Ind. Eng. Chem. Res. 2009, 48, 4719–4730
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Feasibility of Reactive Distillation for Fischer-Tropsch Synthesis. 3. Seethamraju Srinivas, Ranjan K. Malik, and Sanjay M. Mahajani* Department of Chemical Engineering, Indian Institute of Technology, Powai, Mumbai, 400076 India
Reactive distillation (RD), a proven reactive separation method that can enhance yields and improve product selectivity in multiple reactant/product systems, was shown to be feasible for FTS in our earlier papers using a simplified kinetics [Srinivas et al. Feasibility of Reactive Distillation for Fischer-Tropsch Synthesis. Ind. Eng. Chem. Res. 2008, 48, 889-899; DOI 10.1021/ie071094p] as well as a detailed kinetics [Srinivas et al. Feasibility of Reactive Distillation for Fischer-Tropsch Synthesis. 2. Ind. Eng. Chem. Res. 2009; DOI 10.1021/ ie801887m]. In-built thermodynamic procedures of Aspen Plus, along with a detailed kinetic model that predicts product distribution, were used in performing the simulations. In this paper, we present detailed parametric studies like effect of reflux ratio, H2/CO feed ratio, number of nonreactive stages, etc. Conversion, yield, olefin-to-paraffin ratio and product distribution are the parameters used for comparison among the different cases studied. Within RD mode for FTS, some of the alternate column configurations like those with a side-heat removal and a side-draw, a hybrid column with reactive and nonreactive stages, are also explored and investigated. 1. Introduction Fischer-Tropsch synthesis provides a product distribution rather than a single product. To make the overall process economically viable, FT product selectivity and yield are important, which vary according to the syngas feed composition, the type of reactor, and the catalyst employed, and the operating conditions like temperature, pressure, gas velocities used, etc. In the conventional reactors, the selectivity is either toward wax (for LTFT process using fixed-bed or slurry reactors) or gasoline (for HTFT reactors of the fixed- or circulating-fluidized bed type). A proper combination of both the processes helps in changing the product mix. However, the final product necessitates a large number of downstream operations in either of the processes. In our previous works (DOI 10.1021/ie071094p, DOI 10.1021/ie801887m),1,2 we have shown with the help of simulations that an alternative reactor configuration in the form of reactive distillation (RD) is a potential candidate for FTS to increase the selectivity toward the desired fraction. The other advantages RD offers for FT were also listed, notable among them being the utilization of heat of reaction to maintain saturation conditions, the possibility of having multiple product streams as side-draws with each rich in a fraction like gasoline, diesel, etc., and the reduction in the plant size. In this work, parametric studies on the model used in our earlier paper (DOI 10.1021/ie801887m)2 are reported with the aim of demonstrating the flexibility and aforesaid advantages that RD can offer for FTS by manipulation of various parameters. Both design and operating parameters are looked at. The effect on conversion and selectivity are also elucidated in each case. Different possible column configurations like a hybrid column with both reactive and nonreactive stages and columns with side-heat removal and side-draws in RD (see Figure 1) are simulated using Aspen Plus, and their potential is presented with the observations explained on the basis of the major influencing factors considered. 2. Parametric Studies 3
Base Case Simulations. Aspen Plus is chosen to perform the rigorous computations for the refluxed rectifier configuration * To whom correspondence should be addressed. E-mail: sanjaym@ iitb.ac.in. Tel.: (022) 2576 7246. Fax: (022) 2572 6895.
through the “RadFrac” module. The simulations were performed with the following components: CO, H2, CO2, H2O, paraffins from C1 to C30, and olefins from C2 to C20. Thermodynamic properties and calculations are taken care of by choosing the “PRMHV2” option. The detailed FT kinetic model of Wang et al.4 was incorporated using an external user-defined FORTRAN code to enable reactions. Further details regarding the components, kinetics, and VLE model can be found elsewhere.2 Vapor-liquid-free water calculations are enabled in all the simulations to take care of water, which is a major reaction product. The feed flow-rate to all the simulations was 1000 mol/h of syngas (CO + H2) at 250 °C and 25 atm, with three different H2/CO ratios used: 2, 1, and 0.67. A total of six stages, with four reactive stages, a nonreactive reboiler, and a nonreactive condenser, are present. The condenser is specified to be of partial vapor-liquid type having both a vapor and a liquid withdrawal. Free water calculations are enabled only in the condenser and not on the trays. There is a liquid withdrawal from the column bottom, thus giving a total of three exit streams from the RD column. Feed being in the
Figure 1. Schematic of RD configuration (refluxed rectifier) suggested for FTS. (a) Fully reactive column with side-draws; (b) hybrid column with side-draws and side-duties: (1) feed below the reactive section; (2 and 3) partial condenser product streams; (4) side-draws from reactive/nonreactive stages; (5) bottoms product stream; (6) nonreactive stage; (7) reactive stage with catalyst loaded; (8) side-duty to remove reaction exothermic heat from the reactive stage.
10.1021/ie801888v CCC: $40.75 2009 American Chemical Society Published on Web 04/21/2009
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vapor phase is introduced below the last stage of the column. The column is operated at 25 atm and the condenser temperature is 35 °C. To have reasonable extent of reaction on every stage, the catalyst loading on the reactive stages from top to bottom is 1.9, 1.8, 0.25, and 0.05 kg, giving a total of 4 kg of catalyst hold-up.1,2 Reflux ratio is varied to obtain convergence during the simulation runs. The reflux ratio and feed ratio are 0.215 and 2, respectively, for the base case. The fixed-bed reactor simulation is performed using the “RPlug” module2 and the slurry reactor using the “RCSTR” module.2 Henceforth, CSTR relates to the slurry reactor, FBR to the fixed-bed reactor, RR to the reactive distillation process of FT, and feed ratio to the ratio of H2/CO in the feed. Each carbon number corresponds to the sum of contributions from both paraffin and olefin. Light gas (C1-C4), naphtha (C5-C7), gasoline (C8-C12), diesel (C13-C18) and waxes (C19+) are the components considered. The temperatures in the column for the base case on the reactive stages are 208.58, 240.28, 209.79, and 203.13 °C, respectively, from top to bottom. The bottoms stream leaves at 216.27 °C. Conversions of CO and H2 are 40.36 and 10.92, respectively. Yields of light gas, naphtha, gasoline, diesel, and wax correspond to 0.13, 0.12, 0.13, 0.35, and 0.22 kg/h, respectively. The values of the chain-growth parameter on the respective stages and the composition profile are presented in our earlier work.2 Using the refluxed rectifier configuration, the effect of various parameters on the performance is analyzed on the basis of the following factors: reactant conversion, product selectivity, olefinto-paraffin ratio (henceforth referred to as the o/p ratio), amount of CO2 generated, utility ratio, and the tray temperatures. It is to be noted that water is a significant product of the FT reaction and takes part in the WGS reaction generating CO2 and H2. However, the concentration of water in the liquid phase either on mole or mass basis inside the column was found to be too small in all the simulation results to considerably affect the performance criteria like conversion, selectivity, etc. In other words, the column performance is insensitive