Simulator for Coal Liquefaction Based on the NEDOL Process

Mitsui Engineering and Shipbuilding Company Ltd., ST Nishikasai Bld., 8-4-6, ... Chiba 293-8511, Japan, Mitsui SRC Development Company Ltd., Nibiki Bl...
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Ind. Eng. Chem. Res. 2001, 40, 210-217

PROCESS DESIGN AND CONTROL Simulator for Coal Liquefaction Based on the NEDOL Process Hidenobu Itoh,† Masatake Hiraide,† Akira Kidoguchi,† Masaki Onozaki,‡ Hirohito Ishibashi,‡ Yasuki Namiki,‡ Koichi Ikeda,§ Kenji Inokuchi,| and Shigeharu Morooka*,⊥ Mitsui Engineering and Shipbuilding Company Ltd., ST Nishikasai Bld., 8-4-6, Nishikasai, Edogawa-ku, Tokyo 134-0088, Japan, Nippon Coal Oil Company Ltd., KS Bld., Sanban-cho, Chiyoda-ku, Tokyo 102-0075, Japan, Nippon Steel Corporation, 20-1, Shintomi, Futtsu, Chiba 293-8511, Japan, Mitsui SRC Development Company Ltd., Nibiki Bld., 1-15, Awajicho, Chiyoda-ku, Tokyo 101-0063, Japan, and Department of Applied Chemistry, Kyushu University, Fukuoka 812-8581, Japan

Direct coal liquefaction test plants have been constructed and successfully operated under subsidies from the New Energy and Industrial Technology Development Organization. To analyze the data, which had been collected on these plants, a simulator was constructed for the material and enthalpy balances in the preheaters and reactors. This simulator was evaluated in a process supporting unit (PSU) with a capacity of 1 ton of dry coal/day and a pilot plant (PP) with a capacity of 150 tons of dry coal/day, from the standpoint of vapor-liquid equilibria, reaction rate constants, and hydrodynamics in the reactors. Reaction rate coefficients were determined in the PSU by iteration until the calculated yields at the outlet of the third reactor coincided with the observed data. The kinetic data were then applied to the PP, which was operated under conditions similar to those of the PSU, and the liquefaction product yields at the outlet of the third reactor were calculated. The calculated values were in general agreement with the observed data. 1. Introduction The production of liquid fuels via the direct liquefaction of coal has represented an important goal in the field of energy research since the beginning of the 20th century.1 In Japan, several direct coal liquefaction processes were integrated into the NEDOL process in 1984 with support from the New Energy and Industrial Technology Development Organization (NEDO). In this project, a process supporting unit (PSU) with a capacity of 1 ton/day and a pilot plant (PP) with a capacity of 150 tons/day were successfully constructed and operated using a variety of coals. A huge amount of data, including product yields, hydrogen consumption, distillation curves of produced oil, physical properties of narrow-cut oil fractions, temperature and pressure distributions, andflow properties in the reactors, has been collected and is currently in a database. It was found that the hydrodynamic behaviors in the liquefaction reactors were considerably different from those of bubble columns, which were operated using air-water systems.2,3 To design large-scale reactors for future use, the causes and effects among these data need to be systematically investigated. In this study, reaction schemes, vapor-liquid equilibria, and flow properties * To whom correspondence should be sent. Fax: +81-92651-5606. E-mail: [email protected]. † Mitsui Engineering and Shipbuilding Co., Ltd. ‡ Nippon Coal Oil Co., Ltd. § Nippon Steel Corp. | Mitsui SRC Development Co., Ltd. ⊥ Kyushu University.

in the PSU reactors were investigated using Tanitoharum coal. The reaction rate constants were then determined on the basis of the product distributions at the outlet of the reactors. Using the data evaluated in the case of the PSU, it was possible to simulate the liquefaction performance of the PP, and the calculated and experimental values were found to be in agreement. 2. The NEDOL Coal Liquefaction Process Figure 1 shows an outline of the NEDOL process, which consists of (i) a coal-slurry preparation stage, (ii) a preheater, (iii) reactors, (iv) a vapor-liquid separator, (v) atmospheric and vacuum distillation towers, and (vi) a solvent hydrogenation stage which uses a Ni/Mo catalyst. Table 1 shows the dimensions of the preheaters and reactors used in the PSU and PP. The PSU preheater consisted of the heat-transfer lines, which were heated by electromagnetic induction. The preheater for the PP was a combination of a heat exchanger and heat-transfer lines and was capable of heating the slurry to 663-702 K. At the coal-slurry preparation stage, coal was powdered and mixed with hydrogen-rich gas (recycle gas), a hydrogenated solvent (recycle oil), and a disposable catalyst. The catalyst was finely ground pyrite (average diameter ) 0.7 µm). The mixture was introduced into the preheating section and then into the coal liquefaction reactors, which consisted of three bubble columns connected in series for both the PSU and PP. A part of the recycle gas was independently heated and introduced into the first reactor. The temperature rise due

