Biodiesel Production by Esterification of Hydrolyzed Soybean Oil with

Article pubs.acs.org/IECR

Biodiesel Production by Esterification of Hydrolyzed Soybean Oil with Ethanol in Reactive Distillation Columns: Simulation Studies Guilherme Duenhas Machado,† Fernando Luiz Pellegrini Pessoa,§ Marcelo Castier,∥ Donato A. G. Aranda,§ Vladimir Ferreira Cabral,‡,* and Lúcio Cardozo-Filho† †

Departamento de Engenharia Química, Universidade Estadual de Maringá, Av. Colombo 5790, Maringá PR 87020-900, Brazil Departamento de Engenharia de Alimentos, Universidade Estadual de Maringá, Av. Colombo 5790, Maringá PR 87020-900, Brazil § Departamento de Engenharia Química, Escola de QuímicaUniversidade Federal do Rio de Janeiro, Av. Horácio Macedo 2030, Rio de Janeiro RJ 21941-909, Brazil ∥ Chemical Engineering Program, Texas A&M University at Qatar, P.O. Box 23874, Doha, Qatar ‡

ABSTRACT: Biodiesel conventional production process, by alkaline transesterification reaction, have disadvantages such as complex products separation and high feedstock costs. In this regard, production of biodiesel by esterification of fatty acids into a reactive distillation column has proved to be promising for overcoming some of these drawbacks. However, only simulation works that consider only one type of fatty acid reagent are available in the literature, and not simulations based on a real fatty material for this process. In this way, this work presents steady-state computational simulations of fatty acid esters (biodiesel) production in a reactive distillation column by esterification reaction of a new feedstock that represents the fatty acids composition of the soybean oil (hydrolyzed soybean oil) with anhydrous ethanol. Sensitivity analyses showed that the best operating conditions were the minimum reflux ratio of 0.001 and 15 theoretical stages. As to thermal analysis, it was noted that the process is optimized by increasing the energy consumption of reagent instead of the reboiler. The low thermal load on this equipment can be used in order to avoid exposure of the bottom product at elevated temperatures. Conversions close to 99% were possible with the proper choice of these operating conditions. The results show the technical feasibility of this process, and such data can be useful for the design of biodiesel processes.

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INTRODUCTION Since commercial biodiesel production started in the 1990s in some European countries,1 this biofuel, as described by the U.S. EPA,2 has spread around the world, achieving significant production in Latin America, Southeast Asia, and the USA and becoming a promising opportunity in some African countries.3 Few technical changes have occurred in the transesterification process in the last decades.4 In fact, yields higher than 99% are usually obtained mainly using low acidic feedstocks such as rapeseed and soybean oils.5 Raw materials with larger free fatty acids content (FFA) need to be pretreated.6 Usually, a preesterification step is carried out with corrosive acid catalysts such as sulfuric acid.7 This way, FFA is reduced and nondesirable soap production decreases in the transesterification step. However, in addition to chemicals extra costs, sulfate ions are responsible for incrustations that increase operating costs for maintenance. Palm oil biodiesel is a typical case for a preesterification process, since FFA in crude palm oil is usually about 5%.7 The same treatment is necessary for other palm crops such as “macauba” (Acrocomia aculeata)8 and wet environment oils such as algae oils.9 Instead of reducing the FFA in crude vegetable and animal oils, one could convert those raw materials in fatty acids by means of a noncatalytic hydrolysis.10 This route produces a very pure glycerol easily separated from fatty acid obtained in a top phase. After this step, fatty acids can be transformed into esters during an esterification reaction. © XXXX American Chemical Society

Reactive distillation columns can contribute to the search for a favorable alternative technology for biodiesel production. Its advantages, such as those obtained in the Eastman Chemicals11,12 case study process for methyl acetate synthesis, motivate further work with this approach. Papers such as those of Machado et al.,13 Smith et al.,14 and He et al.15 and especially those of Kiss et al.16−22 show that reactive distillation columns have potential for application in biodiesel production. In the computational simulations presented in these studies, the authors consider systems in which there is only one reactive fatty acid, which are known not to exist isolated in nature, since vegetable oils contain many kinds of fatty acids. In this way, the proposal of models that can adequately represent a mixture of biodiesel esters formed from vegetable oils is needed for better representation of the process. Such an issue is the main motivation for this work. Therefore, this study aims at contributing to this topic by presenting new computational simulations of biodiesel production from a hydrolyzed soybean oil heterogeneously catalyzed esterification reaction in reactive distillation columns. The modeling used by Machado et al.13 for systems with only one reactive fatty acid is extended herein using a Received: March 12, 2013 Revised: June 11, 2013 Accepted: June 12, 2013

