Ind. Eng. Chem. Res. 1997, 36, 5439-5446
5439
Modeling of Sulfided Zinc Titanate Regeneration in a Fluidized-Bed Reactor. 2. Scale-Up of the Solid Conversion Model J. T. Konttinen,*,† C. A. P. Zevenhoven,‡ and M. M. Hupa‡ Carbona Corporation, P.O. Box 610, FIN-33101 Tampere, Finland, and Department of Chemical Engineering, A° bo Akademi University, Lemminka¨ isenkatu 14-18 B, FIN-20520 Turku, Finland
Hot gas desulfurization with regenerable metal oxide sorbents is an essential part of the simplified integrated gasification combined cycle process which is currently being demonstrated at commercial scale. As part of a sulfur removal process development, Carbona Corporation is developing fluidized-bed reactor models for scale-up. In the first part of this work, the parameters for an uniform conversion model were determined to describe the conversion rate of ZnS regeneration with oxygen. A method using the uniform conversion model for fluidized-bed application is presented in this paper. The apparent activation energy for the global reaction rate constant obtained from pilot-scale fluidized-bed tests is 200-210 kJ/mol. When using these reaction rate constants as inputs of a sensitivity analysis for a large-scale reactor model, in comparison with the rate constants reported earlier, some differences can be observed. The significance of the differences are discussed. It is stated that any zinc sulfate formed in fluidizedbed regeneration at temperatures 500-600 °C will decompose, forming SO2, at the actual regeneration temperatures of the bed (725-800 °C) due to neglible O2 partial pressure. Introduction H2S removal with regenerable sorbents at high temperature and pressure, within the context of a simplified integrated gasification combined cycle (IGCC) process, has been investigated wordwide (Gupta and Gangwal, 1992; Harrison, 1995; Salo et al., 1995). With the use of regenerable sorbents, the amount of solid waste produced is minimized and the fuel-bound sulfur can be recovered as a commercial product, such as elemental sulfur (Portzer et al., 1995). Zinc titanate appears to be the leading sorbent for high-temperature highpressure (HTHP) sulfur removal in fluidized-bed reactors (Harrison, 1995; Salo et al., 1995). Two different zinc titanate sorbents, D (UCI-5) and E (ZT-4-L), were tested in HTHP pilot-scale fluidizedbed sulfur removal reactors (Salo et al., 1995; Konttinen et al., 1996). A schematic diagram of the reactor system is shown in Figure 1. In the beginning of the test runs, a batch of fresh zinc titanate sorbent was fed into the sulfur removal reactor chamber. The H2S-containing coal gasifier gas then fluidized the sorbent bed, causing sulfur removal to take place at 450-650 °C and 10-20 bar according to the following reaction (Lew, 1990): 1
/xZnxTiyO(x+2y)(s) + H2S(g) f ZnS(s) + y/xTiO2(s) + H2O(g) (1)
Both sorbents were in contact with real coal gasifier gas, which contained impurities such as gasification fines, alkali and trace metals, HCl, ammonia, among others. High, up to 99%, sulfur removal efficiency could be achieved with both sorbents. The contact with coal gas for 5-6 days had no detrimental effect on sorbents’ physical or chemical properties. A detailed analysis of the results on experimental testing and modeling of * Author to whom correspondence should be addressed. Telephone: +358-3-358 0314. Fax: +358-3-358 0325. E-mail:
[email protected]. † Carbona Corporation. ‡ A ° bo Akademi University. S0888-5885(97)00036-5 CCC: $14.00
Figure 1. Schematic diagram of the pilot-scale fluidized-bed sulfur removal and sorbent regeneration reactor test rig.
sulfur removal (reaction 1) at different scales are reported earlier (Konttinen et al., 1996, 1997a,b). During continuous operation, the zinc sulfide containing sorbent (as a result of reaction 1) was transported pneumatically from the sulfur removal reactor to the regeneration reactor and regenerated at 600-750 °C and 10-20 bar, thus releasing gaseous SO2 via the following reaction:
ZnS(s) + 1.5 O2(g) f ZnO(s) + SO2(g)
(2)
The regenerated sorbent was then transported back to the sulfur removal reactor. The SO2 produced upon regeneration can be treated further to yield elemental sulfur (Portzer et al., 1995). At regeneration conditions, besides reaction 2, undesired zinc sulfate can form via the following side reaction (Siriwardane and Woodruff, 1995; Konttinen et al., 1997c):
ZnO(s) + SO2(g) + 0.5 O2(g) f ZnSO4(s)
(3)
In the ambient-pressure laboratory-scale tests (Kont© 1997 American Chemical Society
5440 Ind. Eng. Chem. Res., Vol. 36, No. 12, 1997 Table 1. Chemical and Physical Properties of the Sorbents Used in the Pilot-Scale Tests (Salo et al., 1995) property zinc (wt %) bulk density (g/cm3) particle density (g/cm3) mercury pore volume (cm3/g) porosity surface area (m2/g) median pore diameter (Å) average particle size (µm)
sorbent D sorbent E sorbent D sorbent E (fresh) (fresh) (used)a (used)a 46.8 1.47 2.26 0.21 0.475 1.4 7000 309
1.64 2.14 0.22 0.47 3.5 2700 239
41.6 1.51 2.44 0.172 0.42 1.2 7000 285
1.61 2.38 0.17 0.405 5.5 2200 237
a The properties of the fluidized-bed material samples taken close to the time of the regeneration tests presented in Figures 3 and 4.