to the liquid phase mole fraction of water. The idea behind tracking the CO2 generation rate is to follow the direction of WGS reaction as the parameter under study is varied. WGS reaction plays the important role of maintaining the H2/CO balance for a H2-deficit syngas. Being an equilibrium reaction in nature, it is favored in the forward direction at low temperature and low partial pressures of H2 and CO2 and in the reverse direction at high temperature and high concentrations of H2 and CO2. The effect of operating and design parameters on conversion and product selectivity is discussed in the following sections. The operating parameters considered are reflux ratio, condenser pressure and temperature, feed temperature, and the presence of CO2 in feed. The design parameters studied include feed ratio, catalyst distribution, number of reactive stages, number of nonreactive stages and their location, and the effect of sidedraw/side-duties. Effect of Reflux Ratio. Variation in reflux ratio over the range of 0.212-0.355 results in the trends shown in Table 1 and Figure 2. Conversion of the reactants, utility ratio, and both the liquid and vapor phase H2/CO ratios increase with an increase in the reflux ratio. The H2/CO ratios shown are the average values calculated from the reactive stages. The amount of CO2 formed as a result of the water-gas shift reaction also increases with an increase in the reflux ratio. This is due to the increased FT reaction rate leading to a rapid rate of water
Table 1. Effect of Reflux Ratio reflux ratio
Xco (%)
XH2 (%)
utility ratio
liquid H2/CO
vapor H2/CO
CO2 generated
0.212 0.215 0.225 0.235 0.245 0.26 0.28 0.31 0.35 0.355
39.76 40.36 42.39 44.44 46.53 49.69 54.02 61.38 71.28 76.82
10.74 10.92 11.53 12.17 12.84 13.93 15.55 19.21 28.56 35.82
0.55 0.55 0.55 0.56 0.56 0.57 0.58 0.64 0.81 0.95
2.326 2.336 2.371 2.408 2.446 2.509 2.605 2.806 3.164 3.395
2.436 2.446 2.477 2.511 2.547 2.604 2.693 2.878 3.198 3.407
0.0649 0.0659 0.0692 0.0725 0.0758 0.0809 0.0877 0.0989 0.113 0.118
formation and the decrease in hydrogen content, both of which enable the WGS reaction in the forward direction. An increase in reflux ratio leads to a rise in both the average reactive stage temperature and the bottom stage temperature as seen in Figure 2. The rise in the reactive stage temperature is more sensitive at the higher reflux ratios. Under distillation conditions in the absence of reaction, it is expected that with an increase in the reflux ratio, the temperature would decrease on some stages and increase on others as per the separation achieved. The same was verified by feeding a liquid stream consisting of FT components to an RD column with the reactions switched off, and it followed the trend expected. In the reactive case, further explanation for the rise in temperature with increase in reflux ratio is made difficult due to the factors involved as follows: (i) The partial condenser. Changing the reflux ratio changes the vapor to liquid distillate rates so that the reflux needs are met with an adjustment in the condenser duty too. (ii) Unlike in normal distillation, where an increase in reflux requirement can be met by an increase in reboiler duty, the same appears difficult here owing to the absence of the reboiler. The higher vapor load can only be met by the reaction which in turn needs lighter components to be generated as they remain in the vapor phase. This is possible at a higher reaction temperature. On the other hand, for liquid load to improve, condensable components in the form of heavier hydrocarbons are necessary. (iii) No particular trend was observed in the reflux composition to account for the increase in temperature when reflux ratio increases. At this juncture, we leave this effect of reflux as an observation rather than trying to explain it. The reason can possibly be found with the help of dynamic simulations, by moving from one steady-state value of reflux ratio to another, and observing the change in temperature and composition profiles in the column. An increase in reflux ratio increases the net yield (Figure 3) arising out of the higher conversion of the reactants. Since the increase in reflux ratio is synonymous with an increase in the average reactive stage temperature, both the liquid
Figure 2. Effect of reflux ratio on temperatures and condenser duty.
Ind. Eng. Chem. Res., Vol. 48, No. 10, 2009
Figure 3. Effect of reflux ratio on yields and chain growth parameter.
Figure 4. Effect of reflux ratio on yield distribution.
yields and the chain growth parameter R show a decreasing behavior. The decrease in either case is more drastic at higher values of reflux ratio (Figure 3). For a reflux ratio in the range of 0.212-0.26, the variation in liquid yield is very small (0.04 kg) and hence, it appears to be constant in Figure 3. This is because the chain-growth parameter falls sharply (below 0.8) as the reflux ratio increases beyond 0.28 leading to a fall in yields of diesel and wax, which are the main components in liquid yield. The yield distribution in Figure 4 also follows the temperature effect as expected with diesel and wax present at low reflux values but dominated by a light gas fraction at the higher reflux values. As for the o/p ratio, the average over all the fractions was almost constant. The o/p ratio in the case of the component fractions did not follow a specific trend: it increased for a few of them and decreased for the rest with an increase in the reflux ratio. The limitations2 pointed out in modeling olefin readsorption factor β need to be overcome to study the effect of reflux ratio on the product o/p ratios. To conclude, reflux ratio affects the yields and the conversion, and can be varied to alter the product distribution. Higher reflux ratios, however, reduce the liquid yield considerably. Effect of Catalyst Distribution. The catalyst distribution on the stages can also alter the product distribution and is illustrated through two sets of simulations, leading to four cases. Cases 1 and 3 have a loading of 1.9, 1.8, 0.25, and 0.05 kg catalyst on the reactive stages from top to bottom. Cases 2 and 4 have an equal catalyst weight of 1 kg on all the reactive stages. While cases 1 and 2 are simulated at a reflux ratio of 0.28, cases 3 and 4 correspond to a reflux ratio of 0.33. In all the cases, the total loading was fixed at 4 kg. All the other specifications remain the same in all the cases. Table 2 summarizes the results of these simulations. There is a considerable difference in the reactant conversions, average reactive tray temperature, and condenser duty between the equal and distributed catalyst loading cases at each reflux ratio. The average reactive temperature alone does not help in explaining the decrease in conversion with an increase in the reactive temperature at the reflux ratio of 0.28. As expected, at the higher reflux of 0.33, with a decrease in average reactive temperature, there is a corresponding decrease in conversion and increase in “R”. However, it is observed that for the case of 0.28, the temperatures on the individual trays follow a different trend. A
