Ammonia Pyrolysis and Oxidation in the Claus Furnace - American

The modified Claus process is commonly used in oil refining and gas processing ... typical modified Claus plant feed of H2S and CO2, NH3 is also often...
0 downloads 0 Views 88KB Size
Download PDF
144

Ind. Eng. Chem. Res. 2001, 40, 144-151

Ammonia Pyrolysis and Oxidation in the Claus Furnace W. D. Monnery, K. A. Hawboldt, A. E. Pollock, and W. Y. Svrcek* Department of Chemical and Petroleum Engineering, The University of Calgary, Calgary, Alberta T2N 1N4, Canada

The modified Claus process is commonly used in oil refining and gas processing to recover sulfur and destroy contaminants formed in upstream processing. In oil refining, in addition to the typical modified Claus plant feed of H2S and CO2, NH3 is also often present. NH3 is a process contaminant and must be destroyed in the front-end furnace of the modified Claus plant, otherwise, it poses a risk of poisoning the catalyst beds and plugging off downstream equipment because of the formation of ammonium salts. In this paper, the pyrolysis and oxidation of NH3 was studied under Claus furnace temperatures and residence times. Experimental data was taken, and new reaction rate expressions were developed for NH3 pyrolysis and oxidation. The derived rate expressions are outlined as follows: the NH3 pyrolysis rate expression r ) A exp(-Ea/RT)PNH31.25, where A is 0.004 21 mol s-1 atm-1.25 cm-3 and Ea is 16.5 kcal mol-1, and the NH3 oxidation rate expression r ) A exp(-Ea/RT)PNH3PO20.75, where A is 4430 mol s-1 atm-1.75 cm-3 and Ea is 40.0 kcal mol-1. The rate expression for NH3 pyrolysis matched experimental data within 13% and matched well with published data. The rate expression for NH3 oxidation matched experimental data within 10%. Introduction The modified Claus sulfur recovery process is the most common process used to remove sulfur from acid gas streams occurring in oil and gas processes. This process is comprised of a high-temperature furnace followed by catalytic reactor(s). The reactions occurring in the furnace are numerous, and several authors have attempted to delineate the important ones (Clark et al., 1997). As part of our ongoing study, we have previously published kinetic data and reaction rate expressions for H2S cracking/reassociation (Hawboldt et al., 1999a) and the Claus reactions (Hawboldt et al., 1999b). The reaction furnace has three functions: the conversion of one-third of the H2S to SO2 for downstream catalytic processing, the destruction of any contaminants which may foul downstream equipment, and the production of elemental sulfur, which can account for up to 70% of the inlet sulfur for a “straight-throughtype plant”. The types of contaminants in the sour gas stream depend on the source of the feed gas. For instance, in oil refinery operations, NH3 is formed as a byproduct of denitrogenation operations, such as hydrocracking and hydrotreating. Subsequent sour water stripping results in a sour gas feed stream containing NH3, which is then directed to the sulfur recovery facility for destruction. The desired NH3 destruction reactions in the furnace are the complete pyrolysis and/or oxidation of NH3

2NH3 T N2 + 3H2

(1)

3 2NH3 + O2 T N2 + 3H2O 2

(2)

Under excess oxygen conditions and incomplete mixing, the following reaction may also occur (Clark et al., 1997): * To whom correspondence should be addressed.

5 2NH3 + O2 f 2NO + 3H2O 2

(3)

Incomplete pyrolysis or combustion of NH3 in the furnace results in NH3 and NO carryover into the catalyst beds. Ammonia can form ammonium salts, which can plug or foul the catalyst beds, other equipment, or piping. Although the formation of SO3 occurs in the catalyst bed regardless of the presence of NO, the presence of NO in the beds acts as a catalyst for the conversion of SO2 to SO3, which in turn causes catalyst sulfation (Garside and Phillips, 1962). Of the primary causes of catalyst activity loss, catalyst sulfation is regarded as the most significant (Grancher, 1978). It is therefore critical to convert as much of the NH3 to N2, H2, and H2O as possible. One of the obstacles to designing and operating the furnace properly, i.e., determining optimal residence times and temperatures, is the absence of kinetic rate expressions for the key reactions (Monnery et al., 1993). For ammonia destruction, an empirical rule of thumb that is generally agreed with in industry is that temperatures greater than 1200-1250 °C are required. However, there appears to be some debate as to how much residence time is required and how ammonia oxidation competes for oxygen versus hydrocarbon and H2S oxidation (Johnson and Rempe, 1997; Goar, 1994). As such, ammonia pyrolysis and oxidation kinetic data are badly needed. In this paper, we will present new experimental data and new kinetic rate expressions for the NH3 pyrolysis and NH3 oxidation reactions at Claus front-end furnace (FEF) conditions and residence times. Literature Review NH3 Pyrolysis. Davidson et al. (1990) performed a series of high-temperature pyrolysis experiments at a temperature range of 2000-3200 K, a pressure range of 0.8-1.1 atm, and an NH3 concentration range of 0.11.0%. Under these conditions, NH3 is completely con-