tinen et al., 1997c), the rate of the undesirable zinc sulfate formation reaction at relevant regeneration conditions was found not to be significant. The pressurized thermogravimetric analyzer (PTGA) test results obtained with ZnO (sorbent D) reacting with SO2 and O2 (Yrjas et al., 1996; Konttinen et al., 1997c) indicate that when SO2 is removed, the sulfate formed via reaction (3) starts to decompose. According to Siriwardane and Woodruff (1995), this sulfate decomposition takes place also when oxygen is removed, indicating that if either of the gaseous reactants SO2 or O2, necessary for sulfate formation are not present, sulfate formation is not possible. The samples of sorbents D and E taken during the pilot-scale fluidized-bed regeneration tests (Salo et al., 1995; Konttinen et al., 1996) showed neglible sulfate sulfur accumulation into the sorbent in continuous operation. The significance of sulfate formation in a fluidized-bed reactor modeling will be studied further in this text. In the first part of the sorbent regeneration work (Konttinen et al., 1997c) the Arrhenius parameters of the uniform conversion model were determined on the basis of laboratory-scale test results of ZnS regeneration with oxygen (reaction 2). A reliable zinc sulfide conversion rate model, together with fluidized-bed mass and energy balances, can be used to predict the performance of a large-scale regeneration reactor. Due to an exothermal reaction and a narrow temperature operating window (700-750 °C) requirement in regeneration, it is essential to have a steady-state kinetic model for reactor sizing as well as a dynamic model for process control design. In this paper the application of the ZnS regeneration rate model into HTHP pilot-scale fluidizedbed regeneration test data will be reported and discussed. Experimental Section Table 1 shows the properties of the zinc titanate sorbents D and E that were tested in HTHP pilot-scale fluidized-bed sulfur removal reactors, both for fresh and fluidized-bed samples (Salo et al., 1995). Figure 2 shows a closer schematic diagram of the pilot-scale regeneration reactor. The insulated reactor chamber is located inside a pressure vessel. In the beginning of the regeneration the reactor chamber was pressurized and preheated with regeneration gases (air and steam or nitrogen) at 450-530 °C and at the same pressure as the sulfidation reactor, i.e., 10-20 bar. The gases enter through a porous plate located at the bottom of the reactor which acts as the fluidization grid. After the regeneration reactor was preheated, 20-50 kg of
Figure 2. Schematic diagram of the pilot-scale fluidized-bed regeneration reactor and the related auxiliary equipment. The symbols to be used later in reaction rate parameter determination and modeling are included.
the zinc sulfide containing sorbent (as a result of sulfur capture reaction 1) was transported pneumatically from the sulfidation reactor to the regeneration reactor. The oxygen (from air) causes the exothermal regeneration reaction 2 to start, increasing the temperature of the regeneration reactor from about 500 °C up to 700-800 °C. As the temperature of the bed rises, measurable amounts of gaseous SO2 are formed. Figure 3 shows the temperature increase and SO2 release as a function of time with sorbent D from two different pilot-scale test periods during the regeneration reactor start-up. (In this test oxygen measurement was not yet available.) Figure 4 shows the SO2 release, oxygen content in the off-gas, and the temperature increase with sorbent E. After the regeneration reactor bed reaches a temperature of 650-750 °C, the continuous transport of the sulfided sorbent from the sulfidation reactor to the regeneration reactor and the regenerated sorbent from the regeneration reactor to the sulfidation reactor is started. The regeneration reactor off-gases are cooled to 250 °C immediately after leaving the reactor pressure vessel. A slipstream sample of the regenator off-gas containing SO2 and O2 can be taken for continuous gas measurement after filtration and pressure reduction. Figures 3 and 4 show the bed temperature for three different vertical positions. The temperatures at two lower locations (0.14 and 0.38 m) follow each other very closely, while the temperature at the highest location (1.13 m) is clearly lower when having a smaller bed inventory. The fluidized-bed height in Figure 4, period 2 was about 0.7 m, which indicates that the temperature at 1.13 m was the temperature of the gas in the freeboard section above the fluidized bed. The uniform vertical temperature profile during constant bed inven-
Ind. Eng. Chem. Res., Vol. 36, No. 12, 1997 5441
Figure 5. Dynamic response of the regeneration reactor O2 measurement in step response test at 500 °C and 15 bar.
