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Numerical Study on Coal Gasification in the Downer Reactor of a Triple-Bed Combined Circulating Fluidized Bed Yongpan Cheng†,‡ and Chi-Hwa Wang*,† †

Department of Chemical and Biomolecular Engineering, National University of Singapore, 4 Engineering Drive 4, 117585, Singapore NUS Environmental Research Institute, National University of Singapore, 5A Engineering Drive 1, #02-01, 117411, Singapore

‡

ABSTRACT: Owing to the advantages of cocurrent gas−solids flow without back mixing, short residence time, and uniform residence time distribution, the downer has emerged as an ideal reactor for gasification with high selectivity. In this study, the coal gasification in the down reactor of a triple-bed combined circulating fluidized bed was studied through computation fluid dynamics simulations using an Eulerian−Lagrangian method. The influences of nozzle arrangement for coal feeding, coal particle, and air/steam/coal feeding rates on the gasification were investigated. It was found that the tangential arrangement of feeding resulted in comparable H2 production with the normal arrangement, and higher CO production. When coal particle size was smaller than 500 μm, the particle size had little influence on the produced gas composition. In contrast, with increasing particle size beyond 500 μm, coal particles could not be gasified completely due to shorter residence time, leading to decreasing production of CO and H2. Steam gasification had a higher volume fraction of CO and H2, as well as higher char conversion ratio. With increasing coal feeding rates, the volume fraction of CO increased monotonically, while that of H2 increased first and then approached a constant value due to limited moisture availability in the coal samples. With increasing air feeding rates, more char and volatiles could be decomposed into light gases. As a result, the volume fraction of CO increased first and then started to decrease. These results had great significance in designing a better downer reactor with improved efficiency. gasification was provided. Chen et al.5,6 performed a simulation on a two-hundred ton per day two-stage air blown entrained flow gasifier developed for the IGCC process with the Multi Solids Progress Variables (MSPV) method. The predicted gas temperature profile and the exit gas composition were in general agreement with the measured values, and the effects of various operating conditions, such as heterogeneous reaction rate, coal type, particle size, and air/coal partitioning to the two stages, were studied in the gasification process. Choi et al.7 studied the coal gasification process in a slurry feed type, an entrained flow coal gasifier where the whole process is divided into several simplified stages, such as slurry evaporation, coal devolatilization, and two-phase reaction coupled with turbulent flow and twophase heat transfer. Because of the total heat loss through the gasifier and uncertain kinetics for the heterogeneous reactions, the measured quality of the syngas was better than the calculated one when the O2/coal ratio was increased, while the carbon conversion agreed well. Watanabe and Otaka8 simulated the gasification for an entrained flow coal gasifier with the capacity of two tons per day, comprising integrated studies of pyrolysis model, char gasification model, and gas phase reaction model. The influence of air ratio on gasification performance, such as the carbon conversion efficiency, amount of product char, heating value of the product gas, and coal gas efficiency were studied in detail together with the validation of the experimental data. Vicente et al.9 investigated the gasification in a commercial entrained flow coal gasifier, with the following coal particle

1. INTRODUCTION As a process to reduce the emission of CO2 and alleviate global warming, coal gasification has been the focus of great attention in recent research activities worldwide. The gasification can convert various carbon-based feed stocks to clean synthetic gas, mainly a mixture of hydrogen and carbon-monoxide. Subsequently, the synthetic gases can be further used to produce other chemicals or act as fuel to produce the electricity through the integrated gasification combined cycle (IGCC). In this cycle, the gas turbine and/or steam turbine is integrated with coal gasification, the produced synthetic gases are burned with the compressed air in the gas turbine to produce high pressure and hot gases for driving the generator to produce electricity. The hot exhaust gases from the gas turbine are sent to the steam generator for heat recovery to produce steam, which is used to drive another generator to produce electricity. The waste heat from the exhaust gases from the steam turbine can be recovered to the coal gasification process. With the heat recuperation, the efficiency of IGCC can be increased to around 52% with low pollutant emissions, compared to 43−45% efficiency and higher emissions in the traditional coal combustion power plants.1 Coal gasification process has been well studied in the literature,2−4 and there are three basic reactor types, namely fixed bed, fluidized bed and the entrained flow bed. Because of its high coal throughput, insensitivity to coal type and its simplicity in design and technology, the entrained flow bed has been chosen as a promising way for coal gasification. With the emergence of high performance computing, a numerical simulation has been widely adopted to simulate the coal gasification process. Silaen and Wang1 numerically studied the oxygen-blown coal gasification process inside a generic entrained flow gasifier; the effect of turbulence and devolatilization models on the coal © 2014 American Chemical Society