10.1021/ie0004085 CCC: $20.00 © 2001 American Chemical Society Published on Web 11/21/2000

Ind. Eng. Chem. Res., Vol. 40, No. 1, 2001 211

Figure 2. Scheme for the initial reactions in the preheater.

Figure 1. Outline of the NEDOL coal liquefaction process. Table 1. Dimensions of the Preheaters and Reactors dimension

PSU

PP

i.d., m length,b m heating method

Preheater 0.0143 94 electromagnetic induction

0.0142/0.072a 560 heat exchanger and fired heater

i.d., m height,c m

0.175 1.75

Reactor 1.0 11

a The PP preheater consisted of a slurry phase exchanger (multiple-path, tube i.d. ) 0.0142 m) and a furnace heater (singlepath, tube i.d. ) 0.072 m) and was capable of heating the slurry to 702 K at maximum. b Includes the connecting piping, which was heated. c Does not include the upper and lower conical sections.

Table 2. Properties of Tanitoharum Coal Proximate Analysis moisture, wt % ash, wt % dry coal volatile matter, wt % dry coal fixed carbon, wt % dry coal

14.5 5-7 46-47 49

Ultimate Analysis carbon, wt % daf coal hydrogen, wt % daf coal nitrogen, wt % daf coal oxygen (diff.), wt % daf coal sulfur, wt % daf coal H/C atomic ratio density, Mg/m3

76.6 5.6 1.4 16.2 0.2 0.87 1.4

to the liquefaction was controlled by introducing the recycle gas at a temperature of 323 K into the reactors. After the liquefaction, the slurry was separated by distillation, and a portion was treated in the hydrogenation stage, while the remainder was returned to the coalslurry preparation stage as the solvent. A variety of coals were tested in the development of the NEDOL process. In the present study, however, only the results for Tanitoharum coal are considered. Table 2 shows the properties of Tanitoharum coal used in the PSU and PP. 3. Outline of the Liquefaction Simulator The simulation was performed using a multipurpose process simulator (Aspen Plus). The factors taken into consideration were reaction rates, physical properties of components, vapor-liquid equilibria, hydrodynamics in the reactors (gas holdup, axial dispersion of slurry, etc.), accumulations of fine solid particles in the reactors, enthalpy balances in consideration of heats of reactions, and axial temperature distributions in the reactors.

These data are reported elsewhere.4-7 In the present paper, only material balances based on reaction rates and vapor-liquid equilibria in the reactors are simulated using the following assumptions: (i) Each reactor is well mixed, both horizontally and vertically. Thus, no axial distributions with respect to concentrations, gas holdup, and temperature exist in the reactor. (ii) The coal consists of three components: CA, which is rapidly liquefied to preasphaltene, CB, which is liquefied at a finite rate, and CI, which is inert. The lumping of coal is based on liquefaction reactivity and is different from the original concept of Nagaishi et al.,8 who evaluated the components intrinsic to each coal. The coal components, which are determined in the present study, should be used under specified conditions. (iii) Catalyst particles are suspended in the liquid phase homogeneously with no slip velocity. All reactions except for the water gas shift reaction are governed by irreversible first-order kinetics in the slurry phase. The water gas shift reaction is assumed to be equilibrated among H2O, CO, and CO2 in the vapor phase. The other reactants are not reactive in the vapor phase, which contains no catalyst. (iv) Vapor-liquid equilibria are always established. (v) The liquid products are divided into O1, O2, O3, and PAAO on the basis of boiling point (O1 ) C4 gases and oil with the boiling point below 493 K; O2 ) oil with the boiling point of 493-623 K; O3 ) oil with the boiling point of 623-811 K; and PAAO ) a mixture of preasphaltene, asphaltene, and oil with the boiling point above 811 K). 4. Simulator for the preheater 4.1. Reaction Model for the Preheater. Figure 2 shows the scheme for initial reactions, which can occur in the preheater. Coal component CA is decomposed to PAAO, oil, and gas via parallel reaction paths. Component CB is also converted into PAAO. Consecutive reactions, PAAO to oil and to gas, are neglected in the initial reaction stage. The kinetic constants for these reactions were estimated by autoclave reaction tests. Because details have been reported elsewhere,9 only the main points are addressed in this paper. The coal was pulverized into particles of less than 100 mm in size, and a 3 g sample was suspended in 7 g of the equilibrated recycle solvent, which was produced in the PSU. Figure 3 shows the distillation curve of a typical recycle solvent. A powdered iron sulfide catalyst was mixed at a concentration of 3 wt % based on the mass of dry ash free (daf) coal. The mixture was then placed in an autoclave with an internal capacity of 55 cm3 and was pressurized with hydrogen to 10.1 MPa. The autoclave was then heated at a rate of 50 K/min to a prescribed temperature. After the temperature was maintained for the prescribed period, the reaction was quenched by