10.1021/ie990764r CCC: $20.00 © 2001 American Chemical Society Published on Web 12/01/2000

Ind. Eng. Chem. Res., Vol. 40, No. 1, 2001 145

Figure 1. Apparatus schematic.

sumed within a 1 ms residence time. A detailed reaction mechanism was proposed, consisting of 21 free-radical reactions. These experiments were performed behind reflected shock waves. The severity of the temperatures and the mechanism of the experiment (shock tube) prevent direct comparison with data generated in our experiments. However, these studies do give an indication of trends and therefore the significance of intermediate species (NH2, NH, N2H2, etc.). Clark et al. (1998) conducted a study on NH3 pyrolysis at FEF conditions to determine the effects of other chemical species occurring at furnace temperatures, pressures, and residence times. In general, both H2S and H2O were found to inhibit NH3 pyrolysis, whereas SO2 enhances NH3 pyrolysis. Their results also showed oxidation being much more rapid than pyrolysis. NH3 Oxidation. Published studies have indicated that the oxidation of NH3 is a much more rapid process than pyrolysis (Clark et al., 1998; Lindstedt and Selim, 1994; Miller et al., 1983; Fujii et al., 1981). In the Fujii et al. (1981) study, the reaction was investigated at temperatures between 810 and 2100 K, pressures from 1.1 to 8.4 atm, NH3 concentrations from 0.5 to 75%, and O2 concentrations from 0.5 to 90%, with the balance of the gas being argon. Miller et al. (1983) used a reaction model consisting of 98 reactions to determine the reaction mechanism for NH3 oxidation. The model predictions were compared with experimental data generated by MacLean and Wagner (1967). The developed kinetic model predictions were in good agreement for the lean to moderately rich gases. However, for the rich flames, the agreement was poor, possibly indicating inadequate modeling of ammonia pyrolysis. The model and experimental results also showed very low concentrations (approximately 0.001 mole fraction) of NO in the post-flame gases. In 1994, Lindstedt and Selim proposed a reaction mechanism consisting of five steps based on experimental data generated by MacLean and Wagner (1967) and Vandooren (1992). Both Lindstedt and Selim (1994) and Miller et al. (1983) based their reaction mechanisms on flame studies, and therefore temperatures were not stated. Other studies on NH3 oxidation have focused on the reduction of NO by NH3 in NH3-NO-O2 systems. Lyon and Benn (1978) were the first to conduct a kinetic study of the reduction of NO by NH3 in the presence of O2 at temperatures between 872 and 980 °C and pressures between 1.07 and 2.14 atm. Provided O2 was present in large excess compared to NO, NH3 addition was