Received: Revised: Accepted: Published: 6624

January 2, 2014 March 27, 2014 April 2, 2014 April 2, 2014 dx.doi.org/10.1021/ie500013y | Ind. Eng. Chem. Res. 2014, 53, 6624−6635

Industrial & Engineering Chemistry Research

Article

knowledge, there is limited availability of related studies in this specific area. This has been the motivation for the present study. In this manuscript, the statement for the physical problem will be given first, followed by the description of the mathematical method. On the basis of the numerical results presented, some elaborated analysis will be performed. Finally recommendations and conclusions will be provided for performing coal gasification in the downer rector.

processes/associated closures simulated: drying, volatilization, heterogeneous reactions of combustion and gasification, particle drag, turbulent dispersion, as well as heating-up. Slezaket al.10 simulated the effects of coal particle density and size fraction on the coal gasification in an entrained flow gasifier. Kumar and Ghoniem11 described the construction, validation, and application of a multiscale model of entrained flow gasification; the accuracy of the integrated gasifier model was demonstrated by comparing its prediction with experimental data from pilot and research-scale entrained flow gasifiers. Although the entrained flow gasifier is widely applied in industry, its cold gas efficiency and energy efficiency are quite low. Furthermore, it requires a large amount of oxidant in order to create entrainment. On the other hand, the fluidized bed gasifier started to gain more attention from the researchers due to the high heat transfer efficiency inside the bed. By using the computational particle fluid dynamics (CPFD) methodology based on the multiphase particle-in-cell model (MP-PIC), Snider et al.12 studied the energy transport and chemistry inside the dense fluidized bed gasifier where the interdependencies of fluidization, thermal, and chemistry behavior were described in detail. Also with the CPFD methodology Wang et al.13 and Xie et al.14 developed a comprehensive three-dimensional model to simulate the coal gasification in a fluidized bed gasifier. The flow patterns, gas velocities, particle velocities, composition profiles of gas product, and distributions of reaction rates were provided from the simulation results, and the exit gas compositions were found to be in good agreement with the experiments. Grabneret al.15 presented the modeling results for the pressurized fluidized bed gasifier of 4800 tons per day at a pressure of 33 bar. The formation of flow pattern, turbulence, product gas composition, temperature, and radiation heat transfer were investigated, and the influence of diameter variation on the flow patterns at constant operating conditions was presented. Chejne et al.16 also modeled the coal gasification process in a pressurized fluidized bed in which the temperature, converted fraction, and particle size distribution for solids were predicted, and gasification performance was optimized by changing the excess amount of air, particle size distribution, coal type, and reactor geometry. In both entrained flow gasifier and fluidized bed gasifier, the produced tar during the devolatilization process of coal gasification may adhere onto the coal particle surface, which severely hinders contact between the solid coal particles and reactant gases around them. Henceforth, the efficiency of coal gasification may be reduced greatly. As a novel type of reactor, the downer reactor is ideal for the devolatilization process. In the downer, the gas phase and solid particulate phase flow concurrently in the direction of gravity and undergo the pyrolysis process. Due to the aid of gravitational acceleration, very short residence time and narrow distribution of residence time can be realized in the downer. Based on this idea, the Triple-Bed combined Circulating Fluidized Bed (TBCFB) was proposed, which integrated a downer for coal pyrolysis, a bubbling fluidized bed for gasification of char, and a riser for partial oxidation of unreacted char.17−22 The heat generated from the combustion can be transferred by the circulation of inert solids particles back to the downer and bubbling fluidized bed for pyrolysis and gasification. Fushimi et al.23 and Cheng et al.24 studied the mixing characteristics between injected coal particles and circulating inert heat-carring silica sand with different nozzle arrangements above the downer reactor. On the basis of those studies, it was necessary to conduct further comprehensive studies on the coal gasification in the downer reactor. To the best of the authors’