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Ind. Eng. Chem. Res., Vol. 40, No. 1, 2001 Table 3. Kinetic Constants of the Initial Stage Reactions and Components in Tanitoharum Coal (Temperature ) 723 K, Total Pressure ) 17 MPa) Kinetic Constant, min-1 kCA ) 0.29 kCAO ) 1.5 kCIG ) 0.08 kCB ) 0.11 activation energy ) 95.5 kJ/mol Coal Component, wt % daf Coal Basis CA ) 60.1 CB ) 30.0 CI ) 9.9 Table 4. Operating Conditions for the PSU

Figure 3. Distillation curve of a typical recycle oil for Tanitoharum coal.

placing the autoclave into water at room temperature. The product was sampled after cooling. Within 5 min after the reaction started, 50-60 wt % of the Tanitoharum coal was converted to PAAO, oil, and noncondensable gases. The amount of unreacted coal decreased gradually, and more than 90 wt % of each coal was reacted within 60 min at 723 K. The amount of PAAO, on the other hand, increased rapidly, and PAAO and oil fractions increased in parallel. Carbon dioxide was the major component of the noncondensable gas. The mass of the total unreacted coal at time t, WC(t), is expressed by

WC(t) ) WCA(t) + WCB(t) + WCI

(1)

where WCA(t), WCB(t), and WCI are the masses of the CA, CB, and CI components in the reactor at time t, respectively. The reaction rates of CA and CB are expressed by

dWCAt(t)/dt ) -kCAtWCA(t)

(2)

dWCB(t)/dt ) -kCBWCB(t)

(3)

kCAt ) k0CAt exp(-ECAt/RT)

(4)

kCB ) k0CB exp(-ECB/RT)

(5)

where kCAt is the sum of kCA, kCAO, and kCIG, which are the reaction rate coefficients of the paths to PAAO, oil, and gas, respectively. k0CAt and k0CB are the frequency coefficients for reactions 4 and 5, respectively, and ECAt and ECB represent their activation energies. Table 3 shows the kinetic constants and coal components, which were obtained for Tanitoharum coal in the autoclave reactor at a temperature of 723 K and a pressure of approximately 17 MPa. All of the activation energies were assumed to be the same as the value obtained for Wandoan coal (95.5 kJ/mol), the reactivity of which was examined at temperatures of 693, 713, and 723 K. 4.2. Verification of the Reaction Model for the Preheater. The validity of the initial reactions was examined in the preheater of the PSU, using the kinetic constants shown in Table 3. As listed in Table 1, the preheater consisted of a coil of 0.0143 m i.d., the total

recycle gas fraction of hydrogen, vol % makeup slurry slurry feed rate, kg/h coal concentration in slurry, wt % dry coal basis catalyst (pyrite powder) in slurry, wt % dry coal basis pressure in reactors, MPa Gr/Lf,a m3(STP)/kg of slurry Gr/Lf,b m3(STP)/kg of slurry temperature bottom of the first reactor, K top of the first reactor, K bottom of the second reactor, K top of the second reactor, K bottom of the third reactor, K top of the third reactor, K mean residence time of slurry,c min average superficial gas velocity, cm/s average gas holdup

87 106 39 3.0 16.8 0.7 0.6 734 740 732 738 735 738 129 1.4-1.5 0.16-0.17

a (Volumetric flow rate of recycle gas totally fed to the reactors)/ (mass flow rate of makeup coal slurry). b (Volumetric flow rate of recycle gas fed to the preheater inlet)/(mass flow rate of makeup coal slurry). c Determined by the NAT method.