A

dx.doi.org/10.1021/ie400806q | Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

Industrial & Engineering Chemistry Research

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component i in the vapor leaving stage j, the flow rate in the vapor sidestream, and the flow rate to the next stage, respectively. Fi,j denotes the flow rate of component i in the feed stream to stage j, νi,k denotes the stoichiometric coefficient of component i in reaction k, ξk,j represents the extent of reaction k in stage j, and nr represents the number of independent chemical reactions. We assume the streams leaving each stage are in phase equilibrium. The fugacity coefficients of each component in the vapor phase and of the saturated vapor at the temperature of each stage are assumed to be equal to 1, and the Poynting factor correction is neglected. In eq 2, xIi,j and xi,jII are the mole fractions of component i in the vapor and liquid streams leaving stage j, Pj denotes the pressure of stage j, Psat i,j is the saturation pressure of component i on stage j, and γIIi,j is the activity coefficient of component i in the liquid phase leaving stage j. We assume the chemical reactions occur in the liquid phase only. In eq 3, vIIj denotes the molar volume of an ideal liquid solution in stage j, kk,j is the kinetic constant of reaction k in stage j, αi.k represents the kinetic order of component i in reaction k. In the energy balance (eq 4), (Zj + 1)HIj , (Rj + 1)HIIj , and HFj are the enthalpy flow rates of the vapor and liquid streams leaving stage j and the enthalpy flow rate of the feed stream to stage j, respectively. In addition, Qj denotes the rate of heat addition to stage j and Ej is the ratio between the molar flow of vapor and liquid leaving stage j. Equation 5 specifies the behavior of the condenser and reboiler, based on variable Ej, defined as the ratio between the molar flow of vapor and liquid leaving the stage j. In this study, values of E1 ≠ 0 and EN = 0 are used in the extremes of the column; that is, a partial reboiler and a total condenser are considered in all cases. The value of Ej is calculated for the internal stages of the column (from 2 to N − 1). The equations and unknowns are organized as described in detail by Alfradique and Castier.24 The formulation adopted here uses the Newton−Raphson method to solve the set of nonlinear equations shown previously (eqs 1−5). The Thermath package was used to obtain Fortran subroutines that implement these equations and their derivatives with respect to the process variables and the excess Gibbs energy model used in the simulation. A Fortran program with about 10800 lines of code was used in the simulations of this work. In our research group, we usually develop the programs instead of using commercial software. Our computational code and procedures were validated in a previous publication,13 where the results obtained by us were compared with experimental data and simulation results using the Aspen Plus software. Thermodynamic Modeling. Most simulation studies on biodiesel production in reactive distillation columns13,16−22 consider the esterification reaction of a pure fatty acid with an alcohol. However, from a practical viewpoint, pure fatty acids are not available in nature, requiring complex processes to achieve purification, which increases reactant costs. Moreover, these fatty acids are constituents of triglycerides present in different types of vegetable/edible oils, such as corn, canola, sunflower, palm, soybean, and others. For convenience, in this paper was used a representative mixture of the major fatty acids of soybean oil, widely available in Brazil and often used for biodiesel production. In this case, soybean oil must pass initially by a hydrolysis process. So, the triglycerides present in the oilseed react with

pseudocomponent approach to represent a mixture of fatty acids of soybean oil. In the simulations, the formulation used assumes steady-state operation and the reaction rates are explicitly considered in the model of each stage. Sensitivity analyses are performed in order to verify the best operating conditions for the reactive distillation column.

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METHODOLOGY The methodology used here was adapted from the work of Machado et al.13 and Alfradique and Castier,24 who consider there is no chemical equilibrium in stages but assume steady state operation. In this case, reaction rates are considered explicitly in the model of each stage. The general stage scheme used in this work is shown in Figure 1:

Figure 1. Schematic of each theoretical stage along the reactive distillation column.

From this stage scheme, the model equations for the entire reactive distillation column are formulated as follows: Mass balance (mb) f im, j = (R j + 1)niII, j + (Zj + 1)niI, j nr

−

(niII, j + 1

+

niI, j − 1

+ Fi , j +

∑ vi ,jξk ,j) = 0 k=1

(1)

Phase equilibrium (pe) f ieq = ln(xiI, jPj) − ln(xiII, jγi II, j Pisat ,j ) = 0 ,j

(2)

Reaction rate (rr) ⎛ x II ⎞ i,j ⎟ − ln ξk , j = 0 II ⎟ v ⎝ j ⎠

nc

f kr, j = ln kk , j =

∑ αi , k ln⎜⎜ i

(3)

Energy balance (eb) f jh = (R j + 1)HjII + (Zj + 1)HjI − (HjII+ 1 + HjI− 1 + H Fj + Q j) = 0