2. MODEL DESCRIPTION The schematic diagram of the downer reactor is shown in Figure 1, which was part of the triple-bed combined circulating fluidized

Figure 1. Schematic diagram of the downer in a triple-bed combined circulating fluidized bed.

bed. Gasifying agents, such as air or steam, were injected from the top of the downer. In contrast, the cold coal particles were injected from the two nozzles mounted near the top of downer, then gasification was started after heat was absorbed from the downer. Finally the produced synthetic gases, CO2, volatiles, tar, or other hydrocarbon gases flowed out through the bottom of the downer, and the unreacted char went to the bubbling fluidized bed for further gasification. The nozzles were arranged normally or tangentially to the downer. During the gasification process in the downer, the flow involved both gas and solid phases. The chemical reactions of gasification involved homogeneous reactions among gases as well as heterogeneous reactions between gases and solids. Hence, it was a multiphase reactive flow problem solved by an Eulerian− Lagrangian method in the present study. The Eulerian method was used to model the gas continuous phase, as the solid fraction in the downer was quite low, there were few interactions among coal particles, and the Lagrangian method was used to solve the solid discrete phase. The gas phase calculation was first implemented, followed by the calculation for the solid phase. Subsequently, the paths of the particles were determined by the force balance on the particles calculated by the flow field, and the heterogeneous reactions were calculated. After completing the calculation of the solid phase, the gas phase was updated with the mass and energy change caused by the solid phase, and then the next iteration of gas and solid phase interaction started. 2.1. Continuous Phase. In the Eulerian−Lagrangian method, the Eulerian method was adopted to simulate the 6625

dx.doi.org/10.1021/ie500013y | Ind. Eng. Chem. Res. 2014, 53, 6624−6635


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where Rep = ρdp|u − up|/μ, CD = a1 + a2/Rep + a3/Rep2; a1, a2, and a3 are constants that apply over several ranges of Re given by Morsiand Alexander.25 The temperature of the coal particles affected the gasification performance of coal particles in the downer, and the temperature variation was determined by the convective heat transfer between coal particles and surrounding gases, and the reactive heat and thermal radiation, as shown below.

continuous (gas) phase. The governing equations for the mass, momentum, and energy conservation were shown below: ∂ (ρui) = Sp,m ∂xi

(1)

∂p ∂ ∂ (ρuiuj) = ρgj − + (τij − ρui′u′j) + Sp,mom ∂xi ∂xj ∂xi

(2)

∂ ∂ ⎡ ∂T ⎤ (ρuih) = ⎢λ ⎥ + Srad + S h + Sp,h ∂xi ∂xi ⎣ ∂xi ⎦

mpCp (3)

∂xj

=

QG =

∂xj

=

+

∂uj ⎞ ∂ui ∂ ⎡⎢ ∂ε ⎤⎥ ε ⎛ ∂u ⎟ + Cε1 ⎜⎜ i + ραεμeff k ⎝ ∂xj ∂xi ⎟⎠ ∂xj ∂xj ⎢⎣ ∂xj ⎥⎦ ε2 −R k

(6)

2.3. Discrete Phase. Because of the heterogeneous reactions, devolatilization, and moisture vaporization, the mass of coal particles would be changed, which was calculated as dmp

=

dt

dmC ‐ O2 +

+

dmC ‐ CO2

dt dm vapor

dt

+

dmC ‐ H2O dt

dt

= FD(u − u p) +

gi(ρp − ρ) ρp

18μ C DRep ρp d p2 24

dt dmC ‐ H2O dt

dmC ‐ CO2 dt

HC ‐ CO2

HC ‐ H2O

(11)

(12)

volatiles → 0.299tar + 0.107CO + 0.078CO2 + 0.028H 2 + 0.359H 2O + 0.089N2

+ Fi

(R1)

(8)

tar → 0.645C(s) + 0.032CO + 0.32CH4

where the three terms on the right-hand side represented the drag force of the fluid on the particle, the gravity, and the other body forces including thermophoretic force, Brownian force, Saffman’s lift force, etc. The drag force was calculated according to the spherical drag law as FD =

HC ‐ O2 +

where A = 4.92 × 105 and E = 7.4 × 107 J/kg mol. After the volatiles were released from the coal particle, they were decomposed into various light gases (such as CO, H2, and CO2, etc.) through thermal cracking. The composition of the volatiles could be determined according to the proximate and elemental analysis of the coal. According to Syamlal and Bissett,28 the following volatiles decomposition and tar cracking reactions were assumed in the study.