effective in reducing NO at all temperatures and pressures of the study. However, when the NH3 concentration was comparable to O2, the reduction of NO was inhibited. Miller and Bowman (1989) came to the same conclusions and, in addition, found that the process occurs in the narrow temperature range of 1100-1400 K in the absence of other additives. They also found that the presence of H2O had a slight inhibiting effect on NO reduction. Dou et al. (1992) proposed a simple kinetic model that predicts the rate of NO and NH3 conversion in the NO-NH3-O2 system. Major conclusions from the Dou et al. (1992) study indicate that NH3 conversion increases with temperature, whereas NO conversion increases with temperature until approximately 1200 K, after which conversion decreases. Furthermore, NO conversion rapidly levels off within 50 ms, indicating that longer residence times will not affect NO conversion. In their experiments, O2 was in large excess. Clark et al. (1998) performed oxidation studies on NH3 at residence times between 0.34 and 0.55 s and inlet concentrations of NH3 and O2 of 4 and 3%, respectively. The study found that between 700 and 1100 °C conversion of NH3 is below 10%, whereas at temperatures greater than 1100 °C, the conversion increases to between 60 and 100%. New Experimental Data Generation Experimental Apparatus. The experimental apparatus, in Figure 1, consists of pressurized gas cylinders delivering reactants through a controlled flow system to a flow reactor housed in a high-temperature furnace, a quench system, and a gas chromatograph for analyzing reactants and products. The apparatus was designed by Fookes (1996) specifically for the kinetic study of the gas-phase reactions occurring in the Claus FEF. Reactant gases are dropped from cylinder pressures to the reactor pressure across the regulators. Subsequently, flow rates are controlled by Linde FM4660 mass flow controllers (MFC) connected to a main console. MFCs were calibrated prior to experimental work; however, to ensure flow rates were accurately known, an Alexander Wright wet test meter was also used. Ultimately, flows were consistent to within 1.0%. Temperatures were measured with thermocouples, also accurate to within 1.0%. Quartz flow reactors were used in these experiments to avoid possible metal or ceramic catalytic effects. Each reactor is 5.0 mm in diameter, and three different

146

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

lengths were used to provide residence times from about 0.5 s to about 2.0 s, at temperatures between 800 and 1300 °C, typical of Claus front-end reaction furnace conditions. The reactors have a preheat length, followed by a mixing chamber, and then an isothermal hot zone. The reactors are coiled for two reasons: to have sufficient length to attain the desired range of residence times in a compact arrangement and to create secondary flow, which reduces axial diffusion or Poiseuille flow, thereby enhancing plug flow (Nauman, 1977; Truesdell and Alder, 1970). The reactor was designed to satisfy the conditions Cutler et al. (1988) set out for the design of a plug-flow reactor. Cutler et al. (1988) showed that, even in laminar flow, the error in deriving kinetic data from assuming plug flow is typically a few percent, up to a maximum of 11%, providing certain dimensional and heat- and mass-transfer criteria are met. Indeed, Cutler et al. (1988) showed that in attaining kinetic data there is more error incurred because of systematic error, such as thermocouple precision. Furthermore, to verify that the reactor operated in plug flow, a computational fluid dynamic (CFD) model of the reactor was developed by British Oxygen Corp.. The results of this simulation showed that the flow regime closely approximates plug flow (Hawboldt, 1998). Last, we verified the apparatus by comparing experimental results to accepted model results of ethane pyrolysis and matching these to within 5% (Hawboldt et al., 1999b). On the basis of the Cutler criteria, ethane pyrolysis results, and the results of the CFD simulation, the reactor was considered to operate in plug flow with negligible error. Product gases from the quartz reactors were quenched rapidly against water in a double pipe exchanger, typically down to less than 100 °C within 5-10 ms. This was shown to provide an accurate sample of reactor products (Hawboldt et al., 1999a). Finally, before product gases went to the gas chromatograph, they passed through a separator and filter to remove solid sulfur. A key characteristic of this reactor system is that it operates nearly isothermally, which simplifies data analysis because temperature variations within the reacting system do not have to be corrected for. Isothermality was ensured with a rapid preheat zone and by conduction of experiments under dilute conditions. By keeping reactant concentrations lower than 5%, the effect of any temperature changes due to the heat of reaction are minimized, if not eliminated (Hawboldt, 1998). Figure 2 shows the experimental temperature profile for the 3.05 m reactor as a plot of the deviation from the furnace setpoint temperature as a function of length. This figure clearly shows that the temperature rises rapidly to within 30 °C of the setpoint within the first 25 cm of length, after which the deviation falls rapidly, being within 1.2% at 1.0 m. Experimental Conditions. Pyrolysis. Our experimental conditions were at much lower temperatures than those of the Davidson et al. (1990) study. Thirtyseven pyrolysis experiments were performed using temperatures between 850 and 1200 °C, NH3 concentrations from 0.5 to 2.0%, and residence time ranges from 50 to 800 ms (Hawboldt, 1998). Helium was the gas chromatograph (GC) carrier gas for the majority of the experiments; however, to measure the produced N2 and H2 and close the material balance, the carrier gas was switched to argon. Oxidation. Twenty-one NH3 oxidation experiments were performed. Temperatures were varied between 850