length of which was 94 m. The recycle gas was fed at a flow rate of 63.6 m3/h at standard temperature and pressure (STP), and the slurry was fed at a flow rate of 106 kg/h. The mean residence time of the slurry in the preheater was 52 s. This value was much shorter than the mean residence time of the slurry in the whole PSU reactor, 130 min, as shown in Table 4, which is described in section 5.4. Thus, the simulation for the preheater can be considerably simplified, although the procedure is fundamentally the same as that of the reactors described in section 5. The simulation of the reactions in the PSU preheater was performed under the following conditions: (i) The total length of the preheater was divided into 10 sections, which were connected in series. For each section, temperature, partial pressures, and concentrations were assumed to be constant. (ii) The vapor-liquid equilibria in the preheater were estimated from the equations described in section 5.2. (iii) The slip velocity between the liquid and solid particles in the slurry phase was neglected. The flow rates of the gas and the liquid at each local position were estimated from the vapor-liquid equilibria and the reaction rates. The gas holdup in the horizontal preheater was estimated from the equation derived by Lockhart and Martinelli.10 (iv) The feed was prepared using powdered Tanitoharum coal, which was mixed with the recycle oil at a concentration of 39 wt %. The mixture also contained a finely ground pyrite catalyst at a concentration of 3 wt

Ind. Eng. Chem. Res., Vol. 40, No. 1, 2001 213

Figure 4. Yields at the outlet of the PSU preheater (from Kidoguchi et al.9).

C1, C2, and C3 gases. Coal is decomposed to fraction PAAO in the reactors, and fraction PAAO is then decomposed to fractions O1, O3, OG, IOG1, IOG2, and IOG3. Most of fraction O2 is vaporized immediately after it is produced. Thus, the reaction of fraction O2 cannot proceed on the surface of the catalyst particles in the slurry phase. The hydrogenation path from fraction O2 to fraction O1 is then neglected. Two paths to fraction O2 are possible, from fractions PAAO and O3. However, these paths cannot be separated using the final products. The path from O3 to O2 is adopted in the present simulation because of a better agreement with the data. Hydrogen is largely consumed in the hydrogenation of fraction PAAO. All of the above reactions occur in the slurry phase except for the water gas shift reaction in the gas phase and are assumed to be expressed by firstorder reaction kinetics. Thus, the material balances in each reactor can be described as follows:

-(kCA + kCA0 + kCIG)mCAVslsl ) QimCAi - QomCA -(k1 + k8)mCBVslsl ) QimCBi - QomCB

(6) (7)

[kCAmCA + k1mCB - (k2 + k4 + k5 + k6 + k7)mPAAO]Vslsl ) QimPAAOi - QomPAAO (8) (kCA0mPAAO + k2mPAAO - k3mO3)Vslsl ) QimO3i - QomO3 - FO3 (9) Figure 5. Scheme for reactions in reactors.

% (daf coal basis). The recycle oil, with the distillation curve shown in Figure 3, was also used in this test. (v) The preheater was heated from 353 to 643 K (case 1), 683 K (case 2), and 723 K (case 3). The temperature profile along the preheating tube was determined from the experimental data of the skin temperature. The total pressure was 17.3 MPa throughout the preheater. (vi) The outlet flow of the preheater was introduced directly into the vapor-liquid separator, and the composition of the mixture at the preheater outlet was determined. Figure 4 shows the compositions which were determined for the mixture, as well as those estimated by the simulator. The yields of the components were calculated based on the mass of the daf coal, which was mixed with the recycle oil. In cases 1 and 2, the calculated yields of the unreacted coal were found to coincide with those observed. The agreement between the observed and the calculated yields was also confirmed for the PAAO yield. In case 3, the observed yield of oil was higher, whereas the PAAO yield was lower, than the calculated yields. This suggests that the reaction of PAAO proceeded at a faster rate than that of the estimated value. The conversion of PAAO to oil was neglected in the simulator for the initial reactions, but a slight conversion might occur. 5. Simulator for Reactors 5.1. Reaction Model in Reactors. Figure 5 shows the liquefaction scheme in the reactors, assuming three components of coal (CA, CB, and CI), one intermediate fraction (PAAO), three fractions of oil (O1, O2, and O3), and five gas fractions (IOG1, IOG2, IOG3, OG, and H2). Fraction IOG1 consists of CO and CO2, IOG2 consists of H2O, IOG3 consists of H2S and NH3, and OG consists of

k3mO2Vslsl ) QimO2i - QomO2 - FO2

(10)

k4mPAAOVslsl ) QimO1i - QomO1 - FO1

(11)

k5mPAAOVslsl ) QimOGi - QomOG - FOG

(12)

