Numerical Simulation of Industrial Opposed Multiburner CoalâWater

Article pubs.acs.org/IECR

Numerical Simulation of Industrial Opposed Multiburner Coal−Water Slurry Entrained Flow Gasifier Zhonghua Sun, Zhenghua Dai, Zhijie Zhou, Qinghua Guo, and Guangsuo Yu* Key Laboratory of Coal Gasification of Ministry of Education, East China University of Science and Technology, Shanghai 200237, People’s Republic of China ABSTRACT: A comprehensive three-dimensional numerical model is developed for simulation of the industrial opposed multiburner (OMB) coal−water slurry (CWS) entrained flow gasifier. The CWS gasification process is divided into several submodels, including water evaporation, coal pyrolysis, and homogeneous and heterogeneous char reactions. The “effectiveness factor” method is used to extrapolate the intrinsic char reactivity data to industrial CWS gasifier conditions. Numerical simulations with the proposed models are performed on the industrial OMB gasifier, and the predicted temperature and gas compositions at the outlet of the gasifier are in good agreement with industrial operating data. The profiles of flow field, temperature, and gas composition in the gasifier are described. In addition, the effects of the slurry concentration and oxygen/ coal ratio on the gasifier performance are also studied, which reveal that the lower oxygen/coal ratio or higher slurry concentration leads to higher cold gas efficiency, while the carbon conversion slightly changes about 99%. An “effectiveness factor” (EF) method, conducted by Liu et al,8 was developed to extrapolate char−CO2 reaction data to hightemperature condition, which was also introduced to a pressurized entrained plug flow coal gasification model to investigate the reaction kinetics and char structure on the gasification rate.9 HLA et al.10 performed the EF method into CFD to study the effects of coal type on the gasifier performance, and the predictions were validated by the experimental data, while the EF method has not been performed and validated in the numerical simulation of the industrial CWS gasifier, which operates at high temperature (about 1200−2000 °C) and pressure (4 or 6.5 MPG). In this paper, a three-dimensional steady-state numerical model is developed to simulate the gas−solid turbulent reacting flow in the OMB CWS entrained flow gasifier. The EF method is used to extrapolate the intrinsic char reactivity data to industrial CWS gasifier conditions, and the predictions are compared with industrial operating data. The flow, temperature, and species distributions are investigated; in addition, the effects of operating parameters such as oxygen/coal ratio and slurry concentration on the gasifier performance are studied.

1. INTRODUCTION Gasification of coal provides a means of generating a wide range of products such as power, chemicals, substitute natural gas (SNG), and transport fuels, which offers a high efficiency and low emission way to meet energy needs. By 2015, world gasification capacity is projected to grow by more than 70% and China is expected to achieve the most rapid growth.1 In China, on the basis of research of the cold model and hot bench model, the opposed multiburner (OMB) coal−water slurry (CWS) gasification process (presented in Figure 1) was developed by the Institute of Clean Coal Technology (ICCT), East China University of Science and Technology (ECUST), and successfully applied to synthesis chemicals such as methane and ammonia. So far, there are 43 gasifiers in design and 19 gasifiers in operation across China.2 The OMB gasifier (presented in Figure 2) is featured with four horizontal impinging burners in order to intensify mixing feeds and increase the residence time of solids. For optimizing design and scaling up of industrial gasifier, it is necessary but extremely difficult to measure flow, temperature, and species distributions inside the gasifier. Numerical simulation on coal gasification was developed by researchers3−7 in past decades, which offers an effective technique for predicting gasification characteristics and optimizing operation parameters. In most previous studies, the Eulerian−Lagrangian method was used to describe gas- and particle-phase flow. The standard k−ε turbulence model3,5 or realizable k−ε turbulence model6 was used to evaluate the gas flow in the different constructions of gasifier. The mixture fraction model3 or eddy-break-up model5 was adopted to simulate the competing interactions between the chemical reaction and turbulent mixing. Additionally, knowledge of the char reactivity is also essential for modeling the gasification process and designing an efficient gasifier. The intrinsic reactivity data of different coal are typically obtained under conditions where temperature and pressure are lower than those in the industrial entrained flow gasifier. © 2012 American Chemical Society

2. COAL REACTION SUBMODELS Under high temperature and high pressure in the gasifier, there are complicated physical and chemical processes occurring to CWS particles: slurry atomization, water evaporation, coal pyrolysis, and homogeneous and heterogeneous char reactions. 2.1. Slurry Atomization. The burner is a three-channel airblast atomizer. The CWS (annular channel) is fed into the gasifier and then accelerated by high-speed pure oxygen (the first and third channels) to atomize fully into droplets. The droplet Received: Revised: Accepted: Published: 2560