Figure 3. Pilot-scale fluidized-bed regeneration reactor performance data during two start-up periods with sorbent D at 15 bar.
shows the result of an O2 step response test, in which the regeneration reactor air inlet control valve was closed at time t ) 0. The test was accomplished with a mixture of preheated air and N2 at 500 °C and 15 bar without a sorbent in the regeneration reactor. The slope of the step response test data in Figure 5 (on the righthand side y-axis) indicates that the values of gas measurements have to be shifted by approximately 2 min in order to match with other data. The process measurements data is processed as 1 min average values by the data collection system; therefore, the effect of the 1-min mixing period shown in Figure 5 has no practical importance. External Mass Transfer in Pilot-Scale Tests
Figure 4. Pilot-scale fluidized-bed regeneration reactor performance data during two start-up periods with sorbent E at 15 bar.
tory of the fluidized bed indicates that the mixing of the solids in the bed region is very good. The long pipelines from the regeneration reactor, in addition to gas measurement tubing, give a dead time in SO2 and O2 measurements which has to be accounted for before plotting gas measurement values in the same time scale with other process measurements. Figure 5
The basic approach toward the modeling of the chemical performance of fluidized-bed reactors is to describe the flow of gas and solids by separate phases. Usually the reactive gas in its bulk concentration is one phase and the reactive solid has its own phase (Froment and Bischoff, 1990; Kunii and Levenspiel, 1991; Gupta and Gangwal, 1992). Kunii and Levenspiel (1991) have introduced a three-phase model, where the gas flow through the fluidized bed of solids is assumed to follow plug-flow behavior and the solid phase is perfectly stirred, which means that the concentration of solid reactant and the solid conversion level are not dependent on bed position. In the model, the effect of three phases in the bed can be lumped into one gas conversion rate parameter, which is a combination of the reaction rate constant and mass-transfer effects between the three phases for the purpose of calculating the concentration of gas at the bed exit. However, in order to apply the Kunii-Levenspiel model, several empirical input parameters have to be determined, such as the bubble diameter (as a function of bed height) and minimum fluidization velocity of the solid particles in a fluidizedbed. Empirical correlations for the determination of these parameters can be found in the literature (Froment and Bischoff, 1990; Kunii and Levenspiel, 1991), but the modeling results are very sensitive to their values. For best results, these hydrodynamic input parameters should be determined experimentally, which in this study was not possible.
5442 Ind. Eng. Chem. Res., Vol. 36, No. 12, 1997
On the basis of the uniform vertical temperature profiles of the pilot-scale fluidized bed shown in Figures 3 and 4, it is assumed that the temperature of the ZnScontaining particles is uniform at any position of the bed. This is supported by the assumptions of Kunii and Levenspiel (1991) in corresponding situations.
dioxide at the fluidized-bed exit as a result of reaction 2 can be calculated with the following:
ySO2,e )
Application of Uniform Conversion Model for a Large-Scale Fluidized-Bed Reactor Equations for Obtaining the Rate Parameters from Pilot-Scale Data. The uniform conversion model used to describe the rate of the ZnS reaction (Konttinen et al., 1997c) has a constant bulk-gas concentration as input. It cannot be directly applied into the fluidizedbed reactor data, because the gas concentration in a fluidized-bed reactor varies as a function of height. Kunii and Levenspiel (1991) and Szekely et al. (1976) suggest different forms of average concentration between the gas inlet and exit concentration for steadystate calculations, followed by the calculation of the performance of the reactor by iteration between gas and solid molar balances. The predictions will lead to incorrect results if the solid in the bed is so reactive that only the lower part of the bed is reacting, which means that the upper part of the bed is practically free of reactive gas. Another way to describe the nonideal flow of the gas is to model it as several perfectly mixed tanks in series (Levenspiel, 1989). In the following, this idea is applied for the experimental results of batchwise operation of the pilot-scale regeneration reactor (Figures 3 and 4). For each individual tank the following mass balance should be satisfied (1.5 mol of oxygen gas reacts with 1 mol of solid zinc sulfide (in sorbent particle)):
(
)
mol of O2 in × gas conversion ) time solid conversion (4) 1.5(mol of ZnS) time
(
)
By applying the uniform conversion model as solid conversion rate expression the same way as presented in sulfidation reactor modeling (Konttinen et al., 1997b), the mass balance can be shown to give:
X O 2j )
R 1+R
where R )
1.5nZnSkf(1 - XZnS)CO2,0 FO2,0N
)
yO2,0 (1 - exp(-R)), 1.5
(
where R )
where yO2,0 is the volume fraction of oxygen at the regeneration reactor inlet and kf is a global reaction rate constant (cm3/(mol s)) including chemical reaction and possible mass-transfer effects. The illustration of the symbols used in eq 6 can be found in Figure 2. Large-Scale Steady-State Regeneration Reactor Model. Equation 6 can be applied into the modeling of a continuously operated large-scale fluidized-bed regeneration reactor by the following way. A mass balance of the ZnS in sulfided zinc titanate sorbent particles can be written as follows:
dnZnS ) FZnS,0 - FZnS,e - rZnS dt
where FO2,0 is the oxygen initial inlet flow (mol/s), XO2j the conversion of oxygen in jth tank, nZnS the amount of zinc sulfide in the fluidized bed (mol), XZnS the overall conversion of solid (zinc sulfide) in the bed, t the time (s), and N the number of vertical tanks in series. Vgas is the volumetric flow of gas through the fluidized bed (cm3/s). Because of nitrogen and steam diluents in the regeneration inlet gas, the gas expansion factor due to the regeneration reaction at fluidized-bed reactor conditions is 0.99-1, so the assumption of constant volumetric flow is within experimental error. In order to eliminate parameter N, the common assumption of the plug flow of gases through the fluidized bed (Szekely et al., 1976; Kunii and Levenspiel, 1991) can be applied, the same way as presented earlier (Konttinen et al., 1997b). The volume fraction of sulfur
(7)
where FZnS,0 and FZnS,e are the molar flows of ZnS entering and exiting the reactor (mol/s) and rZnS is the molar rate of sulfided zinc titanate reacted to form SO2 (mol/s). At steady-state conditions, the time derivative equals 0. By replacing
FZnS,e ) FZnS,0(1 - XZnS) and rZnS )
FO2,0XO2 1.5
(8)
we get:
1.5FZnS,0XZnS ) FO2,0XO2
(9)
where XO2 is the total conversion of O2 to SO2 at the fluidized-bed exit. It can be determined by rearranging eq 6:
XO2 ) (1 - exp(-R)) 1.5nZnSkf(1 - XZnS) (5) VgasN
)
1.5nZnSkf(1 - XZnS) (6) Vgas
(10)
In a perfectly stirred regeneration reactor with a steady sorbent circulation rate, the average concentration of ZnS in the fluidized bed equals the ZnS concentration at the exit of the reactor. Using symbols, the molar amount of ZnS in the fluidized bed at steady-state conditions can be determined with the following equation:
FZnS,e nZnS ) F nZnS ) FZnS,0(1 - XZnS)τbed msorb Wbed where τbed )
Wbed (11) msorb
where Wbed is the total bed inventory (g) and msorb is the sorbent circulation rate (g/s) in between sulfidation and sorbent regeneration reactors. Thus the parameter τbed equals the mean residence time of the sorbent (s)
Ind. Eng. Chem. Res., Vol. 36, No. 12, 1997 5443
in the fluidized-bed. Equation 6 can now be rewritten (see Figure 2):
XZnS )
FO2,0(1 - exp(-R)) 1.5FZnS,0
where R )
(
)
1.5FZnS,0(1 - XZnS)τbedkf(1 - XZnS) (12) Vgas
For steady-state calculations XZnS has to be solved iteratively, by using Newton’s method, for example. In steady-state conditions, the effect of residence time distribution (RTD) of the sorbent particles in the fluidized-bed reactor is recommended (Froment and Bischoff, 1990; Kunii and Levenspiel, 1991) to be taken into account. The RTD function of the perfectly stirred reactor is
E(t) )
1 -t/τbed e τbed
(13)
The mean conversion of the ZnS in the fluidized-bed reactor is
1-X h ZnS )
∫0∞(1 - XZnS)E(t) dt
Combining eqs 12, 13, and 14 we get
1-X h ZnS )
(
∫0∞ 1 -
(14)
)
FO2,0(1 - exp(-R)) 1 e-t/τbed dt 1.5FZnS,0 τbed (15)
However, since XZnS is included inside the exponential term R, an analytical solution of eq 15 is not possible. The effect of RTD on the perfectly stirred regeneration reactor performance should not be significant (Kunii and Levenspiel, 1991), and therefore as the first approximation of regeneration reactor steady-state performance in this paper, eq 12 will be used instead of eq 15. Rate Constants from Pilot-Scale Fluidized-Bed Data Equation 10 can rearranged to give the values of a global reaction rate constant (kf) at pilot-scale regeneration conditions (see Figure 2):
kf )
-ln(1 - XO2)Vgas 1.5nZnS(1 - XZnS)
where nZnS ) nZnS,0(1 - XZnS) (16)
The values of process variables in equation 16 are available from the experimental data. By using the data of Figures 3 and 4, the temperature dependence of the rate constant can be studied. The conversion of oxygen in the regeneration fluidized bed (XO2) is calculated with the following:
XO2 )
1.5ySO2,e yO2,0
(17)
where ySO2,e is the fractional content of SO2 at the regenerator exit, (plotted in Figures 3 and 4). The variable yO2,0 and gas volumetric flow (Vgas) were calculated using the gas flow measurements (air + steam/nitrogen). The initial zinc sulfide content of the
Figure 6. Arrhenius plot of the ZnS global regeneration rate constants obtained from pilot-scale data with sorbents D and E at 15 bar.