(4)

Vapor/Liquid relationship (vlr) nc

nc

f jvl = (Zj + 1) ∑ niI, j − Ej(R j + 1) ∑ niII, j = 0 i=1

1)nIIi,j,

i=1

RjnIIi,j,

(5)

ni,jII

In eq 1, (Rj + and represent the molar liquid flow rate of component i leaving stage j, the flow rate of the liquid sidestream, and the flow to the next stage, respectively. (Zj + 1)nIi,j, ZjnIi,j, and ni,jI represent the molar flow rate of B



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water to produce glycerol and fatty acids. Such fatty acids formed are known as hydrolyzed soybean oil. Here, hydrolyzed soybean oil is defined as a pseudocomponent that represents a mixture of the following fatty acids with their respective molar compositions: 58.2% of linoleic acid, 25.1% of oleic acid, and 16.7% of palmitic acid. In the same way, biodiesel is defined as a pseudocomponent that represents a mixture of fatty acid esters with the following composition: 58.2% of ethyl linoleate, 25.1% ethyl oleate, and 16.7% of ethyl palmitate. In this way, some considerations regarding the thermodynamic properties of these pseudocomponents are required. Excess Gibbs Energy (GE) Model. The liquid phase was modeled using the group activity model UNIFAC (UNIQUAC functional-group activity coefficients) Dortmund26 equation. We took the weighted average of the three components (fatty acids or fatty acid esters) that constitute the pseudocomponents to compute the matrix of subgroups in the UNIFAC method. The values used here are shown in Table 1.

Table 2. Evaluation of the Antoine Equation Parameters of Hydrolyzed Soybean Oil

CH3

CH2 14 14 12 13 13 11 12.84 11.84

CH 2 0 4 2 0 4 2.83 2.83

COOH 1 1 1 0 0 0 1 0

ln P sat

1

hydrolyzed soybean oil

fi A B C

0.582 14.44 7962.69 −72.25

0.251 11.61 5628.24 −130.85

0.167 13.99 7546.02 −80.59

15.18 8203.37 −69.69

∑ xi ∫ i=1

T

Tref

c p,L i dT + hE

nc

hV =

CH2COO

∑ yi (Δhivap + ∫

T

Tref

i=1

(8)

c p,L i dT )

(9)

Δhivap

where is the molar enthalpy of vaporization of component i in the system, hE is the molar excess enthalpy, and cLp,i is the molar specific heat of component i in the liquid phase. The reference temperature (Tref) used was 298.15 K. The molar enthalpy of vaporization was calculated using the Clausius−Clapeyron equation as follows:

0 0 0 1 1 1

Δhivap = RT 2

0 1

d ln P sat dT

(10)

Kinetic Model. The esterification of hydrolyzed soybean oil (1) with ethanol (2) producing biodiesel (3) and water (4) follows the equation: hydrolyzed + ethanol ⇔ biodiesel + water (1)

(2)

(3)

(4)

(11)

The esterification reaction was considered to be of first order with respect to the concentrations of hydrolyzed soybean oil and ethanol, while the inverse reaction (hydrolysis) follows a first order kinetic with respect to the concentrations of biodiesel and water. From these assumptions, we propose the following pseudohomogeneous model:

(6)

3

∑ fi Pi = f1 P1 + f2 P2 + f3 P3

linoleic acid (3)

nc

In this case, the pseudocomponent vapor pressure was not evaluated by the weighted average of the parameters A, B, and C of the Antoine equation. We generated pseudodata of vapor pressure at certain temperatures, using the following equation: Pisat =

palmitic acid (2)

hL =

Vapor Pressure. The Antoine equation27 was used to calculate the vapor pressure B =A− T+C

oleic acid (1)

fatty acids. The value of the heat capacity of the liquid was calculated for the temperature 450 K. For biodiesel, due to lack of data in the literature, we used only the data of the methyl oleate. The values of specific heat of liquids (cLp,i) and the molar volumes of liquids were obtained from NIST28 and DIPPR29 databases. Enthalpies. The molar enthalpies of the liquid (hL) and vapor (hV) were calculated using the following equations:

Table 1. Matrix Groups Weighted Average of the Dortmund UNIFAC Model for the Hydrolyzed Soybean Oil (1) and Biodiesel (3) Component oleic acid 1 palmitic acid 1 linoleic acid 1 ethyl oleate 2 ethyl palmitate 2 ethyl linoleate 2 Pseudocomponent hydrolyzed soybean oil 1 biodiesel 2

parameter

r=

(7)

where f i is the mole fraction of component i in the pseudocomponent and Pi is the vapor pressure of component i calculated at a given temperature. These calculations were made for different temperatures between 340 and 500 K. So, the pseudodata generated were used to obtain the Antoine parameters of the pseudocomponents. Table 2 presents the Antoine parameters obtained here for the hydrolyzed soybean and biodiesel. The Antoine parameters for pure components were obtained from NIST (National Institute of Standards and Technology)28 and DIPPR (Design Institute for Physical Properties)29 databases. Molar Volume and Heat Capacity of Liquids. The molar volumes and heat capacity of the liquid for hydrolyzed soybean oil were also calculated by the weighted average of the three

1 1 dni = k1C1C2 − k −1C3C4 mcat v dt

(12)

The constants k1 and k−1 in eq 12 obey the Arrhenius equation ⎛ −24117 [J/mol] ⎞ k1 = 16.13 exp⎜ ⎟ (L/(gcat·min · mol)), T ⎝ ⎠ RT (K)

(13)

⎛ −24117 [J/mol] ⎞ k −1 = 0.72 exp⎜ ⎟ (L/(gcat·min · mol)), T ⎝ ⎠ RT (K)

(14)

The rate constants of eq 12 were fitted from the experimental kinetic data of Rocha et al.,30 who studied esterification of C



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From Figure 3, the removal of the desired product, biodiesel, in the bottom is made with considerable purity (96.54%). If we consider the mass fraction, we will have purity above 99% in biodiesel. In this case, the remainder ethanol, which is added to the column in excess, and water produced in the esterification reaction are removed at the top, as desired. The hydrolyzed soybean oil is almost entirely consumed through the column, and its profile concentration shows considerable decline close to stage 13. In this stage, the hydrolyzed soybean oil feeding occurs. In this way, the esterification reaction is favored close to the feed location of hydrolyzed soybean oil, as can be seen in Figure 4. Figure 4 also shows that the reactive zone (6−12 stages) presents higher temperatures than neighboring stages where the reaction does not occur, except the reboiler. The highest temperature in the reactive zone is close to the feed location of hydrolyzed soybean oil, which is fed at a temperature of 480.15 K. Ethanol is fed on stage 6. It ascends the upper stages of the column as vapor, while the hydrolyzed soybean oil in liquid form descends from stage 13. In this case, the temperature of the hydrolyzed soybean oil is significantly higher than the ethanol temperature. In this way, the feed temperature of the hydrolyzed soybean oil governs the temperature profile in this section between stages 8 and 13. The temperature difference between stages 1 (reboiler) and 2 is considerable. In the reboiler, temperature reaches 535.95 K, while in the second stage the temperature is 400.24 K. This temperature difference occurs because stage 1 (reboiler) has a considerable heat transfer rate and the bottom product is mostly liquid and does not return to stage 2, being removed from the bottom of the column. The higher temperature in the reboiler can degrade the biodiesel removed in the bottom product. Between stages 2 and 6, the temperature remains stable. In this section, the separation between reactants and products takes place, and there is no feed of reactants or heat transfer with the surroundings. Therefore, the temperature barely changes in this section. Sensitivity Analysis. This section presents an assessment of the effect of some important parameters on the operation of the reactive distillation column in this example. The following parameters were analyzed: the reflux ratio in the condenser, reboiler heat transfer rate, and hydrolyzed feed temperature. For the sensitivity analysis of the reflux ratio, a range from 0.001 to 0.7 was used. The minimum value of 0.001 was used for allowing numerical convergence, which does not occur for zero reflux ratio in our implementation. For the sensitivity analysis of the reboiler heat transfer rate, values between 0.01 and 57.2 kJ/min were used, with the first value being almost as if the reboiler were turned off and the second value referring to the maximum heat duty limited by the temperature to which the bottom products are exposed. In the case of the hydrolyzed feed temperature, we used values between 390 and 480 K, since this temperature range leads to reaction sector temperatures close to those at which the kinetic constants were obtained. In each sensitivity analysis performed here, all parameters other than the analyzed parameter have the values shown in Table 3. Reflux Ratio. Figure 5 shows the biodiesel mole fraction in the liquid phase along the reactive distillation column as a function of the reflux ratio, and Figure 6 presents the temperature as a function of the reflux ratio. From Figure 5, the increase in the reflux ratio causes a decrease in the biodiesel purity at the bottom of the column. This fact can be explained by the increase in the water concentration, which is an esterification reaction product, in the

hydrolyzed soybean oil using a niobium oxide catalyst. Niobium oxide is very stable under esterification reactions, since its acid sites are not deactivated in contact with water, as are most heterogeneous acid catalysts. It has a selectivity of approximately 100% in esterification reactions. This type of catalyst is inactive for alcohol dehydration reactions under esterification conditions; thus, no important parallel reaction is observed. Leaching of the niobium element is not detected in reaction media, and this is not an important deactivation mechanism under esterification reaction conditions. Several publications of Aranda and collaborators have dealt with this subject.31−33

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RESULTS AND DISCUSSION This section presents new computational simulation results of the esterification reaction between a fatty acid (hydrolyzed soybean oil) and ethanol in a reactive distillation column. The reactive distillation column setup shown in Figure 2 is used in all simulations performed in this paper.