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“weighted” reactive temperature with respect to catalyst loading results in values of 240.51 and 228 °C for cases 1 and 2, respectively. This explains the decrease in the conversion despite an increase in the bottom and average reactive temperature for a reflux ratio of 0.28. At the higher reflux ratio value, there is a marked difference in the utility ratio too. While the liquid yields at the lower reflux ratio for both the equal and distributed loadings are same, the liquid yield in case of the higher reflux ratio for equal loading is double that of the distributed case. The net yields, R and o/p ratio, are also different in the cases considered. As for the product distribution, it is seen that the equal catalyst loading moves the spectrum from a lighter to a heavier end for both the cases of reflux ratio considered. This is indicated by the decrease in light gas fraction and an increase in the diesel and wax fractions. The naphtha and gasoline components remain constant or increase. Increase in R also supports the conclusion made regarding the product distribution. There could be other cases in between that have equal loading on two stages and different loading on the other two, for example, 1, 1, 1.4, and 0.6 kg, etc. Few other such cases were tried out and it was observed that the product distribution changes. This indicates that the best design needs to be worked out by proper optimization studies. Effect of Condenser Pressure and Temperature. The condenser pressure was set to 25 atm in the base case simulations, which is the normal FT reactor operating pressure. While an increase in pressure would favor the reaction owing to a reduction in the number of moles, it would make the separation difficult. To analyze this, simulations were performed for three different pressures: with 18, 22, and 25 atm at a reflux ratio of 0.215, and with 25, 27, and 30 atm at a reflux ratio of 0.3. All the simulations could not be performed at a single value of the reflux ratio because of difficulty in convergence. Tables 3 and 4 present the important results of these simulations. The first three rows/columns in Tables 3 and 4 correspond to a reflux ratio of 0.215 and the rest to a reflux ratio of 0.3. Irrespective of the reflux ratio, similar trends are observed with respect to the condenser pressure. Contrary to expectation, a decrease in pressure increases the conversion of the reactants. This can be justified by the relatively higher average reactive stage temperature and the liquid phase H2/CO ratio at lower pressures, both of which favor an increase in reaction rates. The utility ratio and amount of CO2 generated from the WGS reaction decreases with increasing pressure. A reduction in column pressure increases the vapor load in terms of mass with lighter components present in it and hence it is necessary for the condenser duty to increase (as observed in Table 3) to generate sufficient liquid to circulate reflux to the column. It is interesting to note that the fraction of water leaving as a liquid from the condenser decreases with an increase in pressure. The same effect is realized with an increase in reflux ratio at constant pressure. As expected, the product distribution shifts from the lower molecular weight range (methane) to the higher range (diesel and wax) with an increase in operating pressure corresponding to the decrease in reactive stage temperature (Table 4). The chain growth parameter, R, also increases with an increase in pressure supporting the conclusion made regarding the product yields. With an increase in pressure, the net product yields decrease resulting out of a decrease in reactant conversion, but the liquid yields increase. The o/p ratio follows a falling trend with increase in pressure. Contrary to expectation, the bottom temperature decreases (Table 3) with increase in pressure despite the formation of heavier fractions increasing. This is possible
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Table 2. Effect of Catalyst Distribution case
reflux ratio
Xco (%)
XH2 (%)
av reacn T (°C)
Qc (kW)
utility ratio
liquid H2/CO
vapor H2/CO
Tbottom (°C)
CO2 generated
1 2 3 4
0.28 0.28 0.33 0.33
54.02 49.24 67.09 55.57
15.55 13.59 24.13 16.09
233 237.56 259.94 244.95
-6.56 -6.09 -8.07 -6.64
0.58 0.56 0.73 0.59
2.605 2.745 2.993 3.104
2.693 2.838 3.045 3.199
221.23 222.66 228.71 225.08
0.0877 0.0804 0.107 0.0906
case light gas (mass %) naphtha (mass %) gasoline (mass %) diesel (mass %) wax (mass %) net yield (kg/h) liquid yield (kg/h) 1 2 3 4
30.94 21.88 74.06 39.98
15.88 16.22 9.99 19.40
12.28 13.23 6.25 11.37
28.38 32.53 6.84 20.81
12.53 16.14 2.86 8.43
1.28 1.16 1.68 1.32
0.65 0.66 0.24 0.48
R
o/p ratio
0.86 0.89 0.77 0.83
1.56 1.61 1.50 1.57
Table 3. Effect of Condenser Pressure pressure (atm)
Xco (%)
XH2 (%)
av reacn T (°C)
Qc (kW)
utility ratio
liquid H2/CO
vapor H2/CO
Tbottom (°C)
CO2 formed
18 22 25 25 27 30
53.58 43.86 40.36 58.7 55.08 51.24
17.47 12.17 10.92 17.65 15.74 14.15
252.74 226.22 215.45 240.52 230.22 219.22
-6.56 -5.61 -5.32 -7.01 -6.66 -6.31
0.66 0.56 0.55 0.61 0.58 0.56
2.595 2.406 2.336 2.727 2.628 2.533
2.645 2.497 2.446 2.805 2.722 2.644
228.18 221.25 216.27 222.35 219.59 214.83
0.0865 0.0715 0.0659 0.0949 0.0894 0.0835
Table 4. Effect of Condenser Pressure on Yield Distribution pressure (atm)
reflux ratio
light gas (mass %)
naphtha (mass %)
gasoline (mass %)
diesel (mass %)
wax (mass %)
net yield (kg/h)
liquid yield (kg/h)
R
o/p ratio
18 22 25 25 27 30
0.215 0.215 0.215 0.3 0.3 0.3
66.77 23.29 13.73 41.68 27.50 16.98
12.75 16.67 12.96 15.25 15.16 12.73
6.86 12.43 13.34 11.45 13.03 14.83
9.78 30.91 36.87 23.44 30.45 35.58
3.83 16.70 23.10 8.18 13.87 19.88
1.31 1.03 0.94 1.41 1.31 1.20
0.24 0.58 0.63 0.58 0.71 0.79
0.78 0.89 0.93 0.83 0.87 0.90
2.11 1.71 1.56 1.59 1.48 1.38
owing to the presence of Henry components whose presence can severely affect the bubble point calculations. It was observed that an increase in pressure (that led to a fall in reactant conversion) increased the mole fractions of H2 and CO in the liquid phase causing a fall in the bottom tray temperatures. To conclude, a higher condenser pressure appears favorable for higher liquid yields with a heavier product spectrum at the expense of lower reactant conversions. The appropriate reflux ratio also needs to be chosen, as illustrated by rows/columns 3 and 4 in Tables 3 and 4 which are at the same condenser pressure but different reflux ratio. It is to be noted that from the point of view of operational ease, it is much easier to control a single parameter (viz. the column pressure) rather than the temperatures on different trays. Thus, the choice of column pressure and corresponding reflux ratio is important. The effect of condenser temperature was verified by performing simulations at four values of 5, 15, 25, and 35 °C using a reflux ratio of 0.215. No significant difference in any of the parameters is observed as the condenser temperature varies (more details in Section 1 of Supporting Information). A lower condenser temperature aids in removing most of the water as a decanted liquid instead of allowing it to slip into the exit vapor stream. At the specified reflux conditions, the proportion of water leaving as a condensed liquid decreases rapidly from 72% to 1% as the condenser temperature rises from 5 to 35 °C. Thus, the final choice of the condenser temperature is an independent decision regardless of the process and would depend on the utility available for cooling like cooling water, chilled water, or some refrigerant. Effect of Feed Temperature. The syngas feed to the FT reactors is provided by coal gasifiers or natural gas reformers whose exit temperatures are of the order of 800-1200 °C. Before entering the reactors, the syngas is cooled, cleaned of impurities, and heated back to the required reaction temperature.
This process entails energy losses arising out of the cooling and heating operations. The study in this section aims to examine if the feed gas temperature has a role to play in the FTS conversion and yields in RD. Simulations were performed at temperatures of 200, 225, 250, 300, and 350 °C. These values cover the range of operating values for the LTFT (low temperature FT) and the HTFT (high temperature FT) processes. Figures 5-8 depict the trends seen for different parameters as a result of these simulations.
Figure 5. Effect of feed gas temperature on conversion and utility ratio.
Figure 6. Effect on reactive/bottom temperatures and condenser duty.
Ind. Eng. Chem. Res., Vol. 48, No. 10, 2009
Figure 7. Effect of feed gas temperature on product yields.