July 15, 2011 December 24, 2011 January 11, 2012 January 11, 2012 dx.doi.org/10.1021/ie201542q | Ind. Eng.Chem. Res. 2012, 51, 2560−2569

Industrial & Engineering Chemistry Research

Article

Figure 1. Overview of the OMB CWS gasification process.

where Po is the local absolute pressure and Xw is the mole fraction of steam. Psat., the saturation pressure of steam, is calculated by Antoine equation:11 Psat. = 0.133289 exp [18.3030 − 3816.44/(Tp − 46.13)]. kc is the mass-transfer coefficient and obtained from the Sherwood number correlation:12,13

Sh = kcd p/Di ,m = 2.0 + 0.6Re1/2Sc1/3

(2)

2.3. Coal Pyrolysis. The coal pyrolysis process is assumed to be two steps: coal devolatilization and volatiles simultaneous decomposition.

coal → volatiles + char(C(s) ·ash)

(R2)

volatiles → α1CH4 + α 2CO + α3CO2 + α4H2 6

Figure 2. Schematics of industrial OMB gasifier.

+ α5N2 + α6H2S

The coal devolatilization rate is determined by Kobayashi model:14

dm v = (m p − ma )[Y1A1exp( − E1/RTp) dt + Y2A2exp( − E 2 /RTp)]

45 < dp < 70 65 4.0 16

70 < dp < 380 170 1.6 46

380 < dp < 680 450 4.7 3

2.2. Water Evaporation. In the gasifier, the atomized CWS particles are quickly heated and evaporated. The evaporation process can be described as

coal−water → coal + H2O(g)

CO + 0.5O2 → CO2

(R1)

H2 + 0.5O2 → H2O

The evaporation rate is governed by bulk steam partial pressure and steam saturation pressure at the particle surface.

⎛ Psat.(Tp) P ⎞ dm w = − kc⎜⎜ − X w o ⎟⎟A pM w RT ⎠ dt ⎝ RTp

(3)

where the yield factors are Y1 = 0.3 and Y2 = 1, the activation energies are E1 = 104.6 kJ/mol and E2 = 167.4 kJ/mol, and the frequency factors are A1 = 2 × 105 s−1 and A2 = 1.3 × 107 s−1. The products of volatile decomposition are supposed to be composed of CO, CO2, H2, CH4, N2, and H2S. The composition of each species is determined by the ultimate analysis of the coal. The change of particle diameter during coal devolatilization is evaluated by the swelling coefficient Cw. 2.4. Homogeneous Reactions. Simple global reactions are used to describe homogeneous chemistry, including fuel gases combustion, water-gas shift, and methane−steam reactions.

Table 1. Parameters of Rosin−Rammler Distribution Function 0 < dp < 45 35 1.5 35

(R3)

i=1

size distributions are considered to be consistent with coal particle size distributions at the grinding mill, and coal particle size distributions are measured by Mastersizer2000 and fitted by Rosin−-Rammler expression. The Rosin−Rammler distribution function is Yd = exp((−dp/dav)n), and its the parameters are shown in Table 1.

dp, μm dav n mass fraction, %

∑α = 1

− 283kJ/mol − 242kJ/mol

CH4 + 0.5O2 → CO + 2H2

(1) 2561

− 35.7kJ/mol

(R4) (R5) (R6)

CO + H2O ⇔ CO2 + H2

− 41.1kJ/mol

(R7)

CH4 + H2O ⇔ CO + 3H2

+ 206kJ/mol

(R8)

dx.doi.org/10.1021/ie201542q | Ind. Eng.Chem. Res. 2012, 51, 2560−2569


Article

Table 2. Homogeneous Chemical Reaction Rate reaction rate expression ((kmol/m3)/s)

reaction combustion

17.6

R 4,f = 1.0 × 10

water-gas shift

0.5

exp( − 20130/ T )[CO][H2O]

source 0.25

[O2 ]

Dryer and Glassman16

R5,f = 2.85 × 1016T −1 exp( − 20130/T )[H2]0.25 [O2 ]1.5

Jones and Lindstedt17

R 6,f = 5.44 × 1012 exp( − 24358/T )[CH 4]−0.3 [O2 ]1.3

Westbrook and Dryer18

R7,f = 7.40 × 108 exp( − 36676/ T )([CO][H2O] − [CO2 ][H2]/K7)

methane−water

K7 = exp( − 3.6893 + 7234/1.8/T )