A Lagrangian method was adopted to track the motion of particles in this study. Since the particles in the flow field experienced external forces, the velocity was accelerated or decelerated. The change of velocity was determined by the force balance shown as du p

(G − 4σTp 4)

k = Ae−E /(RT )

dmdevol + dt (7)

dt

4

The thermal radiation was calculated with P-1 model, which was based on the expansion of radiation intensity into an orthogonal series of spherical harmonics.26 2.4. Chemical Reactions. Devolatilization was the process in which the volatile gases were released from the coal particles. Here the devolatilization model could be separated into two steps: (i) release of the volatile gas simulated by the devolatilization model, (ii) decomposition of the released volatiles into various light gases by thermal cracking. The volatile gas was a lump-sum gas representing all the volatile matters, and the release reaction rate was calculated by the single rate devolatilization model.27 In this model, the rate of devolatilization was determined by the amount of volatiles remaining in the coal particle. The kinetic rate of the decomposition reaction of the coal into volatiles and char was defined according to the Arrhenius law (coal → volatiles + char, with zeroth order rate constant k):

⎛ ∂u ∂uj ⎞ ∂ui ∂ ⎡⎢ ∂k ⎤⎥ ⎟ − ρε + ρμt ⎜⎜ i + ραkμeff ∂xi ⎟⎠ ∂xj ∂xj ⎢⎣ ∂xj ⎥⎦ ⎝ ∂xj

− Cε2ρ

dmC ‐ O2

(4)

(5)

∂(ρujε)

εpA p

The reactive heat was calculated by three heterogeneous reactions between carbon and either oxygen, CO2, or steam.

2.2. Turbulence Models. The standard k−ε model was the most commonly used turbulent models with the advantages of robustness, economy, and reasonable accuracy for a wide range of turbulent flows; hence, it was used in the present study, with the governing equations shown as follows: ∂(ρujk)

dt

= hpA p(T − Tp) + Q G +

(10)

The “species transport model” was adopted to calculate the mixing and transport of different chemical species. The equations of the species transport model was given as ∂ ∂ ⎡ ∂Yi ⎤ (ρuiYi ) = ⎢ρD ⎥ + SYi + SP , Yi ∂xi ∂xi ⎣ ∂xi ⎦

dTp

(R2)

The reactions rates for these two reactions were as follows: ⎧ ⎛ E ⎞ ⎪ A exp⎜ − a ⎟εsρ (XVM − X *) XVM ≥ X * ⎝ RTs ⎠ s r1 = ⎨ ⎪ ⎪0 XVM < X * ⎩

(13)

where X* = (X0VM + X0FC)Xmin *

(9) 6626


Industrial & Engineering Chemistry Research ⎧ 867.2 ⎪ ⎪ Ts − 273 * =⎨ X min 100 ⎪ ⎪0 ⎩

(

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Table 1. Arrhenius Chemical Reaction Rates in Our Simulation11

3.914

)

Ts < 1223 Ts ≥ 1223

⎛ −29000 ⎞ r2 = 2.5 × 10 exp⎜ − ⎟εgρ X tar RTs ⎠ g ⎝ 7

CO + H 2O ↔ H 2 + CO2

(R3)

1 CO + O2 → CO2 2

(R4)

1 O2 → H 2O 2

(R5)

(R6)

⎛ 9.23 × 107 ⎞ 0.4 r6 = 2.3T exp⎜− ⎟[O2 ] RT ⎠ ⎝

(R7)

⎛ 1.62 × 108 ⎞ r7 = 4.4T exp⎜− ⎟[CO2 ]0.6 RT ⎠ ⎝

(R8)

⎛ 1.47 × 108 ⎞ r8 = 1.33T exp⎜− ⎟[H 2O]0.6 RT ⎠ ⎝

fixed carbon volatile matter ash moisture HHV

(R6) (R7)