(kCIGmCA + k8mCB)Vslsl ) QimIOG1i - QomIOG1 - FIOG1 (13) k6mPAAOVslsl ) QimIOG2i - QomIOG2 - FIOG2

(14)

k7mPAAOVslsl ) QimIOG3i - QomIOG3 - FIOG3

(15)

The reaction rate constant for hydrogen is calculated from the total balance of hydrogen in the reactor.

k9mPAAOVslsl ) QimH2i - QomH2 - FH2

(16)

In the above equations, Vsl is the volume of the slurry phase in the reactor, Fsl the density of the slurry phase, and mj the mass fraction of component j in the slurry phase. Qi and Qo are the total mass flow rates of the slurry phase at the inlet and outlet of each reactor, respectively. Fj is the mass of component j, which is transferred to the gas phase in the reactor per unit time. Because hydrogen is transferred from the gas phase to the slurry phase, FH2 possesses a negative value. As discussed in section 4, coal fraction CA is assumed to be completely liquefied to fractions PAAO, O3, and IOG3 in the preheater. Thus the concentration of fraction CA in each reactor is negligible.

mCA ) 0

(17)

The total material balance in the reactor is

Qi - Qo )

∑j Fj

(18)

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The mean residence time of the slurry phase in each reactor, τsl, is given as

τsl ) VslFsl/Qo

(19)

The total mean residence time of the slurry is calculated as the sum of τsl in the first, second, and third reactors. 5.2. Vapor-Liquid Equilibria. Hydrogen in the recycle gas is dissolved in the slurry phase, whereas reaction products are vaporized during the course of the liquefaction. The vapor-liquid equilibria can be expressed by the following equations.

Soave-Redlich-Kwong (SRK) equation11 p)

aR RT Vm - b Vm(Vm + b)

(20)

a ) 0.42747R2Tc2/pc

(21)

b ) 0.08664RTc/pc

(22)

R ) {1 + m[1 - (T/Tc)0.5]}2

(23)

m ) 0.480 + 1.574ω - 0.176ω2

(24)

Peng-Robinson (PR) equation12 p)

RT aR Vm - b Vm(Vm + b) + b(Vm - b)

(25)

a ) 0.45724R2Tc2/pc

(26)

b ) 0.0778RTc/pc

(27)

R ) {1 + m[1 - (T/Tc)0.5]}2

(28)

m ) 0.37464 + 1.54226ω - 0.26992ω2

(29)

Benedict-Webb-Rubin (BWR) equation13 p ) RTF+(B0RT - A0 - C0/T2)F2 + (bRT - a)F3 + aRF6 + (cF3/T2)(1 + γF2) exp(-γF2) (30) In eq 30, F ) 1/Vm, and A0, B0, C0, a, b, c, R, and γ are adjustable parameters. Commonly in eqs 20-30, ω is the acentric factor, R the gas constant, T the temperature, p the pressure, Tc the critical temperature, pc the critical pressure, and Vm the molar volume of the gas. Values of ω, Tc, and pc are calculated from the correlations proposed by Watanasiri et al.,14 on the basis of boiling points, specific densities, and molecular weights for the narrow-cut oil fractions. Any parameter for multicomponent systems, Y, can be calculated from

Y)

∑i ∑j xixjYij

(31)

where xi represents the mole fraction of component i and Yij the cross contribution between components i and j. To check the error in the calculation of the vaporliquid equilibria, the product oil which was recovered at the outlet from the first reactor was divided into 3, 4, 6, 8, 12, and 45 fractions on the basis of boiling point. The physical properties of each fraction were then determined. The vapor-liquid equilibria for these fractions, as well as the light components such as O1 and

Figure 6. Gas holdup in reactors: (b) in PP reactors; (O) in PSU reactors. The superficial gas velocity should be calculated under liquefaction conditions.