regeneration reactor bed (nZnS,0) was determined from the sulfur content analysis of bed material samples taken in the beginning of the regeneration tests. XZnS (as a function of cumulative moles of gaseous SO2 produced) was determined from the gas-solid molar balance during pilot test periods. This method has the advantage that both laboratory and pilot-scale data can be modeled using the same reaction rate parameter, thus making it possible to compare results obtained with different-scale test equipment. An example of the use of eqs 16 and 17 in parameter determination can be given by using the experimental data of sorbent E from the start-up period 1 (Figure 4): At the bed temperature of 600 °C (15 bar) the conversion of O2 to SO2 (XO2, eq 17) is 0.74 and Vgas is 11634 cm3/s. The initial zinc sulfide content of the bed (related with bed inventory of 64 kg) is 281 mol, and the conversion of zinc sulfide to zinc oxide at 600 °C is 0.44. By using these values the reaction rate parameter kf (eq 16) equals 118 cm3/(mol s). Figure 6 shows the temperature dependence of the global reaction rate parameters obtained using the pilotscale fluidized-bed data of Figure 3 (sorbent D) and Figure 4 (E) in eqs 16 and 17. The Arrhenius plot shows clear temperature dependence. On the basis of the properties of sorbents D and E shown in Table 1, it can be assumed that the reactivity of the sorbents in regeneration does not differ very much. The uniform conversion rate equation (Konttinen et al., 1997c) also assumes that the total pressure affects the reaction rate only by changing the O2 partial pressure. In comparison with the Arrhenius plot of laboratory-scale data reported previously (Konttinen et al., 1997c), the values of global rate constants given here are generally 1 order of magnitude lower. A probable reason for this could be the mass-transfer limitation between bubble gas and solid in a fluidized bed, such as indicated by the KuniiLevenspiel model. However, the apparent activation energy of the global reaction rate parameter in the pilotscale reactor is 200-210 kJ/mol for both sorbents D and E as bed material. Such high values usually cannot be
5444 Ind. Eng. Chem. Res., Vol. 36, No. 12, 1997
Figure 7. The results of a sensitivity analysis of a large-scale continuously operated fluidized-bed reactor model in sulfided zinc titanate regeneration. The regeneration level and fractional conversion of O2 are plotted as a function of reactor bed temperature.
obtained for mass-transfer control. The value of the activation energy in laboratory-scale tests was found to be about 140 kJ/mol, which is at the same level as in pilot scale tests, indicating chemical reaction to be the rate-controlling step in both scales. Sensitivity Analysis of a Large-Scale Reactor Model The sensitivity of the regeneration reactor model performance on the different order of magnitude parameters can be studied with the help of eq 12. Figure 7 shows some results of the study. Figure 7 includes a new parameter plotted on the left-hand side y-axis, which is
regeneration level ) FS,e/FZnO
(18)
where FS,e is the flow of unreacted sulfur at the regeneration reactor exit (mol/s) and FZnO the corresponding total flow of zinc oxide (or zinc) (mol/s) in the zinc titanate sorbent at the regeneration reactor exit (see Figure 2). In Figure 7, the regeneration level and corresponding conversion of O2 to SO2 (XO2, eq 10) through the fluidized bed are plotted as a function of regeneration bed temperature. Actually, the regeneration level equals the fractional conversion level of zinc titanate to ZnS in sorbent circulation in between sulfur removal and regeneration reactors. For a large-scale reactor application, the inlet molar flow of ZnS has been selected to be 1 mol/s and the mean residence time of the sorbent (τbed, eq 11) in the reactor as 1 h. In IGCC applications, the process pressure is typically about 20 bar and the O2 inlet content about 3-5 vol %. Figure 7 indicates that generally the sulfur level in the sorbent after the regenerator decreases as the bed temperature increases. With this set of input values, the initial fractional sulfur level of the sulfided zinc titanate entering the regeneration reactor is 0.3 which can be seen to decrease below 0.1 at highest bed temperatures. The fractional conversion of O2 (on the right-hand side y-axis) increases respectively and reaches 1 at temperatures of 700-800 °C, which are actually the practical reactor operating temperatures. Significant differences can be observed at temperatures below 700 °C in the
Figure 8. The results of a sensitivity analysis of a large-scale sulfided zinc titanate regeneration reactor model with more favorable input values for regeneration than in Figure 7.