Figure 8. Effect on yields (kg/h), R and o/p ratio.
An increase in the feed gas temperature decreases the conversion of the reactants and the utility ratio (Figure 5). A lower average reactive stage temperature and a lower H2/CO ratio in the liquid phase are responsible for the fall in conversions with an increase in the feed gas temperature. The bottom tray temperature increases with increasing feed temperature (Figure 6) and is expected as it directly receives the increasingly hot feed. The additional heat input to the column in the form of an increase in the feed temperature expects one to observe an increase in the condenser duty as well. The opposite is, however, noticed (Figure 6) with the higher feed temperature requiring lower condenser duty. Apart from the heat input to the column by the feed gas, the other source of heat is the reaction exotherm. At the higher feed temperature, the part of the input energy going out from the bottom stream increases and this in combination with the reduced conversion (compared to the lower feed temperature case), and hence a lower exotherm release, reduces the condenser duty. Figure 7 shows the effect of feed temperature on the product distribution. Owing to the decrease in average reactive stage temperature, the spectrum shifts from the light gas fraction to the diesel and wax fractions. The naphtha and gasoline components remain almost constant. The relatively higher amounts of condensable components in addition to a lower level of exothermic heat release (due to lower rates of reaction) could be the reason for the observed decrease in condenser duty. As expected, the net yields follow the same trend (viz., decrease) as the conversion in Figure 8. R increases with an increase in feed temperature since the average reactive stage temperature is lowered. No significant difference is observed in the o/p ratio as the feed temperature varies. The final choice of the inlet feed gas temperature is dependent on the operations preceding the FT reactors. Nonetheless, a higher feed temperature seems to be favorable for increasing the selectivity toward diesel and wax with a trade-off in the reactant conversions and the net product yields. Effect of Feed Ratio. The feed ratio of H2/CO is a function of the coal used, the gasifier type, and its operating conditions,5 and any adjustments made using processes like shift reactors, etc. It is usually a design parameter and is not expected to vary considerably during normal operation.
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Simulations are performed for three different feed ratios: 2.03, 1, and 0.67. While the former is typical for syngas from natural gas reforming, the latter two result from coal gasification. We thus seek to examine how a given RR configuration performs for FTS with two different feed stocks, that is, natural gas and coal. Important parameters from the simulations are presented in Table 5. There is an increase in the CO conversion and the liquid phase H2/CO ratio as the feed ratio increases with a fall observed in the H2 conversion. The average reactive stage temperature drops too as the feed ratio rises. Thus, the increase in conversion is explained by the greater increase in the liquid phase H2/CO ratio which offsets the temperature effect. The changes in the condenser duty and utility ratio are not so significant. We noted earlier that R is both a function of the reactant partial pressures and the temperature. In this case, these two effects are opposing to each other. With an increase in the feed ratio, the average reactive stage temperature decreases while the ratio of the reactant partial pressures (H2 to CO) increases. The increase in H2 partial pressure dominates over the temperature dependency of R and results in its decrease as the feed ratio increases (Table 5). Correspondingly, the product distribution has increasing amount of lighter componentsslight gas, naphtha, and gasolinesas the feed ratio increases. The lowering in liquid yield despite an increase in conversion is also attributed to the formation of lower molecular weight components. The increased availability of H2 per mole of CO also enhances hydrogenation and termination to paraffins rather than an increase in chain length despite the presence of olefin readsorption. Thus, the o/p ratio is expected to decrease as the feed ratio increases and can be seen in Table 5. It can hence be concluded that a given RD configuration performs satisfactorily for any feed ratio (owing to the yields of gasoline, diesel, and wax fractions accounting to ∼70% in all the cases with only a slight difference in the yields) and the best operating conditions for the desired product selectivity need to be worked out. For example, an increase in reflux ratio at a feed ratio of 1 will help in bringing its conversion levels at par with a case using feed ratio of 2, but will result in a different product distribution. Effect of Reactive Stages. Apart from the catalyst distribution discussed in the earlier section, the number of reactive stages with the total catalyst loading kept constant also alters the product distribution. The number of reactive stages is increased from 4 to 13 while keeping the total catalyst weight fixed at 4 kg. The base case with four reactive stages is simulated at a reflux ratio of 0.33. As the number of reactive stages increased, the reflux ratio had to be changed to obtain convergence. The change made was kept to the minimum (Table 7). All the other specifications remain the same in all the cases. Tables 6 and 7 summarize the results of these simulations. There is a small difference in the reactant conversions (∼4%), average reactive tray temperature (∼6 °C) and bottom temperature (9 °C). No marked difference is observed in the condenser duty and the utility ratio. With an increase in the number of reactive stages, the maximum temperature observed in the column also increased as shown in Figure 9. Owing to a higher reactive stage temperature, the reactant conversions increase too. The increase in conversion corresponds to an increase in the net yields seen in Table 7. It is interesting to note that the liquid yields increase and later decrease as the catalyst loading per stage decreases. It is however expected that the liquid yields would decrease monotonically with an increase in the number of reactive stages because of the higher reactive temperatures. In the product
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Table 5. Effect of Feed Ratio feed ratio
Xco (%)
XH2 (%)
av reacn T (°C)
Qc (kW)
utility ratio
liquid H2/CO
vapor H2/CO
Tbottom (°C)
CO2 formed
0.67 1 2.03
24.86 29.08 43.41
19.53 15.42 11.85
226.83 221.09 218.93
-5.68 -5.6 -5.58
0.52 0.53 0.55
0.66 1.04 2.39
0.69 1.08 2.49
221.35 218.68 217.26
0.0744 0.0724 0.0708
feed ratio light gas (mass %) naphtha (mass %) gasoline (mass %) diesel (mass %) wax (mass %) net yield (kg/h) liquid yield (kg/h) 0.67 1 2.03
13.33 12.26 16.49
14.31 12.97 14.01
10.40 11.41 13.24
40.61 40.04 35.68
21.35 23.31 20.58
1.04 1.02 1.02
0.7 0.71 0.65
R
o/p ratio
0.95 0.95 0.91
3.44 2.48 1.57
Table 6. Effect of Reactive Stages reactive stages
Xco (%)
XH2 (%)
av reacn T (°C)
Qc (kW)
utility ratio
liquid H2/CO
vapor H2/CO
Tbottom (°C)
CO2 formed
4 5 6 7 8 9 10 11 12 13
55.57 56.03 56.39 56.42 56.52 56.9 57.62 58.31 58.79 59.08
16.09 16.07 16.06 15.98 15.97 16.14 16.5 16.86 17.13 17.3
244.95 245.46 245.56 245.65 245.65 246.45 247.92 249.28 250.35 250.79
-6.643 -6.675 -6.702 -6.702 -6.71 -6.746 -6.817 -6.885 -6.934 -6.964
0.59 0.58 0.58 0.58 0.57 0.58 0.58 0.59 0.59 0.59
3.10 3.07 3.05 3.02 2.98 2.96 2.95 2.95 2.94 2.93
3.2 3.17 3.14 3.11 3.07 3.05 3.03 3.02 3.02 3
225.08 223.52 222.06 220.73 219.29 217.9 216.85 216.41 216.36 216.24
0.0906 0.0914 0.092 0.092 0.0922 0.0927 0.0938 0.0949 0.0956 0.096
Table 7. Effect of Reactive Stages on Yields (kg/h) and o/p Ratio reactive stages
reflux ratio
net yield
liquid yield
R
o/p ratio
4 5 6 7 8 9 10 11 12 13
0.33 0.34 0.347 0.349 0.35 0.35 0.35 0.35 0.35 0.35
1.32 1.33 1.34 1.34 1.34 1.35 1.37 1.39 1.40 1.41
0.48 0.51 0.55 0.57 0.59 0.59 0.57 0.55 0.54 0.53
0.83 0.84 0.84 0.84 0.85 0.84 0.84 0.83 0.83 0.82
1.57 1.68 1.77 1.84 1.9 1.94 1.96 1.98 1.99 2
distribution shown in Figure 10, the light gas fraction decreases and then increases while the diesel component follows the opposite trend. There is no significant change observed in case
of naphtha, gasoline, and wax fractions. The o/p ratio averaged over the entire product range increases with a decrease in the amount of catalyst loaded per stage (Table 7). Figures 11 and 12 represent the liquid phase compositions of gasoline and diesel on the stages. An increase in the number of reactive stages helps in creating a zone in the column where the stages are rich in a single component thereby providing the option for a side-draw (stages 4 to 10 in Figure 11 with 13 reactive stages almost giving a pure gasoline stream) due to the increase in distillation effect. A similar effect can be realized in Figure 12 by the addition of a nonreactive section below the reactive stages to obtain a draw enriched in diesel. It is expected that an increase in the number of reactive stages with the total catalyst weight fixed will result in uniform temperature and reaction rate profiles. However, such an effect was not realized and both the profiles show a maximum before
Figure 9. Temperature profile in the column as number of reactive stages vary.