gases, were calculated from eqs 20-31 under liquefaction conditions. No large differences were found for the division numbers. Thus, the oil was subsequently divided into three fractionssO1, O2, and O3sas shown in Figure 5. The validity of the SRK, PR, and BWR equations was evaluated by comparing the calculated mean residence times with the experimental values, which were determined by the neutron absorption tracer (NAT) method. Because details have been published elsewhere,5 the experimental procedure is only summarized here. Finely divided gadolinium oxide (Gd2O3; average diameter ) 2 µm) was used as the tracer for neutron absorption. A tracer slurry, prepared by suspending the gadolinium powder in the coal oil at a concentration of 50 wt %, was rapidly injected into the feed line to the reactor. At the outlet of the reactor, low-energy neutrons were irradiated horizontally using californium-252 as the neutron source. Neutrons were counted using a 3He counter, which was installed at the opposite side of the pipe from the neutron source, and the neutron intensity was converted to the concentration of the tracer, on the basis of calibration curves, which were prepared prior to the measurement. The mean residence time of the slurry phase in the third PSU reactor was then calculated from the residence time distribution curve and was found to be 2600 s. The mean residence times which were calculated based on the SRK, PR, and BWR equations for this reactor were 2500, 3200, and 1600 s, respectively. Thus, the SRK equation was, hereafter, used for the calculation of the vapor-liquid equilibria. 5.3. Gas Holdup in Reactors. The mean gas holdup in the reactor was determined using three methods:4,5 (i) the differential pressure method, (ii) the gas shutdown method, and (iii) the tracer method. In method i, the gas holdup was calculated from the pressure profile, which was vertically produced in the reactors. In method ii, the gas flow was instantly interrupted, resulting in the formation of a space at the top of the reactor after all of the gas bubbles had been settled. Because the liquid was continuously fed to the reactor, the space was then filled by the liquid. The volume of gas bubbles was calculated from that of the liquid which was fed during this period. Method iii is described in section 5.2. Gas holdup data for the PSU and PP reactors, as determined by the different methods, are shown in Figure 6. The superficial gas velocity on the abscissa is calculated from vapor-liquid equilibria, which are described in section 5.2. Thus, the gas holdup data in the PSU and PP reactors can be typically correlated, based on the superficial gas velocity. The slurry volume in the reactor is then given as

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Vsl ) VR(1 - )

(32)

where VR is the total reactor volume, and  is the gas holdup. 5.4. Verification of the Reaction Model. Table 4 shows typical operating conditions of the PSU reactors. Because the temperature was controlled at the top of each reactor, temperature distributions were observed to some extent along the axial positions. The operating conditions of the PSU preheater were similar to those of case 2 in Figure 4. Tables 5 and 6 show the products at the outlet of the third reactor of the PSU. The values of Tc, pc, and ω in Table 6 were calculated from the correlations proposed by Watanasiri et al.14 Hydrogen was incorporated into liquefaction products in the reactors, and as a result, the total yield on the basis of daf coal mass was increased to 105.5 wt % at the outlet of the third reactor, as shown in Table 5. The yield of oil (C4 gases to the bp 811 K fraction) was 55.7 wt %, and the residue (PAAO + CI) was 14.4 wt %. The simulation for the PSU reactors was then performed based on the reaction scheme shown in Figure 5, using the results of the preheater, shown in Figure 2 and Table 3. The vapor-liquid equilibria described in section 5.2, as well as the gas holdup data described in section 5.3, were also used in the simulation. The iteration was repeated by the following procedure: (i) Calculate the product distribution at the outlet of the preheater, (ii) assume the physical properties of the oil fractions in the reactors from the given product distribution at the outlet of the third reactor, (iii) assume reaction rate constants, (iv) assume the superficial gas velocity and calculate the gas holdup and the mean residence time in the reactors, (v) calculate the compositions in the vapor and liquid phases, (vi) compare the calculated product distribution at the outlet of the third reactor with the operating data, and (vii) repeat the procedure by altering the reaction rate constants, until the calculated data coincide with the data shown in Table 5. Figure 7 shows the calculated product distributions in the feed to PSU reactor 1 (i.e., at the outlet of the preheater) and at the outlets of PSU reactors 1-3. The recycle gas and oil are included, but the coal ash and catalyst particles are excluded from the calculation. Because no quench gas is introduced in the PSU reactors, the total mass remains unchanged. The product distributions observed at the outlet of reactor 3 are in good agreement with those calculated by the simulation. Table 7 shows the kinetic constants, which were thus determined. The large value of k1 suggests that PAAO was more reactive than small-molecule components. Figure 8 shows the yields in the PSU reactors. Hydrogen is incorporated into the liquefaction products, and the total yield based on the daf coal mass increases. It can be assumed that the decomposition of Tanitoharum coal in the PSU reactors was fundamentally completed in the preheater and the first reactor. Fraction PAAO was produced mainly in the first reactor and was further decomposed to oil fractions in the second and third reactors. 5.5. Application of the Simulator to PP. Table 8 shows the operating conditions of the PP. The recycle gas, total pressure, and gas/slurry ratio are fundamentally the same as those in the PSU, but the temperature in the PP reactors was 10-12 K lower than that in the PSU reactors. Tables9 and 10 show the products, which were obtained at the outlet of the third PP reactor. The