Figure 2. Schematic of the reactive distillation column of this work.

Preliminary studies indicated 15 stages are sufficient to achieve suitable conversion. Therefore, the simulated column has 15 stages: 1 partial reboiler, 13 adiabatic plates (7 reactive), and 1 total condenser. The specifications of the feed are presented in Table 3. Figure 3 shows the mole fraction profiles in the liquid phase of each component, and Figure 4 exhibits the temperature and extent of the esterification reaction along the column. Table 3. Specifications of the Reactive Distillation Column variable/component pressure stages condenser reboiler reflux rate reactive sector catalyst feed

hydrolyzed soybean oil anhydrous ethanol

specifications all stages total partial condenser reboiler stages niobium oxide 1.0132 bar 0.366 (mol/min) T = 480.15 K 0.437 (mol/min) T = 351.15 K

1.0132 bar 15 stage 15 stage 1 0.001 0.14 6−12 40 g per stage stage 13 stage 6

D



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Figure 3. Liquid phase composition along the reactive distillation column.

Figure 4. Temperature profile and extent of reaction along the reactive distillation column.

reactive zone. Such a phenomenon allows the occurrence of the hydrolysis reaction, which is not desirable. Figure 6 shows that the increase in the reflux ratio in the condenser decreases the temperature profile along the reactive distillation column. The product obtained at the column top needs to be condensed before it returns to the column. The increase in the reflux ratio increases the amount of liquid that returns to the column. This return of liquid at a lower temperature causes a decrease of the temperature profile of the column. This can be best observed in the temperature of the lower stages of the column. Regarding the temperature of the reboiler, for example, for the reflux ratio 0.001 the temperature is 451.92 K, whereas with a reflux ratio of 0.7 the temperature is 379.57 K. The decrease in the temperature profile of the column does not favor the occurrence of an esterification reaction, since the influence on the reaction temperature is governed by Ahrrenius law. In this way, for this case, lower values of the reflux ratio are better choices for operating the reactive distillation column.

Figure 5. Composition surface of biodiesel (3) in the liquid phase of the column simulated: Effect of reflux ratio.

E



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Figure 6. Temperature surface along the reactive distillation column. Effect of reflux ratio.

Figure 8. Temperature surface along the reactive distillation column in sensitivity analysis: Effect of reboiler.

Reboiler Heat Transfer Rate. Figure 7 shows the influence of the reboiler heat transfer rate on the composition surface of

the column bottom is approximately 397 K. At the other extreme, when the upper limit, 57.2 kJ/min, is employed, the temperature reaches around 599 K. In the last case, such a temperature can promote thermal degradation of biodiesel and also increase the cost of operation of the reactive distillation column, as more energy is required. Figure 9 shows the conversion of the hydrolyzed soybean oil as a function of the reboiler heat transfer rate. The increase in the reboiler heat-transfer rate generates an increase of just about 2.3% conversion of the hydrolyzed soybean oil. Therefore, in terms of conversion, this procedure is not recommended. Thus, there is a need for another variable in order to optimize the process of the reactive distillation column. Energetically, it is possible to vary the temperature of the feed streams of reactants. Feed Temperature. Another key factor for operating a reactive distillation column is the feed temperature of the reactants. In this analysis, we considered only the feed temperature of the hydrolyzed soybean oil as a parameter. Values between 390 and 480 K were used in this assessment. In all simulations, the ethanol stream was fed at temperature of 351.15 K. This temperature has not been modified since the methodology of this study is considered food in just one phase (liquid). Figure 10 shows the biodiesel surface composition in the liquid phase, and Figure 11 shows the hydrolyzed soybean oil conversion as a function of the feed temperature of hydrolyzed . In Figure 10, the increase in the feed temperature of the hydrolyzed soybean increases the purity of biodiesel removed at the bottom of the reactive distillation column. In other words, there is an increase of conversion in the system due to the additional heat load promoted by the hydrolyzed soybean oil feeding. This fact can be better seen in Figure 11. Such results suggest that increasing the feed temperature of hydrolyzed soybean oil can be a better optimization strategy than increasing the reboiler heat transfer rate. This inference can be noted when Figures 9 and 11 are compared. First, the hydrolyzed soybean oil conversion exhibits a wider range in Figure 11 (∼70−95%) than in Figure 9 (∼93−95%). The