Figure 11. Effect on stage composition of gasoline.
Figure 10. Effect of reactive stages on product distribution.
Figure 12. Effect on stage composition of diesel.
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Table 8. Effect of CO2 in the Feed (RR) feed ratio
Xco (%)
XH2 (%)
av reacn T (°C)
Qc (kW)
utility ratio
liquid H2/CO
vapor H2/CO
Tbottom (°C)
CO2 formed
Xco+co2 (%)
1 2.03 2.6
28.81 46.68 47.26
7.81 8.11 6.7
220.2 225.96 219.36
-3.68 -3.89 -3.62
1.24 1.47 1.5
2.66 6.12 7.90
2.77 6.35 8.26
229.04 229.27 226.97
0.026 0.025 0.02
6.31 11.2 11.6
Table 9. Effect of CO2 in the Feed (FBR) feed ratio
Xco (%)
XH2 (%)
Reacn T (°C)
Qrt (kW)
utility ratio
liquid H2/CO
vapor H2/CO
CO2 formed
Xco+co2 (%)
1 2.03 2.6
38.16 57.77 58.44
10.56 10.48 8.84
221 226 220
-2.48 -2.58 -2.33
1.25 1.48 1.53
3.52 -
3.61 10.76 14.27
0.034 0.029 0.023
8.48 10.76 13.87
Table 10. Effect of CO2 in the Feed (CSTR) feed ratio
Xco (%)
XH2 (%)
Reacn T (°C)
Qrt (kW)
utility ratio
liquid H2/CO
vapor H2/CO
CO2 formed
Xco+co2 (%)
1 2.03 2.6
34.25 46.83 47.59
9.77 8.75 7.43
221 226 220
-4.91 -4.34 -4.66
1.27 1.51 1.57
3.35 8.53 11.50
3.43 8.72 11.21
0.03 0.023 0.019
7.69 8.53 11.21
decreasing. The optimum number of reactive stages under the conditions studied seems to be seven to nine if one were interested in maximizing the diesel yields. At higher number of reactive stages, the light gas fraction increases quite rapidly. Simulations beyond thirteen reactive stages did not converge for small changes in reflux ratio and hence were not tried out. To conclude, there appears to be an optimum number of reactive stages that aids in maximizing the yield of a given fraction (diesel in this case) with all the other specifications fixed. 3. Effect of CO2 in the Feed The presence of CO2 in the feed also affects the FT reactions. CO2 hydrogenation is feasible provided the extra hydrogen demand is met and a suitable catalyst with considerable activity is available.6,7 It is proposed that CO2 first gets converted into CO through the reverse WGS reaction, and the CO so formed helps in FT monomer formation.5 This section compares the performance of RR vis-a`-vis the conventional reactors (FBR and slurry reactors) in the presence of CO2 in the feed. The feed ratio is redefined as H2/(CO + CO2) and the utility ratio as moles of H2 consumed per mole of carbon oxides (CO and CO2 together). Three different feed ratios are considered: 1, 2.03, and 2.6, with the CO/CO2 ratio constant in each case. The corresponding H2/CO ratios are 2.5, 5.08, and 6.5, respectively. It is to be noted that the kinetics used was not developed with CO2 in the feed. However, since WGS reaction is included in the kinetics, it is expected to take care of conditions involving CO2 in the feed. In the absence of CO2 in the feed, it is observed that the WGS reaction is always in the forward direction on all the reactive stages in all the simulation results presented earlier. Its presence, however, shows CO2 consumption on some of the stages in the simulations in this section indicating a reverse WGS reaction. This was confirmed by the negative sign associated with CO2 and positive sign associated with CO in the component generation/depletion profile results table on the stages. A comparison of the performance of the three reactors is shown in Tables 8, 9, and 10 and in Figure 13. The conventional reactors are simulated at the average reactive stage temperature values of the RR configuration. As the feed ratio increases, an increasing amount of H2 helps in improving the overall conversion of the carbon oxides and the utility ratio in all the reactors. At the lower and higher feed ratios, this conversion is greatest in the FBR. The RD column gives the highest conversion at the medium feed ratio. An increase in the H2 partial pressure drives the WGS in the reverse direction leading to higher conversion of CO2 to CO, which corresponds to an
increase in the total carbon oxide conversion and decrease in the number of moles of CO2 formed. This trend is evident in all the three reactors as the feed ratio increases. The condenser duty (Qc) or the reactor duty (Qrt) do not follow any particular trend. This is possible owing to the differing rates of exothermic FT reactions and endothermic reverse WGS reaction as the feed ratio increases. The liquid yields from the FBR were not significant and hence the liquid phase H2/CO ratios are not shown. Comparison of these ratios between the RD column and CSTR shows that the values are higher in the CSTR and are enhanced by a factor of 1.3-1.7 as compared to the inlet feed H2/CO ratio. This increased solubility of reactants in the CSTR is explained by the use of C30 as a solvent in it. The bottom stage temperatures decrease in RR (Table 8) with an increase in the feed ratio owing to two reasons: increase in the presence of H2 resulting from its higher amounts in the feed with lowered conversion levels; decrease in R (Figure 13a) which gives a more volatile product spectrum. In all the three reactors, the R value decreases and then remains almost constant with a rise in the feed ratio. As for the products, the light gas and naphtha fractions increase and decrease; the diesel and wax fractions decrease and then increase in all the cases. The gasoline fraction decreases and then increases in the RR, follows the opposite trend in the FBR, and increases monotonically in the case of CSTR. The net yields decrease in the order FBR, CSTR, and RR at a given value of the feed ratio. The liquid yields are slightly higher in the slurry reactor compared to the refluxed rectifier. The individual product yields show some interesting features. The diesel fraction in RR compares well with the FBR at all feed ratios used but fares better than the CSTR at high feed ratios. The fraction of the product going to wax and light gas components is always highest in RD. The gasoline selectivity is almost same for the FBR and CSTR and approximately twice the values obtained in the RD column. Selectivity toward naphtha is marginally higher in RD at a low feed ratio and in the CSTR at high feed ratio values. The o/p ratio decreases with an increase in the feed ratio for all the reactors. These values are 1.82, 1.17, and 1.03 for RR; 2.23, 1.7, and 1.46 for FBR; 2.23, 1.71, and 1.47 for CSTR corresponding to feed ratios of 1, 2.03, and 2.6, respectively. The decrease is visible for each of the fractions as well. It is thus demonstrated that RD performs reasonably well compared to the conventional reactors in the presence of CO2 in the feed owing to comparable levels of conversions of the oxides of carbon (CO + CO2) as well as being competitive in the product yields.