C + H 2O → CO + H 2

(R8)

C H O N

75.3 5.4 15.6 3.7

3. RESULTS AND DISCUSSION 3.1. Model Validation. To validate the model used in our simulation, a preliminary numerical simulation based on the experimental setup by Ocampo31 was conducted. The feeding rates of air, steam, and coal as well as bed temperature in the simulation were kept the same, and the experimental steam gasification and the produced gas composition at the outlet of gasifier were compared as seen in Figure 2. It was found that the production of CO, H2, and CO2 in the simulation was a little

(15)

C + CO2 → 2CO

54.1 41.8 1.5 2.6 2.97 × 107 J/kg

ultimate analysis (wt %)

phase coupled SIMPLE algorithm was used for pressure and velocity coupling. The second-order upwind scheme was adopted to discretize the governing equations, and a residual of less than 1.0−4 for all the variables was imposed as the stopping criterion.

where the equilibrium constant Kp(T) = 0.0265 exp(4.546 × 107/(RT)). The heterogeneous reactions between char and gases occur (O2, CO2 and H2O) as follows: 1 O2 → CO 2

⎛ 1.67 × 108 ⎞ −1 r5 = 6.8 × 1015 exp⎜− ⎟T [H 2]0.25 [O2 ]1.5 RT ⎠ ⎝

proximate analysis (wt %)

⎛ 1.26 × 107 ⎞ r3 = 2.78 × 106 exp⎜ − ⎟ RT ⎝ ⎠

C+

(R5)

Table 2. Properties for Colombian Coal31

The water gas shift reaction was modeled as a chemical equilibrium reaction that favored the forward reaction at relatively low temperature for the production of more H2 and CO2, and reverse reaction at high temperature for the production of more CO and H2O. The reaction rate was as follows30

⎛ [CO2 ][H 2] ⎞ ⎟ × ⎜⎜[CO][H 2O] − K p(T ) ⎟⎠ ⎝

⎛ 1.67 × 108 ⎞ r4 = 2.24 × 1012 exp⎜− ⎟[CO][O2 ]0.25 [H 2O]0.5 RT ⎝ ⎠

(14)

Since the flow of gas phase was turbulent, homogeneous gas phase reactions were controlled by the reaction with lower Arrhenius rate and eddy dissipation rate.29 In this model, three homogeneous gas phase reactions were introduced, including the water−gas shift reaction and exothermic oxidation reaction of CO and H2, as follows:

H2 +

(R4)

The kinetics from R4 to R8 were assumed as an Arrhenius rate,11 as shown in Table 1. 2.5. Simulation Conditions. In this study the Colombian coal was used, and its properties are shown in Table 2. The coal particles at ambient temperature were injected from the nozzles near the top of the downer. Air or steam was fed into the downer as a gasifying agent. The feeding rates of coal, steam and air, particle size, and nozzle arrangement could be varied, so as to study the corresponding gasification performance of coal inside the downer. To obtain the accurate solutions for the gasification inside the downer, fine mesh was used near the nozzle region. After the grid-independent study, the grid number around 12000 was used throughout the simulation. A constant pressure boundary condition was applied at the exit of the downer. The wall was assumed to be no-slip and at constant temperature around 850 °C. The commercially available Computational Fluid Dynamics software FLUENT 14.0 (ANSYS, Inc., USA) was adopted. The

Figure 2. Comparison of outlet gas compositions between numerical simulation and experiment31 at ṁ coal = 8 kg/h, ṁ air = 21.9 kg/h, ṁ steam = 4.7 kg/h, and Tair−steam = 420 °C. This comparison serves as the validation of the present simulation code. RD stands for relative error. 6627



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Figure 3. Distribution of species volume fraction in the upper quarter of the downer for the tangential arrangement at dp = 200 μm, ṁ steam = 4.7 kg/h, ṁ air = 21.7 kg/h and ṁ coal = 8 kg/h.