Figure 7. Composition surface of ethanol (2) in the liquid phase of the column simulated: Effect of reboiler.

ethanol (2) in the liquid phase along the column, while Figure 8 presents the effect of the reboiler heat transfer rate on the temperature surface along the reactive distillation column. The increase in the reboiler heat transfer rate has more influence on the most volatile component, ethanol, as seen in Figure 7. The increase in the reboiler heat transfer rate increases the energy available for this component that is fed in excess into the column. This situation allows ethanol to change to the vapor phase. Consequently, the rest of the energy added by the reboiler is transferred to the column as sensible heat, increasing the temperature of the bottom products, as can be seen in Figure 8. In this case, a very important issue arises, which is the relationship between the reboiler heat transfer rate and the temperature to which the bottom product is exposed at this stage. Figure 8 shows a remarkable change in temperature. When the lower value, 0.01 kJ/min, is used, the temperature in F



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Figure 9. Conversion of hydrolyzed soybean oil as a function of the reboiler heat transfer rate: Effect of reboiler.

manipulation of the feed temperature seems to have more positive impact in the operation of the column. Second, the temperatures reached in the bottom product are considerably lower when the feed temperature of the hydrolyzed soybean oil is increased (364.56−451.92 K). When the reboiler heat transfer rate is increased, the reboiler temperature has values between 397.0 and 598.8 K. In this last case, biodiesel removed from the bottom is exposed to higher temperatures that may degrade the product. Therefore, as a matter of product quality (which is a fuel) and low return on conversion into biodiesel, the optimization strategy for increasing the thermal load on the reboiler is not attractive. The addition of energy in the feed stream of soybean hydrolyzed proves fruitful. Simulation results with unpublished grease fatty acids of soybean oil were consistent compared to previous work. The

Figure 10. Composition surface of biodiesel (3) in the liquid phase of the column simulated: Effect of feed temperature.

Figure 11. Conversion of hydrolyzed soybean oil as a function of the hydrolyzed feed temperature: Effect of feed temperature. G



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Fj = total molar flow rate of the feed stream in stage j fm i,j = mass balance function of component i in stage j f eq i,j = phase equilibrium function of component i in stage j f hj = energy balance function at each stage f vlj = function relating the liquid and vapor streams f rk,j = chemical equilibrium function f i = mass fraction of component i in soybean oil hL = molar enthalpy of liquid hV = molar enthalpy of vapor hE = molar enthalpy excess HIj = total enthalpy of stream I at stage j HIj+1 = total enthalpy of stream I at stage j + 1 HFj = total enthalpy flow rate of feed stream to stage j kk,j = rate constant of reaction k in each stage j W = catalyst mass per reactive stage n = number of components N = reactive distillation column number of stages nIi,j = molar flow rate of component i in stream I of stage j nIIi,j = molar flow rate of component i in stream II of stage j Psat i,j = saturation pressure of component i in stage j Pj = pressure at stage j Qj = heat duty to stage j R = universal gas constant Rj = liquid side stream fraction at stage j Tj = temperature at stage j VIIj = liquid molar volume at stage j xIi,j = mole fraction of component i in stream I of stage j xIIi,j = mole fraction of component i in stream II of stage j Zj = vapor side stream fraction at stage j

sensitivity analysis of the feed temperature was useful as an alternative to optimize the reactive distillation column. It is suggested that the use of the reboiler can be minimized or even eliminated, as found in some studies in the literature. However, these studies introduce the need to use additional processes to recover the reactants at the top and bottom of the reactive distillation column. The strategy for increasing the thermal load in the feed stream of soy hydrolyzed can eliminate these additional processes to recover as found in the literature.

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CONCLUSIONS In this work, a process for biodiesel production by esterification of a hydrolyzed soybean oil with anhydrous ethanol in a reactive distillation column was simulated computationally. The results were consistent with previous studies and the available literature. Initially, the simulation conditions used allow almost complete conversion of the hydrolyzed soybean oil; however, remarkably extreme operation conditions were necessary in this case. So, a sensitivity analysis permitted determination of suitable conditions for column operation. In this situation, conversions above 98% were obtained. Sensitivity analyses showed operating conditions optimized for the column: minimum reflux ratio of 0.001, reboiler off (reactive absorption), temperature of 480 K in the feed stream of hydrolyzed soybean oil (stage 13), and 15 theoretical stages. The reboiler heat transfer rate analysis together with the study of the feed temperature of the hydrolyzed oil soybean showed that a good way to optimize the operation of a reactive distillation column is to increase the temperature of the fatty acid feed and not to increase the reboiler heat transfer rate. From an economic viewpoint, this operation may have advantages in fixed cost, since it minimizes the use of the reboiler and does not imply the use of additional equipment to the process. From the literature, the strategy to optimize the reactive distillation column is minimizing the reboiler use by the insertion of additional units to recover the unreacted alcohol at the top and the unreacted fatty acid in the bottom of the column. On the other hand, the operational costs also can be reduced by using fatty feedstock with high acidity as residual soybean oils. However, economic studies must be performed to prove that this process is economically feasible. The techniques and procedures presented here can be used for the design and optimization of biodiesel production using reactive distillation.