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Figure 13. Effect of CO2 on product yields (a) RR; (b) FBR; (c) CSTR. Table 11. Effect of Nonreactive Stages in the Rectification Zone
case 1 case 2
Xco (%)
XH2 (%)
av reacn T (°C)
Qc (kW)
utility ratio
liquid H2/CO
vapor H2/CO
Tbottom (°C)
CO2 formed
61.38 52.27
19.21 14.9
245.66 226.68
-7.29 -6.38
0.64 0.58
2.81 2.52
2.88 2.61
223 209.77
0.0989 0.0846
light gas (mass%) naphtha (mass%) gasoline (mass%) diesel (mass%) wax (mass%) net yield (kg/h) liquid yield (kg/h) case 1 case 2
50.64 28.07
14.24 18.09
10.47 14.08
19.04 31.32
4. Complex Configurations The complex configurations include a RD column with both reactive and nonreactive sections, side-draws to remove multiple products from different stages, side heat duties to remove the heat of reaction from the reactive stages, and both side-draws and side heat duties. Effect of Nonreactive Stages. The RD column considered in the base case simulations was a fully reactive column. There can however be a hybrid RD column having both reactive and nonreactive sections. The nonreactive section has three placement options in the column: above all the reactive stages in the rectification section (reflux ratio: 0.31), below all the reactive stages in the stripping section (reflux ratio: 0.215) and in between the reactive stages (reflux ratio: 0.30). All the three possibilities were evaluated through simulations. It is also to be noted that the reflux ratio is adjusted to obtain convergence during the simulations and is indicated by the figures in brackets. The aim of adding nonreactive stages is to enhance the distillation effect and see if it effects the reaction and, thus, the conversion and product yields. In the case of liquid phase reactants, selectivity can be improved by removing the product that acts as a secondary reactant in other reactions by providing sufficient number of nonreactive stages.8 In the case of FT, the reactants are gas phase (H2 and CO) and reactions are in the liquid phase. Distillation can result in a solvent on the reactive stages with a higher molecular weight, which in turn increases the solubility of the reactants and the stage temperatures both of which favor higher reaction rates. Case 1 is the base case with a total of six stages with four reactive stages: 2, 3, 4, and 5. Case 2 has an additional stage in the rectifying section and the reactive stages are 3, 4, 5, and 6. All the other specifications
5.61 8.45
1.49 1.25
0.5 0.61
R
o/p ratio
0.81 0.86
1.63 1.94
remain the same in either case. Table 11 compares the results of these simulations. The addition of an extra stage in the rectifying zone results in a drop in the reactant conversions, which can be explained on the basis of the lower average reactive stage temperature and the liquid phase H2/CO ratio. Though it is expected that an increase in distillation effect will lead to a higher temperature at the column bottom, the same is not observed in Table 11. Such a conclusion may not be arrived at here owing to the considerable difference in the conversion of reactants and a change in product distribution. Case 2 has lower light gas content and higher diesel content. Simulations with more numbers of additional nonreactive stages in the rectifying section at the reflux ratio considered failed to converge. On the basis of the converged results, additional nonreactive stages in the rectifying zone show a potential in altering the product spectrum at the expense of a decrease in the net product yields but an increase in the liquid yields. Adding nonreactive stages in the stripping section of the column below all the reactive stages did not have significant effect on the reactant conversions and the product yields or any other variables with respect to the base case as shown in Table 12 and Table 13. It is interesting to note that the mass percent of diesel in the net product initially increases as the stripping section becomes longer and then starts decreasing with further increase in the number of stripping stages. As expected, however, there is a change in the composition profile on the stages as seen from Figures 14 and 15. With an increase in the separation between gasoline and diesel fractions, the fraction of gasoline on any stage increases with increasing height of the stripping section
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Table 12. Effect of Nonreactive Stripping Section stripping stages
Xco (%)
XH2 (%)
av reacn T (°C)
Qc (kW)
utility ratio
liquid H2/CO
vapor H2/CO
Tbottom (°C)
CO2 formed
0 (Base) 1 2 3 4 5 6
40.36 40.71 40.87 40.96 41.01 41.04 41.05
10.92 11.02 11.06 11.08 11.09 11.1 11.1
215.45 213.58 212.7 212.21 211.9 211.7 211.55
-5.32 -5.35 -5.36 -5.37 -5.37 -5.38 -5.38
0.55 0.55 0.55 0.55 0.55 0.55 0.55
2.34 2.33 2.33 2.33 2.33 2.33 2.33
2.45 2.45 2.45 2.45 2.45 2.45 2.45
216.27 217.02 217.54 217.75 217.78 217.7 217.58
0.0659 0.0665 0.0667 0.0669 0.0669 0.067 0.067
Table 13. Effect of Nonreactive Stripping Section on Product Distribution no. of stripping stages
light gas (mass %)
naphtha (mass %)
gasoline (mass %)
diesel (mass %)
wax (mass %)
net yield (kg/h)
liquid yield (kg/h)
R
o/p ratio
0 (base) 1 2 3 4 5 6
13.73 13.65 13.57 13.50 13.45 13.40 13.36
12.96 12.93 12.88 12.85 12.82 12.79 12.77
13.34 13.43 13.60 13.84 14.13 14.45 14.76
36.87 37.23 37.39 37.39 37.29 37.13 36.96
23.10 22.75 22.55 22.41 22.31 22.22 22.15
0.94 0.95 0.95 0.95 0.96 0.96 0.96
0.63 0.64 0.65 0.65 0.66 0.66 0.66
0.93 0.93 0.93 0.93 0.93 0.93 0.94
1.56 1.58 1.6 1.61 1.63 1.63 1.64
(Figure 14a). Further, Figure 14b shows the enrichment of gasoline mass fraction on stage 2 from 94% to 96%. With three stages in the stripping section, the diesel fraction is ∼60% (Diesel3) and further addition of stages in stripping section brings down its value slightly as seen in Figure 15. Thus, the addition of a nonreactive stripping section helps in enriching the products on some of the stages thereby providing a possibility of side-draw from those stages without altering the overall product distribution of the net product. The yields are not hampered either.
Figure 14. Effect of stripping section on the gasoline composition.
Figure 15. Effect of stripping section on the diesel composition.