upper quarter and at z = 0.05 m are shown in Figure 3 and Figure 4, respectively, under the following experimental condition with the tangential nozzle arrangement: dp = 200 μm, ṁ steam = 4.7 kg/ h, ṁ air = 21.7 kg/h, ṁ coal = 8 kg/h. When the coal particles at ambient temperature were injected into the downer, they absorbed heat quickly and started to devolatilize, then they

underpredicted, while the production of CH4 was overpredicted, but the maximum deviation was 23.1%, so the agreement between numerical simulation and experimental results was satisfactory. 3.2. Species Distribution. To perform detailed study inside the downer, the distributions of various gas components at the 6628



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Figure 4. Distribution of species volume fraction at z = 0.05 m of the downer for the tangential arrangement at dp = 200 μm, ṁ steam = 4.7 kg/h, ṁ air = 21.7 kg/h, and ṁ coal = 8 kg/h.

fraction of CO was quite low. Subsequently, the CO volume fraction increased along the downer when O2 was consumed. The variation trend of CO was opposite to that of CO2 because of the water gas shift reaction. In Figure 5, the axial distributions of different species were plotted along the axial location of the downer. The consistent results could be observed when coal particle size was maintained at dp = 200 μm. It could be found that the volume fractions of O2 and H2O decreased sharply, and then approached constant values quickly at z = 0.5 m. Under steady state with steam as the gasifying agent, the volume fraction

were decomposed into volatiles and char, which continued to react with other gases. Because of the strong swirl flow in the downer for the tangential arrangement of nozzles, the volume fractions of steam and O2 in the central regions were quite high, and these components were consumed quickly in the downer near the nozzles. Because the coal particles were quite small, they could easily react with the gasifying agent. CH4 was mainly produced during tar cracking, so its volume fraction became quite stable after the nozzles. As the gas was rich in O2 near the nozzles in the downer, there was little production of CO, and the volume 6629



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Figure 5. Axial distribution of gas volume fractions for steam gasification in the tangential arrangement under different coal particle sizes.

particles as well as the corresponding gasification performance. Figure 7 shows the effect of particle size on the produced gas composition at the outlet of the downer for the steam gasification case under two different nozzle arrangements. As expected, when the coal particle size was smaller than 500 μm, the particle size had little effect on the gas composition at the outlet. This was because, for a small particle size, the reaction rate between the solid coal particles and surrounding gases was high enough to convert the particles into gases completely. In contrast, for large particles, the conversion would become incomplete due to shorter residence time, thus the consumption of gasifying agent, such as O2 and steam, became less, leading to less production of CO and H2. For example, when the particle size was as large as 1 mm, the volume fraction at the outlet was only around 1%, so particle size played a critical role in the gasification in the downer. From Figure 8a to 8e, the produced gas compositions at the outlet of downer were compared under different gasifying agents and nozzle arrangements. It was found that, with increasing coal particle size, the volume fractions of O2 were constant when particle size was smaller than 500 μm, then they increased gradually due to lower consumptions of O2. Similar trends were observed for the variation of steam. The volume fractions of H2 were comparable for the tangential and normal nozzle arrangements regardless of gasifying agent, and the volume fractions of H2 and CO in steam gasification were higher than those in air gasification. The char conversion ratio was crucial for the gasification reaction to proceed in the downer. From Figure 8f it could be found that the char conversion ratio in the steam gasification case was much higher than that in the air gasification case, which might account for a high production of CO and H2 in the steam gasification. For steam gasification, the char conversion ratio was

of H2 was the highest, followed by CO, CO2, O2, CH4, and H2O. The axial distributions of the gas species were also provided in Figure 5 where the coal particle size was increased to 500, 800, and 1000 μm. It was found that with increasing coal particle size, the residence time of coal particles became shorter. Henceforth, the conversion of the coal particles could not complete, and the consumption of O2 and steam became less, leading to lower production rate of CO and H2. When the particle size was larger than 500 μm, the gas species could not reach a stable state, which meant that the length of downer was not long enough for complete conversion of coal particles. These findings provided an example guideline for design of the downer under fixed operating conditions. 3.3. Effect of Nozzle Arrangement. Figure 6 showed the axial distributions of gas volume fractions at dp = 500 μm when the nozzles were arranged either normally or tangentially to the downer. It could be found that the variations of H2 were quite similar in two arrangements: along the downer the volume fraction increased gradually and reached stable values in the downstream, and the volume fractions at the outlet were both around 20%. However, the variation trends of CO were quite different. For instance, in the tangential arrangement, CO volume fraction increased monotonically until it was 17% at the outlet. But in the normal arrangement, the CO volume fraction increased first, but after reaching a plateau at about 10% at z = 3.2 m, it started to decline, and the volume fraction at the outlet was only about 8%. Because of the presence of the water gas shift reaction, the volume fraction of CO2 at the outlet became quite high. 3.4. Influence of Coal Particle Size. In the downer, the coal particles flow downward under the drag force and gravity, so the particle size has a great influence on the residence time of coal 6630



