymak, D. B.; Luyben, W. L. A Quantitative Comparison of Reactive Distillation with Conventional Multiunit Reactor/Column/ Recycle Systems for Different Chemical Equilibrium Constants. Ind. Eng. Chem. Res. 2004, 43, 2493. (17) Cheng, Y. C.; Yu, C. C. Effects of Feed Tray Locations to the Design of Reactive Distillation and Its Implication to Control. Chem. Eng. Sci. 2005, 60, 4661. (18) Dragomir, R. M.; Jobson, M. Conceptual Design of Single-Feed Kinetically Controlled Reactive Distillation Columns. Chem. Eng. Sci. 2005, 60, 5049. (19) Hoffmaster, W. R.; Hauan, S. Using Feasible Regions to Design and Optimize Reactive Distillation Columns with Ideal VLE. AIChE J. 2006, 52, 1744. (20) Chen, C. S.; Yu, C. C. Effects of Relative Volatility Ranking on Design and Control of Reactive Distillation Systems with Ternary Decomposition Reactions. Ind. Eng. Chem. Res. 2008, 47, 4830. (21) Kaymak, D. B.; Luyben, W. L.; Smith, O. J. IV. Effect of Relative Volatility on the Quantitative Comparison of Reactive Distillation and Conventional Multiunit Systems. Ind. Eng. Chem. Res. 2004, 43, 3151. (22) Chen, F.; Huss, R. S.; Malone, M. F.; Doherty, M. F. Simulation of Kinetic Effects in Reactive Distillation. Comput. Chem. Eng. 2000, 24, 2457. (23) Tung, S. T.; Yu, C. C. Effects of Relative Volatility Ranking to the Design of Reactive Distillation. AIChE J. 2007, 53, 1278. (24) Lin, L. C. Effects of Relative Volatility Ranking to the Design of Reactive Distillation: Excess-Reactant Design. M.S. Thesis, National Taiwan University, Taipei, Taiwan, 2007. (25) Reay, D. The Role of Process Intensification in Cutting Greenhouse Gas Emissions. Appl. Therm. Eng. 2008, 28, 2011. (26) Chen, H.; Huang, K.; Wang, S. A Novel Simplified Configuration for an Ideal Heat-Integrated Distillation Column (Ideal HIDiC). Sep. Purif. Technol. 2010, 73, 230. (27) Zhang, X.; Huang, K.; Chen, H.; Wang, S. Comparing Three Configurations of the Externally Heat-Integrated Double Distillation Columns (EHIDDiCs). Comput. Chem. Eng. 2011, 35, 2017.

HHK = heavier than heavy key HI = heat integration HIRDC = heat-integrated reactive distillation column HK = heavy key HP = high pressure (bar) kB = backward specific rate constant (kmol s−1 kmol−1) kF = forward specific rate constant (kmol s−1 kmol−1) L = liquid flow rate (mol/s) LK = light key LLK = lighter than light key LP = low pressure (bar) LR = height of weir (m) Mj = kinetic holdup on reactive tray j (mol) MW = molecular weight of a mixture (g/mol) NR = number of rectifying trays NRX = number of reactive trays NS = number of stripping trays NT = number of total trays P = pressure (bar) Pis = vapor pressure of component i (bar) POC = price of the catalyst ($/kg) QCON = condenser duty (kJ/h) QREB = reboiler duty (kJ/h) R = reflux flow rate (mol/s) Rj,i = reaction rate of component i on the jth tray (mol/s) RDC = reactive distillation column SCON = heat-transfer area of condenser (m2) SREB = heat transfer area of reboiler (m2) T = temperature (K) ΔT = temperature difference (K) TAC = total annual cost ($/year) UCON = overall heat-transfer coefficient of condenser (kJ h−1 K−1 m−2) UREB = overall heat-transfer coefficient of reboiler (kJ h−1 K−1 m−2) V = boilup flow rate (mol/s) VNT = maximal vapor flow rate of a distillation column (mol/s) X = vector of design variables xj,i = liquid mole fraction of component i on tray j yj,i = vapor mole fraction of component i on tray j γ = ratio of heat capacities η = efficiency of compressor ν = stoichiometric coefficient λ = heat of reaction (cal/mol) ΔHv = heat of vaporization (cal/mol)

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AUTHOR INFORMATION

Corresponding Author

*Tel.: +886-3-5927700, ext. 2853. Fax: +886-3-5927310. E-mail: [email protected].

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ACKNOWLEDGMENTS

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REFERENCES

This work is supported by the National Science Council of ROC under Grant No. NSC 99-2221-E-233-008.

(1) Null, H. R. Heat Pumps in Distillation. Chem. Eng. Prog. 1976, 72, 58. (2) King, C. J. Separation Processes; McGraw-Hill: New York, 1980. (3) Smith, R. Chemical Process Design; McGraw-Hill: New York, 1995. 4015



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(28) Huang, K.; Nakaiwa, M.; Wang, S. J.; Tsutsumi, A. Reactive Distillation Design with Considerations of Heats of Reaction. AIChE J. 2006, 52, 2518. (29) Kumar, M. V. P.; Kaistha, N. Internal Heat Integration and Controllability of Double Feed Reactive Distillation Columns, 1. Effect of Feed Tray Location. Ind. Eng. Chem. Res. 2008, 47, 7294. (30) Olujic, Z.; Fakhri, F.; de Rijke, A.; de Graauw, J.; Jansens, P. J. Internal Heat IntegrationThe Key to an Energy-Conserving Distillation Column. J. Chem. Technol. Biotechnol. 2003, 78, 241. (31) Wang, S.; Huang, K.; Lin, Q.; Wang, S. J. Understanding the Impact of Operating Pressure on Process Intensification in Reactive Distillation Columns. Ind. Eng. Chem. Res. 2010, 49, 4269.

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Design and Analysis of Internally Heat-Integrated Reactive Distillation

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