Table 5. Products in the PSU yields, wt % daf coal basis gas water oil (C4 to the bp 811 K fraction) residue total hydrogen consumption, wt % daf coal basis

24.6 10.8 55.7 14.4 105.5 5.5

Table 6. Oil Fractions Produced in the PSU

O1 O2 O3 a

average bp, K

s.g. 60a

MW, g/mol

Tc, K

P c, MPa

ω

487 559 666

0.9385 0.9752 1.0754

137 174 237

715 796 923

3.40 2.86 2.33

0.380 0.456 0.503

Specific gravity at 60 °F.

Figure 7. Distributions in the PSU reactors.

Figure 8. Yields in the PSU reactors. Table 7. Kinetic Constants in the PSU under the Conditions Shown in Table 4 k1 ) +0.046 min-1 k2 ) +0.0048 k3 ) +0.0034 k4 ) +0.0082 k5 ) +0.0051 k6 ) +0.0033 k7 ) +0.000 35 k8 ) +0.0032 k9 ) -0.0017

values of Tc, pc, and ω in Table 10 were calculated from the correlations proposed by Watanasiri et al.14 Because these physical properties were different from those found in the PSU, the vapor-liquid equilibria in the PP were estimated based on the data shown in Table 10.

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Table 8. Operating Conditions for the PP recycle gas fraction of hydrogen, vol % makeup slurry slurry feed rate, kg/h coal concentration in slurry, wt % dry coal basis catalyst (pyrite powder) in slurry, wt % dry coal basis pressure in reactors, MPa Gr/Lf,a m3(STP)/kg of slurry Gr/Lf,b m3(STP)/kg of slurry temperature bottom of the first reactor, K top of the first reactor, K bottom of the second reactor, K top of the second reactor, K bottom of the third reactor, K top of the third reactor, K mean residence time of slurry,c min average superficial gas velocity, cm/s average gas holdup

86 16 000 40 3.0 16.9 0.7 0.3 699 728 727 729 726 728 86 6.4 0.51

a (Volumetric flow rate of recycle gas totally fed to the reactors)/ (mass flow rate of makeup coal slurry). b (Volumetric flow rate of recycle gas fed to the preheater inlet)/(mass flow rate of makeup coal slurry). c Determined by the NAT method.

Table 9. Products in the PP yields, wt % daf basis gas water oil (C4 to the bp 811 K fraction) residue total hydrogen consumption, wt % daf coal basis

17.2 10.2 51.0 26.1 104.5 4.5

Table 10. Oil Fractions Produced in the PP

O1 O2 O3 a

average bp , K

s.g. 60a

MW , g/mol

Tc, K

P c, MPa

ω, g/mol

487 559 666

0.9400 0.9688 l.0663

140 194 258

716 798 928

3.30 2.45 2.05

0.384 0.459 0.692

Specific gravity at 60 °F.

Figure 9. Distributions in the PP reactors.

The kinetic constants obtained for the PSU reactors were applied to the PP reactors without modification. Figure 9 shows the calculated product distributions at the inlet of the first PP reactor and the outlets of the first, second, and third PP reactors. The PP preheater was comprised of a slurry heat exchanger and a furnace heater, and the estimated residence time of the slurry phase was 111 s, which was longer than that in the PSU preheater, 52 s. The total mass percentage at the outlet of each PP reactor increases beyond 100, because

Figure 10. Yields in the PP reactors.