Figure 2. Experimental measurements of the deviation of the temperature of the reactor from set point temperature along the reactor length at a flow rate of 8 SLPM.

and 1200 °C, residence times ranged from 30 to 700 ms, inlet concentrations of NH3 varied between 0.3 and 2.5%, and O2/NH3 ratios varied from 0.85 to 2.55 (Hawboldt, 1998). Argon was used as the dilution gas for this set of experiments. GC measurements of NH3 were made for all experiments, and to check the material balance in selected experiments, N2 concentrations were also measured. For stoichiometric mixtures of NH3 and N2, the agreement between NH3 consumed and N2 produced was within 10% (Hawboldt, 1998). This excellent material balance closure was also found for experiments with O2/NH3 feed ratios greater than the stoichiometric 0.75. Although some NO may be formed and the GC cannot separate NO from N2, the results of equilibrium calculations by both ourselves and Miller et al. (1983) show that the concentration of NO formed comprises a maximum of 0.001 mol % of the product. As such, the excellent material balance that was observed was valid for all experiments. New Experimental Results Pyrolysis. Figures 3 and 4 present the conversion of NH3 with temperature and residence time for the ammonia pyrolysis experiments. As the figures show, between 850 and 1050 °C, the conversion of NH3 was below 25%. At 1150 °C and residence times below 300 ms, conversions are less than 20%; however, at longer residence times, conversions increase to 55%. At 1200 °C, conversions increased substantially and varied from 27 to 83%. Material balances were verified by comparing NH3 conversion measurements with N2 and H2 composition measurements. These measurements agreed to within 10%, indicating a good material balance. Clark et al. (1998) showed similar results, in that at residence times from 0.6 to 1.0 s conversions varied from 10% at 900 °C to 75% at 1200 °C. Oxidation. Figures 5 and 6 show NH3 conversions as a function of residence time and temperature for the oxidation experiments. In general, when compared to pyrolysis, conversion of NH3 via oxidation is faster and

Ind. Eng. Chem. Res., Vol. 40, No. 1, 2001 147

Figure 3. Experimental and predicted conversion of NH3 in pyrolysis experiments as a function of temperature and residence time.

Figure 5. Experimental and predicted conversion of NH3 in oxidation experiments as a function of temperature and residence time.

Figure 4. Experimental and predicted conversion of NH3 in pyrolysis experiments as a function of temperature and residence time.

Figure 6. Experimental and predicted conversion of NH3 in oxidation experiments as a function of temperature and residence time.

occurs to a greater extent for the same temperature. For example, at 850 °C, conversion by pyrolysis never exceeds 20%, whereas at the same temperature, conversion by oxidation is observed at nearly double this value. Direct comparison with the data generated by Lyon and Benn (1978) is difficult because their experiments were performed using a large excess of O2. When compared to the Clark et al. (1998) data, our NH3 conversions were significantly greater at lower temperatures. Clark et al. (1998) predict conversions of 4-10% at residence times from 0.34 to 0.55 s in the temperature range from 700 to 1100 °C, whereas our conversions in this same range were 2-60%. The formation of NO in these experiments could not be substantiated because of the difficulty in analyzing for this component; however, equilibrium calculations suggest negligible quantities to be present.

simulated the reactor. This reactor model contained the proposed kinetic expressions for the reactions, and the parameters of the kinetic model were regressed using our experimental data. Details of the modeling and regression procedure are provided in Hawboldt (1998). Reactor Model. The reactor simulation numerically modeled the experimental apparatus from the inlet through to the quench, solving a set of differential equations to determine the reactor product distribution. The key assumptions in the development of the model are plug flow in the reactor, steady-state operation, and ideal gas behavior. The validity of the plug-flow assumption was previously discussed. Because the apparatus operates at high temperatures and close to atmospheric pressure, the assumption of ideal gas behavior is also valid, given compressibility factors greater than 0.99. The set of differential equations is comprised of the plug-flow reactor design (eq 4), the reaction rate equation (eq 5) in terms of reacting species partial pressures, and an equation for the reactor pressure drop. The