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Figure 6. Axial distribution of gas volume fractions at dp = 500 μm under two different nozzle arrangements. Figure 7. Effect of particle size on produced gas composition.

independent of the nozzle arrangement, while for air gasification, the char conversion ratio in the normal arrangement was higher than that in the tangential arrangement. The lower heating value of produced gas, LHVgas, could be calculated based on the following formula:32 LHVgas = 10.8yH + 12.6yCO + 35.8yCH [MJ/m 3] 2

4

CO2 was reduced, because there was less CO for CO oxidation reaction R4 to occur with the consumption of CO, leading to low CO2 volume fraction, whereas at high steam flow rate, reaction R7 might become dominant in the conversion of CO2 into CO, so the volume fraction of CO2 was reduced. 3.6. Influence of Coal Feeding Rates. Figure 10 showed the influence of coal feeding rates on the produced gas composition under the case of air gasification where the air flow rate ṁ air was 21.7 kg/h and coal particle size dp was 200 μm. As expected, with increasing coal feeding rates, the volume fraction of CO in the produced gases increased because more carbon was available for the reaction of carbon with O2, CO2, and H2O R6−R8, either one could contribute to the production of CO. Moisture in the coal would be vaporized to steam, and further reacted with CO to produce H2 and CO2. Henceforth, the volume fraction of steam decreased greatly at low coal feeding rate with increasing feeding rate. In contrast, the steam volume fraction was kept at low values at high coal feeding rates. Correspondingly, the volume fraction of H2 increased greatly first and then approached almost a constant. This was mainly determined by the water gas shift reaction; that is, when steam was consumed completely, no more H2 could be produced. Because of the oxidation reaction R4 of CO to produce CO2, the volume fraction of CO2 increased first at low coal feeding rates, then started decreasing at high coal feeding rates. This was because more carbon was available for reaction R7, and more CO2 was consumed to produce CO.

(16)

where yH2, yCO, and yCH4 are the mole fractions of hydrogen, carbon monoxide, and methane in the produced gas, respectively. In Figure 8g the lower heating values for the normal and tangential arrangements of nozzles were compared for air and steam gasification. It could be found that when coal particle size was small in the steam gasification case, the lower heating values in the tangential arrangement were higher than those in the normal arrangement, while for air gasification the trend was opposite. 3.5. Influence of Steam Flow Rates. Steam has great influence on the gasification performance in the downer. In Figure 9, the influences of steam flow rates on the produced gas compositions were provided for a specific case study where the air flow rate ṁ air was 21.7 kg/h and the coal particle size dp was 200 μm. It was found that with increasing steam flow rates, the volume fraction of CO in the outlet-produced gases decreased while that of H2 increased. Obviously, this was due to the water gas shift reaction R3. With increasing steam feeding rate, the water gas shift reaction would move forward, so CO was consumed, accompanied by the production of H2 and CO2. It was noted that under low steam flow rate, the volume fraction of 6631



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Figure 8. Comparison of produced gas compositions between air/steam gasification for normal/tangential arrangements.

3.7. Influence of Air Flow Rates. Figure 11 showed the influence of air flow rates on the produced gas composition at the

outlet of the downer, as well as the char conversion ratio for air gasification at the fixed coal feeding rate ṁ coal = 8 kg/h and coal 6632



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volatiles was reduced and the char conversion ratio increased. It was noted that the volume fraction of CO increased with increasing air flow rates. This might occur because more volatiles were converted into CO and other light gases, then the volume fraction of CO was reduced, maybe because more O2 was available to oxidize CO.