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Greek Letters

αi,k kinetic order of component i in reaction k νi,k stoichiometric coefficient of component i in reaction k ξk,j extent of reaction k at stage j Subscripts/Superscripts

II I sat i k F j ref

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AUTHOR INFORMATION

Corresponding Author

liquid vapor saturation component reaction feed stage/component reference

REFERENCES

(1) Yusuf, A.; Hanna, M. A. Alternative diesel fuels from vegetable oils. Bioresour. Technol. 1994, 50, 153−163. (2) Kim, S.; Dale, B. E. Life cycle assessment of various cropping systems utilized for producing biofuels: Bioethanol and biodiesel. Biomass Bioenergy 2005, 29, 426−439. (3) Atabani, A. E.; Silitonga, A. S.; Badruddin, I. A.; Mahlia, T. M. I.; Masjuki, H. H.; Mekhilef, S. A comprehensive review on biodiesel as an alternative energy resource and its characteristics. Renewable Sustainable Energy Rev. 2012, 16, 2070−2093. (4) Salvi, B. L.; Panwar, N. L. Source: Biodiesel resources and production technologiesA review. Renewable Sustainable Energy Rev. 2012, 16, 3680−3689. (5) Hajek, M.; Skopal, F.; Cernoch, M. Effect of phase separation temperature on ester yields from ethanolysis of rapeseed oil in the presence of NaOH and KOH as catalysts. Bioresour. Technol. 2012, 110, 288−291. (6) Atadashi, I. M.; Aroua, M. K.; Abdul Aziz, A. R.; Sulaiman, N. M. N. Production of biodiesel using high free fatty acid feedstocks. Renewable Sustainable Energy Rev. 2012, 3275−3285.

*Telephone: +55-4432614749. Fax: +55-4432614774. E-mail: [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS This work was supported by the following Brazilian Funding Agencies: FAPERJ (Grant No. E-26/100.182/2012), Fundaçaõ Araucária, CNPq, and CAPES.

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NOMENCLATURE A, B, C = Antoine constants Ci = molar concentration of component i Ej = relation between the vapor and liquid streams in stage j Fi,j = molar flow rate of the feed stream of component i to stage j H