Modeling Methodology The determination of the new rate expression was accomplished by developing a mathematical model that

148

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

Arrhenius law (eq 6) represents the specific rate constant.

dF/dV ) rj

(4)

r ) kΠjPjm

(5)

k ) Ae-Ea/RT

(6)

In addition, an equation of the reactor axial temperature profile was included, which was established using experimental data obtained from a temperature calibration reactor (Hawboldt, 1998). Pure component viscosity was calculated using Chung’s method, and the mixing rules of Wilke were used for the prediction of gas mixture viscosity (Reid et al., 1987). The set of algebraic and differential equations that describe the physical system were solved using an adaptive step-size RungeKutta method taken from Press et al. (1992). Optimization Method. Simulated annealing, detailed in Bohachevsky et al. (1986), was used to minimize the objective function:

OF )

∑|xp,i - xm,i|

(7)

Figure 7. Arrhenius plot for NH3 pyrolysis.

where xp,i is the predicted exiting gas conversion and xm,i is the experimental (measured) gas conversion. The optimization method uses the reactor model to calculate the exiting reactor conversions, which are compared to the experimental conversions. The temperature, pressure, and residence time were set at the beginning of each experiment, and the independent parameters used to minimize the objective function were the Arrhenius constant (A) and activation energy (Ea). Pyrolysis Modeling Results. Reaction Order. Previous studies indicated that NH3 conversions under pyrolysis conditions are small; however, equilibrium calculations show 100% conversion of NH3 to N2 and H2, indicating that the pyrolysis of ammonia (eq 1) is not equilibrium limited at Claus FEF temperatures (Hawboldt, 1998). Hence, the following rate equation is proposed for FEF conditions:

r ) kPNH3m

(8)

The order and initial estimates of A and Ea were made prior to parameter optimization. The reaction order (m) was determined by performing experiments where the temperature and residence time were kept constant while the inlet concentration of NH3 was varied from 0.5 to 2.0%. By plotting inlet versus exit NH3 concentrations, we calculated the value of m to be 1.25. This value was verified using the parameter optimization program. On the basis of the derived form of the rate equation, an Arrhenius plot for NH3 pyrolysis is shown in Figure 7. The slope of the line is -11.0, yielding an Ea of 21.0 kcal mol-1, which was used as an initial estimate for the optimization. Parameter Optimization Results. In the optimization, parameters A, Ea, and m were optimized to match NH3 conversions. The resulting rate equation is shown below as eq 9:

r ) Ae-Ea/RTPNH31.25

(9)

where A is 0.004 21 mol cm-3 s-1 atm-1.25 and Ea is 16.5 kcal mol-1. Figures 3 and 4 show the comparison

Figure 8. Clark et al. (1998) NH3 conversions compared to model predictions. The Clark et al. data is based on a residence time (J) range of 0.6-1.0 s.

between the model predictions and experimental data. The average absolute error (AAE) between the model and data is 13%. Data generated by Clark et al. (1998) was also compared to model predictions, and the results are shown in Figure 8. The experimental conditions of the Clark et al. (1998) data were an inlet NH3 concentration of 14.5% and a residence time range of 0.6-1.0 s. The proposed kinetic model was run at 0.6 and 1.0 s and, as shown in Figure 8, the data falls between these two simulations, further verifying our model. Ammonia Oxidation Modeling. Reaction Order. Equations 2 and 3 show the two primary oxidation pathways for NH3. In this study, the focus has been on eq 2, and consequently, the experiments were run at

Ind. Eng. Chem. Res., Vol. 40, No. 1, 2001 149

Figure 9. Natural logarithm of the regressed rate constant versus inverse temperature for the NH3 oxidation reaction.

conditions where eq 3 was minimized. The proposed form of the rate equation is as in eq 10:

r ) kPNH3mPO2n

(10)

Because direct measurement of O2 is difficult with our apparatus, the reaction orders m and n were determined by the method of excess reactants. That is, to determine m, the O2 was kept in excess of the stoichiometric value (0.75 O2/NH3). Although this ratio exceeds the quantity of O2 required to form NO, as stated previously, negligible amounts of NO are formed. In determining the reaction order with respect to NH3, we varied the NH3 levels from 0.44 to 1.35% at a temperature of 1050 °C and a residence time of 70 ms. The reaction order of NH3 was then calculated by plotting the inlet concentration against the outlet concentration. For the reaction order with respect to O2, n, the NH3 was kept constant at 1.37% and the O2 increased from 2.6 to 4.4% at a residence time of 720 ms and 950 °C. The order with respect to O2 was calculated via eq 11:

(

)

NH3,exit ) n log O2 + constant log ln NH3,inlet

(11)

The calculated values of m and n were 0.98 and 0.72, respectively. These values were set to 1.0 and 0.75 in the oxidation kinetic rate expression. Parameter Optimization Results. Previous studies (Fujii et al., 1981; Clark et al., 1998) have suggested that the temperature dependence of the rate constant changes at approximately 1050-1100 °C, indicating that the Arrhenius constant and activation energy must also change. To verify this result, individual rate constants were regressed at each temperature and plotted as shown in Figure 9. The linear relationship between ln(k) and inverse temperature indicates that there is no significant change in the temperature dependence of the rate constant between 850 °C and 1200 °C, a result that is in disagreement with previous work by Fuji et al. (1981). It is worthwhile to note that in the regression of the rate constants for the oxidation reaction, kinetic expressions for both ammonia pyrolysis and oxidation were included in the simulation, because

Figure 10. Comparison of model predictions with Clark et al. (1998) data for NH3 oxidation (residence time ) 0.34-0.55s).

this more accurately describes the chemistry of the system. As such, the apparent change in the temperature dependence is due to the combined effect of pyrolysis and oxidation. Given the above conclusion that the temperature dependence of the rate constant follows the Arrhenius law, the form of the rate equation is as shown in eq 12:

r ) kPNH3PO20.75

(12)

Figures 5 and 6 compare the model predictions to the experimental data for the complete range of temperatures. The AAE for the data set was 10%, indicating an excellent agreement between the predicted and experimental values. The model was compared with the data generated by Fujii et al. (1981). In this study, the residence time required to consume 5% of NH3 was calculated as a function of temperature. According to the study, for a temperature range of 909-1250 K, the residence time to reach 5% was between 10 and 20 ms. Given their experimental conditions as inputs, our model predicts the same range of residence times, suggesting a good fit. Our oxidation data are compared to the data of Clark et al. (1998) in Figure 10. It is immediately apparent that our measured conversions are much higher than those of Clark et al. (1998). Our data follow an NH3 cracking trend until 800 °C, after which oxidation becomes significant. Note that our cracking results are very similar to those of Clark et al. (1998). According to Clark (personal communication, 2000), their oxidation conversion data are much lower than even cracking until oxidation becomes significant at 1100 °C, because of the inhibition effect of substantial amounts of water present. As such, it can be concluded that we did not experience the inhibition effect. It must be noted that, although the experimental feed concentrations and temperature were similar, the apparatus of Clark et al. (1998) is considerably different than ours with a ceramic reactor and a much slower quench. Thus, the reason for the large difference in conversions could be attributed to the formation of larger amounts of water prior to oxidation in their apparatus. However, this

150

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

needs to be investigated in future work, because it is beyond the scope of this paper. Conclusions In this paper, NH3 pyrolysis and oxidation were studied at Claus FEF temperatures, pressures, and residence times. New experimental data were generated in order to develop global kinetic rate expressions for the pyrolysis and oxidation of NH3 at FEF conditions. At conditions similar to those in a Claus FEF, ammonia oxidation can be much more significant than pyrolysis at temperatures of less than 1150 °C. In fact, conversion of NH3 through oxidation varied between 1 and 50% for temperatures between 850 and 950 °C and 50-700 ms, whereas pyrolysis conversions at these conditions never exceeded 20%. The increase in NH3 conversion due to oxidation is even more significant at temperatures between 1050 and 1200 °C, where the conversion varied from 25 to 90% at the same residence times. For ammonia pyrolysis, the newly developed rate expression is

r ) Ae-Ea/RTPNH31.25 where A is 0.004 21 mol cm-3 s-1 atm-1.25 and Ea is 16.5 kcal mol-1. The AAE between the model and the data was 13%. The model was also used to simulate experiments performed by Clark et al. (1998), and the model predictions agreed well with their data. For ammonia oxidation, the following rate equation was developed:

r ) kPNH3PO20.75 where A and Ea were 4430 mol cm-3 s-1 atm-1.75 and 40.0 kcal mol-1, respectively. Model predictions and experimental data showed good agreement with an AAE of 10%. Although oxygen competition needs to be addressed in a more definitive manner to fully judge industrial recommendations for ammonia destruction, simulations based on our new rate expressions give preliminary indications of what minimum temperatures and residence times are necessary to obtain 60 ppm of ammonia in the furnace effluent, as typically desired for ammonia destruction in the Claus FEF. Specifically, temperatures need to be greater than 1200 °C with residence times greater than 1.0 s or greater than 1250 °C with residence times greater than 0.5 s. In industry, it is generally accepted that the furnace should be designed with residence times greater than 0.8 s, and flame temperatures of 1300 °C or higher are desired. As such, guidelines for actual conditions are somewhat more conservative to account for the incomplete mixing and oxygen competition that occur in Claus plant FEFs. In addition, calculated results based on the new rate expressions clearly show that ammonia pyrolysis accounts for significant destruction. Finally, it should be noted that there appears to be an inhibition of ammonia pyrolysis when significant amounts of water are present and under certain experimental conditions. This phenomenon requires further study to delineate its occurrence.

Acknowledgment The authors thank the following supporters of this work: Shell Canada, British Oxygen Corp., StorkComprimo, HEC Technologies, and NSERC. In particular, the authors thank the Gas Research Institute for their continued support of this research project. In addition, we thank Alberta Sulfur Research for their cooperation and contributions to our continuing studies. List of Symbols A ) Arrhenius constant D ) diameter (cm) Ea ) activation energy (kcal mol-1) k ) reaction rate constant K ) equilibrium constant m and n ) reaction order OF ) objective function P ) pressure (atm) R ) gas constant (kcal mol-1 K-1) r ) reaction rate (mol m-3 s-1) T ) temperature (°C or K) t ) time (s) V ) elemental reactor volume (cm3) xm,I ) experimental fractional conversion xp,I ) predicted fractional conversion z ) axial position (m) Greek Symbols τ ) residence time (s) Subscripts l ) species j ) species Acronyms AAE ) average absolute error SLPM ) standard liters per minute

Literature Cited Bohachevsky, I. O.; Johnson, M. E.; Stein, M. L. Generalized simulated annealing for function optimization. Technometrics 1986, 28 (8), 209-217. Clark, P. D.; Dowling, N. I.; Huang, M. Chemistry of the Claus Front-End Reaction Furnace. Hydrocarbon Reactions and the Formation and Destruction of CS2. Proceedings of the Brimstone Sulphur Recovery Symposium, Vail, CO, Sept 23-26, 1997. Clark, P. D.; Dowling, N. I.; Huang, M. Ammonia Destruction in the Claus Furnace. Proceedings of the Brimstone 1998 Sulphur Recovery Symposium, Vail, CO, Sept 15-18, 1998. Culter, A. H.; Antal, M. J.; Jones, M. A Critical Evaluation of the Plug-Flow Idealization of Tubular-Flow Reactor Data. Ind. Eng. Chem. Res. 1988, 27 (4), 691-697. Davidson, D. F.; Kohse-Hoinghaus, K.; Chang, A. Y. Pyrolysis Mechanism for Ammonia. Int. J. Chem. Kinet. 1990, 22 (5), 513-535. Duo, W.; Dam-Johansen, K.; Østergaard, K. Kinetics of the GasPhase Reaction between Nitric Oxide, Ammonia and Oxygen. Can. J. Chem. Eng. 1992, 70 (5), 1014-1020. Fookes, R. B. Development of a High-Temperature Reactor System to Study the Intrinsic Kinetics of Homogeneous Gas-Phase Sulpur Reactions. M.S. Thesis, Department of Chemical and Petroleum Engineering, The University of Calgary, Calgary, Alberta, Canada, 1996. Fujii, N.; Miyama, H.; Koshi, M.; Asaba, T. Kinetics of Ammonia Oxidation in Shock Waves. Eighteenth Symposium (International) on Combustion, Aug 17-22, 1980; PUBL Combust Inst. 1981; 873-883. Garside, J. E.; Phillips, R. F. Pure and Applied Chemistry; Pittman and Sons Ltd.: London, 1962; pp 751-755. Goar, G. B. Improving Claus Plant Performance by Oxygen Enrichment. Proceedings of Sulphur 94 International Conference, Tampa, FL, Nov 6-9, 1994; Vol. 1, pp 109-129.