4. CONCLUSIONS In this study the coal gasification through air and steam gasification was investigated through numerical simulation in the downer reactor of a triple-bed combined circulating fluidized bed. After validation with reliable experimental data, the influences of different nozzle arrangements, coal particle size, and air/steam/coal feeding rates on the gasification performance were investigated to optimize the process. It was found that the tangential arrangement had equivalent H2 volume fraction at the outlet of the downer with the normal arrangement, and higher CO volume fraction. When the coal particle size was smaller than 500 μm, the particle size has little effect on the produced gas composition, whereas with increasing particle size beyond 500 μm, the char conversion was not complete due to short residence time. The volume fractions of CO and H2 were reduced greatly, while those of O2 and steam increased greatly. the steam gasification had higher volume fraction of CO and H2, as well as char conversion ratio. With increasing coal feeding rates, the volume fraction of CO increased monotonically while that of H2 increased first and then approached a plateau due to the availability of moisture in the coal. With increasing air feeding rates, more char and volatiles could be decomposed into light gases. As a result, the observed volume fraction of CO increased first and then started declining. These results may be useful in the design of a high performance downer reactor for coal gasification.

Figure 9. Effect of steam mass flow rates on produced gas compositions for the tangential arrangement at ṁ air = 21.7 kg/h and dp = 200 μm.

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AUTHOR INFORMATION

Corresponding Author

*Tel.: +6565165079. Fax: +6567791936. E-mail: chewch@nus. edu.sg.

Figure 10. Effect of coal mass flow rates on producer gas compositions at ṁ air = 21.7 kg/h and dp = 200 μm.

Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The initial phase of this study was funded by the Economic Development Board (EDB) of Singapore under the Grant No. R261-501-003-414 through the Minerals, Metals, and Materials Technology Center (M3TC), National University of Singapore (NUS). The authors acknowledge the advices offered by Professor Atsushi Tsutsumi (The University of Tokyo) and technical support by Zhenyuan Yin, Wen Wu, Yang Weng, Jayadev S Marol, and Wei Gao. At the later stage of study, this research programme was funded by the National Research Foundation (NRF), Prime Minister’s Office, Singapore under its Campus for Research Excellence and Technological Enterprise (CREATE) programme.

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Figure 11. Effect of air mass flow rates on produced gas compositions in the tangential arrangement at ṁ coal = 8 kg/h and dp = 200 μm.

particle size dp = 200 μm. It was noted that the coal pyrolysis occurred when there was no air feeding into the downer; coal particles were decomposed into char, tar, volatiles, and other gases. With increasing air feeding rates, char and volatiles could be further converted into light gases, so the volume fraction of 6633

NOMENCLATURE A = pre-exponential factor for chemical reactions; particle external surface area, m2 C1ε,C2ε = constants in turbulent model CD = drag coefficient Cp = specific heat, J/(kg·K) d = diameter, m Ea = activation energy for chemical reaction, J dx.doi.org/10.1021/ie500013y | Ind. Eng. Chem. Res. 2014, 53, 6624−6635


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FD = drag force, N g = gravitational constant, m/s2 G = net irradiation at the location of particle, kg/s3 h = enthalpy, J/kg k = turbulent kinetic energy, m2/s2 ṁ = turbulent kinetic energy, kg/s p = pressure, Pa QG = heat source in particle energy equation, J/s r = chemical reaction rate R = universal gas constant, J/(mol·K) Re = Reynolds number Sh = source term due to homogeneous reactions in gas phase energy equation, J/s Srad = radiation source term in gas phase energy equation, J/s Sp,m,Sp,mom,Sp,h,SP,Yi = interphase exchange terms for mass, momentum, enthalpy and species SY = source term in species equation due to gas phase reaction, kg/(m3·s) T = temperature, K u = velocity, m/s x = coordinate, m X = mass fraction y = coordinate, m; mole fraction Y = mass fraction of the species z = coordinate, m αk,αε = constants in turbulent model ε = turbulent energy dissipation rate, m2/s3 particle emissivity σ = Stefan−Boltzmann constant νeff = effective viscosity, kg/(m·s) νt = turbulent viscosity, kg/(m·s) ρ = density, kg/m3 τ = shear stress, N/m2 FC = fixed carbon p = particle VM = volatile matters

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