hydrogen is introduced into the reactor as the quench gas. At the outlet of the third reactor, the calculated product distribution is generally in agreement with that of the experimental results shown in Table 9. However, the calculated value for the residue (unreacted coal and PAAO) at the outlet (30.6 wt %) in the PP is somewhat higher than the observed value (26.1 wt %). Figure 10 shows the yields based on the daf coal mass at the inlet of the first PP reactor and the outlets of the first, second, and third reactors. The conversion of coal to PAAO and that of PAAO to oil proceeded in all of the PP reactors. The calculated values were in good agreement with the observed data. Based on this result, the simulator was applied to design a demonstration plant with a capacity of 2500 tons/day (dry coal basis), as reported by Onozaki et al.7 6. Conclusions A simulator was constructed for direct coal liquefaction plants on the basis of the concept of the NEDOL process. The behavior of the PSU reactors, with a capacity of 1 ton of dry coal/day, was simulated using data, such as vapor-liquid equilibria, hydrodynamic properties in the reactors, and reaction rate constants. The vapor-liquid equilibria were calculated from the SKR equation. The applicability of this equation was confirmed by comparing the calculated mean residence time with the experimental one, determined by the NAT method under the liquefaction conditions. The reaction rate constants in the PSU reactors were calculated based on the product distributions at the outlet of the third PSU reactor. The simulator was then applied to the PP, which had a capacity of 150 tons of dry coal/ day and was operated under conditions similar to those of the PSU. The product distributions, which were estimated by the simulator at the outlet of the third PP reactor, were in agreement with those found from the actual operating conditions. Acknowledgment This work is supported by the New Energy and Industrial Technology Development Organization (NEDO). The authors express their gratitude to NEDO for permission to publish this paper. Literature Cited (1) Shah, Y. T. Reaction Engineering in Direct Coal Liquefaction; Addison-Wesley Pub.: Reading, MA, 1981.

Ind. Eng. Chem. Res., Vol. 40, No. 1, 2001 217 (2) Fan, L.-S. Gas-Liquid-Solid Fluidization Engineering; Butterworth Pub.: Stoneham, MA, 1989. (3) Deckwer, W.-D. Bubble Column Reactors; John Wiley & Sons: Chichester, U.K., 1992. (4) Ishibashi, H.; Onozaki, M.; Kobayashi, M.; Hayashi, J.-i.; Itoh, H.; Chiba, T. Gas Holdup in Bubble Column Reactors of 150 t/d Coal Liquefaction Pilot Plant Process. Fuel 2001, in press. (5) Sakai, N.; Onozaki, M.; Saegusa, H.; Ishibashi, H.; Hayashi, T.; Kobayashi, M.; Tachikawa, N.; Ishikawa, I.; Morooka, S. Fluid Dynamics in Coal Liquefaction Reactors Using Neutron Absorption Tracer Technique. AIChE J. 2000, 46, 1688. (6) Onozaki, M.; Namiki, Y.; Ishibashi, H.; Takagi, T.; Kobayashi, M.; Morooka, S. Steady-State Thermal Behavior of CoalLiquefaction Reactors of NEDOL Process. Energy Fuels 2000, 14, 355. (7) Onozaki, M.; Namiki, Y.; Ishibashi, H.; Kobayashi, M.; Itoh, H.; Hiraide, M.; Morooka, S. A process Simulation of the NEDOL Coal Liquefaction Process. Fuel Process. Technol. 2000, 64, 253. (8) Nagaishi, H.; Moritomi, H.; Sanada, Y.; Chiba, T. Evaluation of Coal Reactivity for Liquefaction Based on Kinetic Characteristics. Energy Fuels 1988, 2, 522. (9) Kidoguchi, A.; Itoh, H.; Hiraide, M.; Kaneda, E.; Ishibashi, H.; Kobayashi, M.; Ikeda, K.; Imada, K.; Kidoguchi, K. Simulation

of Initial Stage Reactions in Direct Coal Liquefaction of Subbituminous Coal. Fuel 2001, submitted for publication. (10) Lockhart, R. W.; Martinelli, R. C. Proposed Correlation of Data for Isothermal Two-Phase, Two-Component Flow in Pipes. Chem. Eng. Prog. 1949, 45, 39. (11) Soave, G. Equilibrium Constants from a Modified RedlichKwong Equation of State. Chem. Eng. Sci. 1972, 27, 1197. (12) Peng, D.-Y.; Robinson, D. B. A New Two-Constant Equation of State. Ind. Eng. Chem. Fundam. 1976, 15, 59. (13) Benedict, M.; Webb, G. B.; Rubin, L. C. An Empirical Equation for Thermodynamic Properties of Light Hydrocarbons and Their Mixtures. I. Methane, Ethane, Propane and n-Butane. J. Chem. Phys. 1940, 8, 334. (14) Watanasiri, S.; Owens, V. H.; Starling, K. E. Correlations for Estimating Critical Constants, Acentric Factor, and Dipole Moment for Undefined Coal-Fluid Fractions. Ind. Eng. Chem. Process Des. Dev. 1985, 24, 294.

Received for review April 17, 2000 Accepted September 8, 2000 IE0004085