Ind. Eng. Chem. Res., Vol. 40, No. 1, 2001 151 Grancher, P. Advances in Claus Technology. Hydrocarbon Process. 1978, 57 (7), 155-160. Hawboldt, K. A. Kinetic Modeling of Key Reactions in the Modified Claus Plant Front End Furnace. Ph.D. Thesis, Department of Chemical and Petroleum Engineering, The University of Calgary, Calgary, Alberta, Canada, 1998. Hawboldt, K. A.; Monnery, W. D.; Svrcek, W. Y. A Study on the Effect of Quench Design on the Quality of Experimental Data. Ind. Eng. Chem. Res. 1999a, 38 (6), 2260-2263. Hawboldt, K. A.; Monnery, W. D.; Svrcek, W. Y. New Experimental Data and Kinetic Rate Expression for H2S Cracking and ReAssociation. Chem. Eng. Sci. 1999b, 55 (5), 957-966. Johnson, J. E.; Rempe, M. L. Reaction Furnace Design and Operation for Processing Ammonia Bearing GasessA Review of Fundamentals and a Case Study. Proceedings of the Brimstone 1997 Sulphur Recovery Symposium, Vail, CO, Sept 2326, 1997. Lindstedt, R. P.; Selim, M. A. Reduced Reaction Mechanism for Ammonia Oxidation in Premixed Laminar Flames. Combust. Sci. Technol. 1994, 99 (4-6), 277-298. Lindstedt, R. P.; Lockwood, F. C.; Selim, M. A. Detailed Kinetic Modeling of the Chemistry and Temperature Effects on Ammonia Oxidation. Combust. Sci. Technol. 1994, 99 (4-6), 253276. Lyon, R. K.; Benn, D. Kinetics of the NO-NH3-O2 Reaction. 17th Symposium (International) on Combustion, Leeds, England, Aug 20-25, 1978; pp 601-610. Maclean, D. L.; Wagner, H. Gg. The Microstructure of the Reaction Zones of Ammonia-Oxygen and Hydrazine Decomposition Flames. 11th Symposium (International) on Combustion, Berkeley, CA, Aug 14-20, 1967; pp 871-878.

Miller, J. A.; Bowman, C. T. Mechanism and Modelling of Nitrogen Chemistry in Combustion. Prog. Energy Combust. Sci. 1989, 15 (4), 287-338. Miller, J. A.; Smooke, M. D.; Green, R. M.; Kee, R. J. Kinetic Modelling of the Oxidation of Ammonia in Flames. Combust. Sci. Technol. 1983, 34, 149-176. Monnery, W. D.; Svrcek, W. Y.; Behie, L. A. Modelling the Modified Claus Reaction Furnace and the Implications on Plant Design and Recovery. Can. J. Chem. Eng. 1993, 71 (5), 711-724. Nauman, E. B. The Residence Time Distribution for Laminar Flow in Helically Coiled Tubes. Chem. Eng. Sci. 1977, 32 (3), 287294. Press, W. H.; Vetterling, W. T.; Teukolsy, S. A.; Flannery, B. P. Numerical Recipes in CsThe Art of Scientific Computing, 2nd ed.; Cambridge University Press: Cambridge, U.K., 1992. Reid, R. C.; Prausnitz, J. M.; Poling, B. E. The Properties of Gases and Liquids, 4th ed.; McGraw-Hill: New York, 1987. Truesdell, L. C.; Adler, R. J. Numerical Treatment of Fully Developed Laminar Flow in Helically Coilded Tubes. AIChE J. 1970, 16 (6), 1010-1015. Vandooren, J. Comparison of the Experimental Structure of Ammonia Seeded Rich-Hydrogen-Oxygen-Argon Flame with Calculated Ones Along Several Reaction Mechanisms. Combust. Sci. Technol. 1992, 84 (1-6), 335-344.

Received for review October 22, 1999 Revised manuscript received September 27, 2000 Accepted September 27, 2000 IE990764R