model performance when using reaction rate constant values from TGA tests: 1.44 × 1011 exp(-140 000/RT) (cm3/(mol s)) (Konttinen et al., 1997c), in comparison with the values of kf of ZT-4-L from pilot-scale tests: 8.36 × 1013 × exp(-201 740/RT) (cm3/(mol s)) (Figure 6). Figure 8 shows the results of a corresponding study, using a set of input values that are extremely favorable for regeneration. The sulfur flow to the regenerator (FZnS,0) is decreased to 30% of that used in Figure 7. The residence time of the sorbent in the bed has been increased to 13.7 h. The fractional initial sulfur content in the sulfided zinc titanate sorbent entering the regenerator reactor is increased to 50%. It can be seen that the operation temperature range for reaching a high regeneration level of the reactor is wider than in Figure 7, covering temperatures of 600-800 °C. The differences of the reactor model performance with different reaction rate parameters are significant at temperatures below 650 °C. The sensitivity analysis of the regeneration reactor steady-state model by eq 12 indicates that the parameter values obtained from different sources (k from TGA and kf from pilot-scale) create differences in model performance. However, as Figures 7 and 8 show, at the practical steady-state regeneration temperatures of 700-800 °C all O2 is consumed and the unreacted sulfur level in the sorbent exiting the reactor is not controlled by reaction kinetics. At temperatures of 550-650 °C, the values of the regeneration level (eq 18) using pilotscale reaction rate constants (kf) are 1.2-2.3 times higher than the corresponding values obtained using laboratory-scale constants (k). The parameter values of XO2 with laboratory-scale rate constants are 1.1-2.4 times higher than those with pilot-scale constants. When considering that the reaction rate constants at 550-650 °C from laboratory-scale tests are 5-14 times higher than those obtained from pilot-scale tests, the order of magnitude effect of this difference to the steadystate model performance is significantly lower. Actually, at temperatures of 450-650 °C, Figures 7 and 8 give a performance window of the regeneration reactor, which can be used as guideline information for further design. On the basis of the empirical experience of sulfided zinc titanate regeneration in bench- and pilotscale fluidized beds (Gupta and Gangwal, 1992; Abbasian et al., 1994; Konttinen et al., 1996), it can be
Ind. Eng. Chem. Res., Vol. 36, No. 12, 1997 5445
suspected that the reaction rate constant parameters preferred to be used in modeling should be the same or close to the values determined from laboratory-scale. In addition to the material balance, an energy balance will have to be employed for steady-state calculations of a regenation reactor model, due to exothermal reaction 2. The energy balance calculations related with results in Figures 7 and 8 indicate that the bed temperatures of 450-800 °C can be achieved by using 300-500 °C as temperatures of the gases and solids entering the regeneration reactor. The operation of the pilot-scale reactor was not optimized for reaction rate parameter determination. In pilot-scale tests it is possible that some part of the oxygen entering the fluidized bed may have been captured by ZnO to form sulfate, not shown in gas measurements, resulting in lower values for the reaction rate constants than those in laboratory-scale tests. Indications on this were found when comparing the measured oxygen contents into and out of the regeneration reactor during test periods such as those presented in Figure 4. Due to inaccuracy in the O2 mass balance in the pilot-scale tests, any further study of this phenomenon was omitted. On the basis of the literature and the laboratory-scale test results reported earlier (Konttinen et al., 1997c) the rate of O2 in reaction to form zinc sulfate, in comparison with the rate of SO2 release, can be assumed insignificant. It is stated that the regeneration of sulfided zinc titanate in a large-scale fluidized bed is kinetically controlled with a relatively high activation energy. The difference between reaction rate constants obtained from laboratory- and pilot-scale tests was studied with the help of a sensitivity analysis of a steady-state regeneration reactor model. The differences in the model performance with different reaction rate parameters were not significant at temperatures 700-800 °C, which is the practical operating temperature window of a large-scale reactor. The amount of the sulfided zinc titanate sample in laboratory-scale tests was 50 mg (Konttinen et al., 1997c) and the amount of sulfided zinc titanate in the pilot-scale fluidized bed was 50-80 kg; thus the scale-up factor between the results obtained is 106. The reasons for the differences and uncertainties will be studied further in the future test runs of largescale regeneration fluidized-bed reactors. The Possibility of Zinc Sulfate Formation in a Fluidized Bed The oxygen partial pressure at different temperatures as a result of SO2 release can be obtained by using eq 10 and the Arrhenius plot in Figure 6. The partial pressure of oxygen decreases and the partial pressure of SO2 increases as a function of bed height. The equilibrium partial pressure of O2 for sulfate formation reaction 3 can be obtained from
pO2,h(eq) )
(
)
2
pSO2,hK
(pO2,0 - pO2,h) 1.5
1
and
where
pSO2,h )
pO2,h ) yO2,0(1 - XO2,h)p (19)