Table 14. Side-Draw Specifications Used case
s5 (mol/h)
s9 (mol/hr)
s13 (mol/hr)
c1 (base) c2 c3 c4 c5 c6
0 0.25 0.34 0.25 0.25 0.3
0 0.25 0.25 0.34 0.25 0.3
0 0.25 0.25 0.25 1.45 0.5
The last option of placing nonreactive stages in between the reactive stages was also analyzed and the details are presented in Section 2 of the Supporting Information. No considerable difference is once again observed in the reactant conversion or product yields or any of the other parameters with the addition of the nonreactive section. However, the expected change in composition profiles of the component fractions is realized. For example, the addition of six nonreactive stages between the third and fourth reactive trays enriches the diesel fraction to ∼92% on stage 11. A judicious choice in the placement of the nonreactive stages thus creates potential in obtaining pure components as side-draws without adversely effecting the conversion or yields in most of the cases as demonstrated. Effect of Side-Duties. These effects were discussed in our previous work1 and the main conclusions that hold good in the present study as well are listed as follows: a. There is no direct relation between the stage temperature and reaction heat to be removed and these numbers are arrived at by estimating the amount of heat released on each stage as per the reaction extent. b. The compensating effects of temperature and the H2/CO ratio decide the conversion of the reactants and the product yields. With side heat-removals, the condenser duty comes down and heat is now available at a higher temperature and can be utilized more efficiently. c. The conversion of both the reactants remains unchanged by altering the side-stream rates. Net mass flows in all the cases are almost equal. There is no significant difference in the liquid phase H2/CO ratios on the reactive stages with variation in the side-draw flows. To study the effect of side-duties, a reflux ratio of 0.11 was used with the following duties: 0.3, -1.9, and -0.13 kW on the second, third, and fourth stages, respectively. It is to be noted that there is heat supply to stage 2 rather than removal. This was considered as the base case and the duties were so adjusted as to maintain a temperature close to 225 °C on all the reactive stages. These turned out to be 213.57, 224.67, 227.79, and
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Table 15. Effect of Side-Draws
c1 c2 c3 c4 c5 c6
Xco (%)
XH2 (%)
av reacn T (°C)
Qc (kW)
utility ratio
liquid H2/CO
vapor H2/CO
Tbottom (°C)
CO2 formed
35.54 33.52 33.01 33.32 33.24 33.06
9.47 8.88 8.74 8.82 8.81 8.76
212.69 218.29 219.16 219.72 223.73 220.62
-2.64 -2.46 -2.41 -2.44 -2.43 -2.41
0.54 0.54 0.54 0.54 0.54 0.54
2.31 2.31 2.31 2.31 2.33 2.32
2.42 2.42 2.42 2.42 2.43 2.42
227.67 234.83 235.63 236 245.89 237.91
0.0583 0.0551 0.0543 0.0547 0.0546 0.0543
case light gas (mass %) naphtha (mass %) gasoline (mass %) diesel (mass %) wax (mass %) net yield (kg/h) liquid yield (kg/h) c1 c2 c3 c4 c5 c6
11.20 10.56 10.81 10.79 11.41 11.09
13.27 12.92 13.12 13.10 13.50 13.31
16.95 15.87 15.83 15.87 15.88 15.86
32.74 32.99 32.53 32.61 32.25 32.23
225.24 °C on stages 2, 3, 4, and 5, respectively. Simulations were then repeated by varying the heat duty of the third stage between -2.1 and -1.2 kW. There was a significant alteration in the product distribution and conversion of the reactants with additional heat removal from the reactive trays, which is discussed further in the section on multiple solutions. Effect of Side-Draws. A refluxed rectifier with the following specifications is used in the present work to study the effect of side withdrawals: reflux ratio of 0.8; a total of 15 stages with reactions enabled on four stages, namely 2, 6, 10, and 14; feed entering on stage 15; stages 6 and 10 having side duties of -2.1 and -0.1 kW, respectively; side draws are provided on stages 5, 9, and 13, respectively. Six cases are analyzed with the flows listed in Table 14.
Figure 16. Effect of side draw on gasoline composition.
Figure 17. Effect of side draw on diesel composition.
Figure 18. Effect of side draw on wax composition.
25.84 27.66 27.71 27.63 27.01 27.51
0.82 0.77 0.76 0.77 0.77 0.76
0.58 0.53 0.52 0.53 0.52 0.51
R
o/p ratio
0.94 0.94 0.94 0.94 0.94 0.94
1.46 1.38 1.37 1.38 1.39 1.38
Convergence failed at the reflux ratio considered when one of the flow rates in each case listed in Table 14 was increased further, and hence, each set corresponds to the best possible case. The results of the simulations are presented in Table 15. Among the side-draw cases, the variation in their flow rates does not seem to considerably effect the conversion of the reactants, the yields and any other variables listed in Table 15, except for the column bottom temperature. The advantage foreseen with the side-draw arrangement is to selectively remove streams enriched in a single fraction like gasoline and diesel and reduce the load on further downstream processing. Figures 16 to 18 depict the composition profiles inside the column in the cases considered for each of the fractions: gasoline, diesel, and wax. The gasoline fraction is highest (∼97%) on stage 2 in Figure 16 in all the cases. Since the first side draw is from stage 5, it does not affect the gasoline composition on the upper stages. However, the gasoline fraction on the stages below stage 5 is decreased considerably with a side-draw removal compared to the base case in Figure 16 and all the cases with side removals perform in a similar manner. In the absence of a side-draw, the peak in the diesel fraction is on stage 13 (∼85%) as shown by the base case in Figure 17. In all the cases with a side-product removal, the peak is more than 85%. This diesel peak reaches a maximum of ∼93% on stage 12 in case 3 corresponding to Table 15. At a high rate of withdrawal from stage 13 corresponding to case 5 in Table 15, the slippage of diesel fraction into the bottoms stream is the least (∼5%) at the expense of a notso-rich diesel composition on stage 13 (∼89% on stage 11 compared to ∼63% on stage 13). Compared to the base case, the presence of side draws helps in increasing the wax composition on the bottom stages as shown in Figure 18. Nearly a pure wax stream (∼95%) can be obtained as the bottom product for case 5 when the rate of withdrawal from stage 13 is the maximum. In all other cases, it is in the range of 56-58% except in case 6 where the wax fraction in the bottom product improves to ∼65%. There is quite an improvement over the base case which gives a bottom product having 44% waxes only by mass. Thus, a judicious choice of the location and rate of side-draw can significantly improve the performance. Significant difference is also observed in the temperature profile and average o/p ratio on the stages in the column for the different cases. As shown in Figure 19a, the tray temperatures are higher than the base case profile in the presence of side-draws. This is so as the bubble point temperature of the stages below the side-draw tray increases owing to removal of more volatile components through the exit stream. The average o/p ratio increases compared to the base case from stages 7 and
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Figure 19. (a) Effect of side draw on temperature profile, (b) effect of side draw on o/p ratio.
Figure 20. (a) Multiple solutions in conversion of CO, (b) multiple solutions in the average reactive stage temperature.
below and is lower than the no side-draw case on the upper stages (Figure 19b). This is an interesting feature and needs more investigation since it can possibly help in manipulating the o/p ratio for each of the product side-draws.
temperature. The mass percentages at the high temperature solution are 8.38, 10.45, 16.05, 32.48, and 32.64, respectively, for the light gas, naphtha, gasoline, diesel, and wax fractions. The corresponding values for the low temperature branch are 6.98, 9.26, 12.94, 34.55, and 36.27, respectively. Figure 20 indicates that the solution at the jumps depends on the solution branch that the initial guess lies in. Though these differences are not very significant, it proves the existence of multiple solutions and the need to carefully analyze the results of the steady-state simulation. As pointed out earlier, the region where the sensitivity analysis fails needs to be more carefully analyzed to check for the existence of multiple or no solutions in the case of other parameters like reflux ratio, condenser temperature, etc.