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(7) Aranda, D. A. G.; Santos, R. T. P.; Tapanes, N. C. O.; Ramos, A. L. D.; Antunes, O. A. C. Acid-Catalyzed Homogeneous Esterification Reaction for Biodiesel Production from Palm Fatty Acids. Catal. Lett. 2008, 122, 20−25. (8) Alves, M. B.; Medeiros, F. C. M.; Suarez, P. A. Z. Cadmium Compounds as Catalysts for Biodiesel Production. Ind. Eng. Chem. Res. 2010, 49, 7176−7182. (9) Halim, R.; Danquah, M. K.; Webley, P. A. Extraction of oil from microalgae for biodiesel production: A review. Biotechnol. Adv. 2012, 30, 709−732. (10) Angel, A. A.; Cruz, Y. R.; Aranda, D. A. G. Hydroesterification of Nannochloropsis oculata microalga s biomass to biodiesel on Al2O3 supported Nb2O5 catalyst. Nat. Sci. 2012, 4, 204−210. (11) Pöpken, T.; Steinigeweg, S.; Gmehling, J. Reactive Distillation for the Synthesis and Hydrolysis of Methyl Acetate using Structured Catalytic Packings: Experiments and Simulation. Ind. Eng. Chem. Res. 2001, 40, 1566. (12) Steinigeweg, S.; Gmehling, J. n-Butyl Acetate Synthesis via Reactive Distillation: Thermodynamic Aspects, Reaction Kinetics, Pilot-Plant Experiments and Simulation Studies. Ind. Eng. Chem. Res. 2002, 41, 5483. (13) Machado, G. D.; Aranda, D. A. G.; Castier, M.; Cabral, V. F.; Cardozo-Filho, L. Computer simulation of fatty acid esterification in reactive distillation columns. Ind. Eng. Chem. Res. 2011, 50, 10176− 10184. (14) Silva, N. L.; Santander, C. M. G.; Batistella, C. B.; Maciel Filho, R.; Maciel, M. R. W. Biodiesel Production from Integration Between Reaction and Separation System: Reactive Distillation Process. Appl. Biochem. Biotechnol. 2010, 161, 245. (15) He, B. B.; Singh, A. P.; Thompson, J. C. A Novel ContinuousFlow Reactor Using Reactive Distillation for Biodiesel Production. Trans. ASAE 2006, 49, 107. (16) Kiss, A. A.; Omota, F.; Dimian, A. C.; Rothenberg, G. C. Solid acid catalysts for biodiesel productionTowards sustainable energy. Adv. Synth. Catal. 2006, 348, 75. (17) Kiss, A. A.; Omota, F.; Dimian, A. C.; Rothenberg, G. The Heterogeneous Advantage: Biodiesel by Catalytic Reactive Distillation. Top. Catal. 2006, 40, 141. (18) Dimian, A. C.; Bildea, C. S.; Omota, F.; Kiss, A. Innovative Process for Fatty Acid Esters by Dual Reactive Distillation. Comput. Chem. Eng. 2009, 33, 743. (19) Kiss, A. Novel Process for Biodiesel by Reactive Absorption. Sep. Purif. Technol. 2009, 69, 280. (20) Kiss, A. A.; Dimian, A. C.; Rothenberg, G. Biodiesel by Catalytic Reactive Distillation Powered by Metal Oxides. Energy Fuels 2008, 22, 598. (21) Kiss, A. A. Separative Reactors for Integrated Production of Ethanol and Biodiesel. Comput. Chem. Eng. 2010, 34, 812. (22) Kiss, A. A. Heat-Integrated Process for Biodiesel by Reactive Absorption. Adv. Synth. Catal. 2010, 348, 75. (23) Chen, F.; Huss, R. S.; Malone, M. F.; Doherty, M. F. Multiple Steady States in Reactive Distillation: Kinetic Effects. Comput. Chem. Eng. 2000, 24, 2457. (24) Alfradique, M. F.; Castier, M. Automatic Generation of Procedures for the Simulation of Reactive Distillation Using Computer Algebra. Comput. Chem. Eng. 2005, 29, 1875. (25) Castier, M. Automatic implementation of thermodynamic models using computer algebra. Comput. Chem. Eng. 1999, 23, 1229. (26) Gmehling, J.; Li, J.; Schiller, M. A Modified UNIFAC Model. 2. Present Parameter Matrix and Results for Different Thermodynamic Properties. Ind. Eng. Chem. Res. 1993, 32, 178−193. (27) Hala, E.; Boulblik, T.; Fried, V. Vapour Pressure of Pure Substances; Elsevier: Amsterdam, 1984; Vol. 17, p 972. (28) NIST Chemistry WebBook. NIST Standard Reference Database Number 69; 2008. (29) DIPPR. Information and Data Evaluation Manager. Public Version 1.2.0; 2000. (30) Rocha, L. L. L.; Ramos, A. L. D.; Antoniosi Filho, N. R.; Furtado, N. C.; Taft, C. A.; Aranda, D. A. G. Production of Biodiesel

by a Two-Step Niobium Oxide Catalyzed Hydrolysis and Esterification. Lett. Org. Chem. 2010, 7, 571. (31) Chenard, G.; Reyes, Y.; Leite, G. G.; Aranda, D. A. G.; Arceo, A. A.; Furtado, N.; Taft, C. A. Hidroesterification of Nannochloropsis oculata microalgàs Biomass to biodiesel using the catalyst pure Niobium Oxide (Nb2O5), Niobium Oxide supported on alumina (Nb2O5/Al2O3) and Niobium Oxide impregnated with Phosphoric Acid (Nb2O5H3PO4). Curr. Org. Chem. 2013, 17, 1. (32) Gonçalves, A. J.; Ramos, A. L. D.; Rocha, L. L. L.; Domingos, A. K.; Monteiro, S. R; Peres, J. S.; Furtado, N. C.; Taft, C. A.; Aranda, D. A. G. Niobium Oxide Solid Catalyst: Esterification of Fatty Acids and Modeling for Biodiesel Production. J. Phys. Org. Chem. 2011, 24, 54. (33) Aranda, D. A. G.; Gonçalves, J. A.; Peres, J. S.; Ramos, A. L. D.; Melo, C. A. R.; Antunes, O. A. C.; Furtado, N. C.; Taft, C. A. The Use of Acids, Niobium Oxide, and Zeolite Catalysts for Esterification Reactions. J. Phys. Org. Chem. 2009, 22, 709.

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