where pO2,h and pSO2,h are the partial pressures of O2 and SO2 at certain bed height as a result of the regeneration reaction; pO2,h(eq) is the equilibrium partial pressure of O2 as a result of sulfate formation at a given
SO2 partial pressure and K is the equilibrium constant for the sulfate formation reaction 3. According to eqs 10 and 19, the partial pressure of O2 can go below to the equilibrium value required for sulfate formation at actual regeneration temperatures (725 °C) at higher parts of bed, because all O2 is consumed in the lower parts due to SO2 release. By assuming that the concentration and temperature of the solid reactant in the bed is uniform (supported by bed temperatures in Figures 3 and 4), the bed material spends most of its time at the region where O2 is not available. This makes it possible for any zinc sulfate possibly formed via reaction 3 at lower temperatures to decompose at the actual regeneration temperatures of 700-750 °C. The zinc titanate sorbent samples from the regeneration reactor taken during the tests showed no sign of permanent sulfate formation (Salo et al., 1995; Konttinen et al., 1996). All this indicates that the regeneration of a ZnS-containing sorbent in a fluidized bed is very favorable. Conclusions The uniform conversion model was applied for modeling of ZnS regeneration. The parameters of the model were determined earlier on the basis of laboratory-scale results. A method using the model for pilot-scale fluidized-bed reactor is presented. In this method it is assumed, based on literature, that the flow of gas through the bed of solids follows plug flow behavior. The apparent activation energy for the global reaction rate constant is 200-210 kJ/mol for two different sorbents, which is higher than that obtained earlier with laboratory-scale results: 140 kJ/mol. Due to the differences in the reaction rate parameters, a sensitivity analysis of a large-scale steady-state regeneration reactor model is performed using rate constants obtained from pilot-scale in comparison with constants from laboratory-scale. Differences in the model performance can be observed, but at practical steady-state operation temperatures of 700-800 °C, these appear not to be significant, since the rate of sulfided zinc titanate regeneration is not controlled by reaction kinetics. With respect to the scale-up factor of 106 between laboratory- and pilot-scale results, the model and its parameters at temperatures of 550-650 °C give a performance window of a large-scale regeneration reactor that can be used as guideline information. The reasons for the differences will be studied further in the future test runs of large-scale regeneration fluidized-bed reactors. It is stated that the reason for neglible zinc sulfate formation in regeneration in a pilot-scale fluidized bed is that at sufficiently high temperatures the partial pressure of oxygen (required for sulfate formation) at the exit of the bed is practically zero so all the possible sulfate formed at lower temperatures will decompose. Acknowledgment The work on the development and testing of regenerable sulfur removal sorbents and processes has been carried out as a part of the Combustion and Gasification Research Program “LIEKKI 2” and Energy and Environmental Technology Program “SIHTI 2”, which are partly financed by the Technology Development Centre of Finland. The pilot-scale testing of sorbents D (UCI5) and E (ZT-4-L) was a part of the Cooperative Research and Development Agreement between Envi-
5446 Ind. Eng. Chem. Res., Vol. 36, No. 12, 1997
ropower Inc. (the predecessor of Carbona Corporation) and the U.S. Department of Energy. Mr. Thomas Dorchak of the Federal Energy Technology Center, Morgantown, WV, is gratefully acknowledged for his valuable contribution. The development of the sulfur removal reaction models was funded partly by the Imatran Voima Foundation. The authors acknowledge the valuable contribution of Dr. Javad Abbasian of the Institute of Gas Technology, Des Plaines, IL, in arranging the chemical and physical analysis of the fresh and used sorbents. Notation C ) gas concentration (mol/cm3) eq ) thermodynamic equilibium conditions F ) flow of gaseous or solid reactant (mol/s) h ) fluidized bed height (m) k, kf ) kinetic reaction rate constant according to UCM (cm3/(mol s)) and global reaction rate constant in eq 6 (cm3/(mol s)) msorb ) sorbent circulation rate (g/s) nZnS ) the amount of zinc sulfide in fluidized bed (mol) N ) total number of vertical slices in series p ) pressure (bar) r ) reaction rate, (mol/s) t ) time (s) Vgas ) gas volumetric flow (m3/s) Wbed ) bed inventory (g) X ) fractional conversion XO2,j ) conversion of reactant oxygen gas in vertical slice j y ) volume fraction of gas and also stoichiometric coefficient in eq 1 R ) parameter defined by eqs 5, 6, and 12 τbed ) solid residence time parameter (s) Subscripts e ) at fluidized bed exit f ) at fluidization conditions g ) gas phase h ) at fluidized bed height h j ) order number of the vertical slice s ) solid phase 0 ) initial
Literature Cited Abbasian, J.; Bachta, R. P.; Wangerow, J. R.; Mojtahedi, W.; Salo, K. Advanced High-Pressure Bench-Scale Reactor for Testing with Hot Corrosive Gases. Ind. Eng. Chem. Res. 1994, 33, 9195. Froment, G. F.; Bischoff, K. B. Chemical Reactor Analysis and Design, 2nd Ed.; John Wiley & Sons: New York, 1990. Gupta, R. P.; Gangwal, S. K. Enhanced Durability of Desulfurization Sorbents for Fluidized Bed Applications. Topical Report, 96U-4274, Contract DE-AC21-88MC25006; Research Triangle Institute: Research Triangle Park, NC, November 1992.