5. Multiple Solutions Multiple steady-states (MSS) for MTBE synthesis in RD were reported for the first time by Jacobs and Krishna9 using Aspen Plus. An extensive review on the nonlinear dynamics studies in RD was done by Kienle and Marquardt.10 During the course of sensitivity analysis, it is a common practice to vary the parameter of interest in both increasing and decreasing directions and observe its effect on other process variables. Sometimes, a unique solution is not obtained in both the directions indicating the possibility of multiple solutions. A base case for simulations involving side-duty had the following specifications: reflux ratio of 0.11; side duties of 0.3, -1.9, and -0.13 kW on the second, third, and fourth stages, respectively. No convergence in simulations was achieved as the side duty of the third stage was increased beyond -1.2 kW and decreased below -2.1 kW. For simulations in between the range of -1.2 and -2.1 kW as the third stage side-duty, different solutions were obtained in some of the cases for the increasing and decreasing branches. The results in the case of CO conversion and the average reactive stage temperature are presented in Figure 20a and Figure 20b, respectively. The unsteady-state solution that completes the “S-shape” could not be achieved using the steady-state simulations. It can be realized by the use of continuation algorithms or through dynamic simulations. Figure 20a shows two branches for the CO conversion, a higher conversion branch that corresponds to the higher average reactive stage temperature in Figure 20b and a lower conversion solution that matches with the lower average reactive temperature. The fraction of liquid yield as a part of the total yield is almost the same for both the branches. The net and liquid yields are 0.99 and 0.77 kg/h at the high conversion level; and 0.76 and 0.6 kg/h at the lower conversion limit. There is a slight change in the product distribution too with the diesel and wax fractions being higher at the low
6. Conclusion In continuation with our earlier work (DOI 10.1021/ ie801887m),2 parametric studies were performed in this work using a “refluxed rectifier” configuration without a reboiler for FTS in RD. The flexibility that is possible for FTS in RD by variation of operating parameters like reflux ratio, condenser pressure, etc. and of design parameters like catalyst loading, number of nonreactive stages, etc. is demonstrated. In all the cases, the effects on reactant conversion, yields, and product distribution are explained on the basis of the observations made. An increase in reflux ratio leads to an increase in conversion at the expense of liquid yields. The reactive stage temperature and the product distribution are altered too. Increasing the condenser pressure led to higher liquid yields with a heavier product spectrum, but reduced the conversion. The condenser temperature did not seem to have any effect on the conversion or product distribution. Though the feed temperature is decided by the overall process conditions, it is observed that a hotter inlet feed shifts the product selectivity toward diesel and wax with a trade-off in the conversion and net product yields. The advantage of using side-draws to selectively remove products rich in a single component, and the use of side-duties to maintain a desired temperature on the reactive stages in the column is also illustrated through examples. Owing to the number of
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parameters one can play with, it is difficult to arrive at an optimum configuration by variation of a single parameter. Using the knowledge from the single parameter sensitivity analysis, further simulations and optimization are necessary to suggest the best possible configuration. Use of algorithms that permit homotopy continuation need to be tried to solve the convergence problems that are faced. Such algorithms are also a better tool to investigate the presence of multiple steady states (MSS) reported in this study. The effect of using a feed containing CO2 on the performance of RD is studied and in comparison with the conventional reactors, RD is found to perform reasonably well even in this case. Supporting Information Available: Effect of condenser temperature; effect of nonreactive stages; composition profiles. This material is available free of charge via the Internet at http:// pubs.acs.org. Nomenclature and Abbreviations RD ) reactive distillation RR ) (reactive) refluxed rectifier FTS ) Fischer-Tropsch synthesis LTFT ) low temperature Fischer-Tropsch process HTFT ) high temperature Fischer-Tropsch process FBR ) fixed-bed reactor CSTR ) slurry reactor WGS ) water-gas shift reaction o/p ) olefin-to-paraffin Qc ) condenser duty (kW) Qrt ) reactor duty (kW) Xco ) conversion of CO XH2 ) conversion of H2 Xco+co2 ) overall conversion of carbon oxides (CO and CO2) av reacn T ) average temperature of the reactive stages (°C)
Tbottom ) bottom temperature in the RD/RR column (°C) R ) chain-growth probability parameter β ) olefin readsorption parameter
Literature Cited (1) Srinivas, S.; Malik, R. K.; Mahajani, S. M. Feasibility of Reactive Distillation for Fischer-Tropsch Synthesis. Ind. Eng. Chem. Res. 2008, 48, 889–899. (2) Srinivas, S.; Malik, R. K.; Mahajani, S. M. Feasibility of Reactive distillation for Fischer-Tropsch Synthesis. 2. Ind. Eng. Chem. Res. 2009; DOI 10.1021/ie801887m. (3) Aspen Plus User Manual, Aspen Plus version 2004.1; Aspen Technologies Inc.: Cambridge, MA, 2004. (4) Wang, Y. N.; Ma, W. P.; Lu, Y. J.; Yang, J.; Xu, Y. Y.; Xiang, H. W.; Li, Y. N.; Zhao, Y. L.; Zhang, B. J. Kinetics Modeling of FTS over an Industrial Fe-Cu-K Catalyst. Fuel 2003, 82, 195–213. (5) Srinivas, S.; Khadse, A.; Aghalayam, P.; Malik, R. K.; Mahajani, S. M. Effect of Gasification Conditions on FT Fuels and Power Production; PCC: Pittsburgh, PA, 2008. (6) Kim, J. S.; Lee, S.; Lee, S. B.; Choi, M. J.; Lee, K. W. Performance of Catalytic Reactors for the Hydrogenation of CO2 to Hydrocarbons. Catal. Today 2006, 115, 228–234. (7) Riedel, T.; Claeys, M.; Schulz, H.; Schaub, G.; Nam, S. S.; Jun, K. W.; Choi, M. J.; Kishan, G.; Lee, K. W. Comparative Study of FTS with H2/CO and H2/CO2 Syngas Using Fe and Co Catalysts. Appl. Catal. A 1999, 186, 201–213. (8) Agarwal, V.; Thotla, S.; Mahajani, S. M. Attainable Regions of Reactive Distillation - Part I. Non-azeotropic Single Reactant Systems. Chem. Eng. Sci. 2008, 63 (11), 2946–2965. (9) Jacobs, R.; Krishna, R. Multiple Solutions in Reactive Distillation for MTBE Synthesis. Ind. Eng. Chem. Res. 1993, 32, 1706–1709. (10) Kienle, A. Marquardt, W. Nonlinear Dynamics and Control of Reactive Distillation Processes. In ReactiVe Distillation: Status and Future Directions; Sundmacher, K., Kienle, A., Eds.; Wiley VCH: Germany, 2003.
ReceiVed for reView December 8, 2008 ReVised manuscript receiVed February 25, 2009 Accepted March 25, 2009 IE801888V