Harrison, D. Control of Gaseous Contaminants in IGCC Processes, An Overview. In Twelfth Annual International Pittsburgh Coal Conference, Pittsburgh, PA, September 11-15, 1995; Chiang, S.-H., Ed.; Pittsburgh Coal Conference: Pittsburgh, PA, 1995; pp 1047-1052. Konttinen, J. T.; Mojtahedi, W.; Abbasian J. Coal Gas Desulfurization at the 15 MWth Pressurized Gasification Pilot Plant. In High Temperature Gas Cleaning; Schmidt, E., Ga¨ng, P., Dittler A., Eds.; G. Braun Printconsult GmbH: Karlsruhe, Germany, September 1996; pp 756-766. Konttinen, J. T.; Zevenhoven, C. A. P.; Hupa, M. M. Hot Gas Desulfurization with Zinc Titanate Sorbents in a Fluidized Bed. 1. Determination of Sorbent Particle Conversion Rate Model Parameters. Ind. Eng. Chem. Res. 1997a, 36 (6), 2332-2339. Konttinen, J. T.; Zevenhoven, C. A. P.; Hupa, M. M. Hot Gas Desulfurization with Zinc Titanate Sorbents in a Fluidized Bed. 2. Reactor Model Ind. Eng. Chem. Res. 1997b, 36 (6), 23402345. Konttinen, J. T.; Zevenhoven, C. A. P.; Yrjas, P.; Hupa, M. M. Modeling of Sulfided Zinc Titanate Regeneration in a FluidizedBed Reactor. 1. Determination of the Solid Conversion Rate Model Parameters. Ind. Eng. Chem. Res. 1997c, 36, 5432-5438. Kunii, D.; Levenspiel, O. Fluidization Engineering, 2nd ed.; Butterworth-Heinemann: Boston, 1991. Levenspiel, O. The Chemical Reactor Omnibook. OSU Book Stores: Columbus, OH, 1989. Lew, S. High-Temperature Sulfidation and Reduction of Zinc Titanate and Zinc Oxide Sorbents. Ph.D. Dissertation, Massachusetts Institute of Technology, October 1990. Portzer, J.; Gangwal, S. K.; Dorchak, T. Slipstream Testing of HotGas Desulfurization with Sulfur Recovery. In Twelfth Annual International Pittsburgh Coal Conference, Pittsburgh, PA, September 11-15, 1995; p 1073. Salo, K.; Konttinen, J. T.; Ghazanfari, R.; Feher, G.; Lehtovaara, A. Mojtahedi W. Pilot Scale Experience on IGCC Hot Gas Cleanup. In Proceedings of the Advanced Coal-Fired Power Systems ‘95 Review Meeting, Volume 1, June 1995; U.S. Department of Energy, Morgantown Energy Technology Center: Morgantown, WV, 1995. Siriwardane, R. V.; Woodruff, S. FTIR Characterization of the Interaction of Oxygen with Zinc Sulfide. Ind. Eng. Chem. Res. 1995, 34, 699-702. Szekely, J.; Evans, J. W.; Sohn H. Y. Gas-Solid Reactions; Academic Press: New York, 1976. Yrjas, K. P.; Hupa, M. M.; Konttinen, J. T. Regeneration of Sulfided Zn-Titanate by Oxidation under Pressure. In High Temperature Gas Cleaning; Schmidt, E., Ga¨ng, P., Dittler A., Eds.; G. Braun Printconsult GmbH: Karlsruhe, Germany, September 1996; pp 514-527.
Received for review January 15, 1997 Revised manuscript received September 2, 1997 Accepted September 2, 1997X IE9700369
X Abstract published in Advance ACS Abstracts, October 15, 1997.