Bustamante and Enick19

R8,r = 3.0 × 108 exp( − 15100/ T )([CH 4][H2O] − K8[CO][H2]3 )

Jones and Lindstedt17

K8 = 5.12 × 10−14 exp( − 27347/T )

In the complex turbulent reacting flow, the net homogeneous reaction rate is controlled by chemical reaction rate Ri,r and turbulent mixing rate Ri,t, and the minimum of the above is taken as the limiting rate. As shown in Table 2, the Arrhenius rate is used to describe the detailed chemical reaction. Ri,t is calculated by the eddy-break-up (EBU) model:15

⎛ Y Yi ,p ⎞ ε i ,r ⎟⎟ , R i ,t = υ′i ,rMi ,rC rρ min⎜⎜ ′ ′ k M M υ υ ⎝ i ,r i ,r i ,p i ,p ⎠

where x is the carbon conversion of char particle, Ai is the frequency factor, Ei is the activation energy, Pi is the partial pressure of the gasifying agent, and n is the apparent order. The extrapolated char gasification rate to high pressure will be overpredicted by the nth-order model due to not taking the saturation of adsorbed surface complexes22 and the inhibition by produced gases into account. And it requires further exploration of the mechanisms for different coals by the experimental study in future work. With the increase of reaction temperature, the change in the rate-determining step from a chemical reaction control to a pore diffusion hindered is shown in Figure 3. The

(4)

where υi,r′ and υi,p ′ are stoichiometric coefficients for reactant and product, respectively. Yi,r and Yi,p are mass fractions of the reactant and product.Cr is an empirical constant equal to 4.0. The water-gas shift and methane−steam reactions are assumed to proceed to chemical equilibrium at all locations, and the forward reaction rate and equilibrium constant are used to calculate the chemical reaction rate.

C + 0.5O2 → CO

− 111 kJ/mol

(R9)

C + CO2 → 2CO

+ 172 kJ/mol

(R10)

C + H2O → CO + H2

+ 131 kJ/mol

(R11)

2.5. Heterogeneous Char Reactions. Heterogeneous reactions of char with oxygen, carbon dioxide, and steam are rate-determining steps in the coal gasification processes. According to previous studies,20,21 char reaction rate depends on the coal properties, thermal history, partial pressure of gasifying agent, total pressure, and so on. To obtain more reliable reactivity data, the pyrolysis char is prepared at 1400 °C by drop-tube furnace under atmospheric pressure at the ICCT laboratory. The reactivity data under chemical reaction control are obtained by pressurized thermogravimetric analyzer (PTGA; P, 0.1−1 MPa; T, 800−1000 °C) and fitted well by random pore model as shown in

Figure 3. Rate controlling regimes for heterogeneous char reaction.

effectiveness factor, η, as presented in eq 7, is introduced to extrapolate reactivity data to high temperature in regime II. The modified Thiele number, ϕ, and correction function, fc, conducted by Hong et al.,23 are used to calculate the effectiveness factor.

1⎡ 1 1 ⎤ η = fc ⎢ − ⎥ ϕ ⎣ tanh(3ϕ) 3ϕ ⎦

dx = ηA i Pi n exp( − Ei /RT )(1 − x) dt 1 − ψ ln(1 − x)

2

⎡ ⎤0.5(1 − n) 1/2 ⎥ fc = ⎢1 + ⎣ 2ϕ2 + 1/(2ϕ2) ⎦

(5)

x = 1 − m ci / m c 0

(6)

φ=

dp 6

(7)

n−1 (n + 1) A i exp( − Ei /RTP)(1 − x) 1 − ψ ln(1 − x) ρpRTpPi 2 McDeff 2562

(8)

(9)



Article

μ ⎞ ∂ε ⎤ ∂ ∂ ⎡⎛ ⎢⎜μ + t ⎟ ⎥ + ρC1S ε ( ρ̅ uĩ ε) = ∂xi ∂xi ⎢⎣⎝ σε ⎠ ∂xi ⎥⎦

where the effective diffusivity, Deff, based on the polymodal model,24 can be calculated by the following equations:

Deff = εmacro2Dmacro + εmeso2Dmeso

− ρC 2

2

+ (1 − εmacro − εmeso) Dmicro + 2εmacro(1 − εmacro − εmeso) 2 × (1/Dmacro + 1/Dmicro) + 2εmeso(1 − εmacro − εmeso) 2 × (1/Dmeso + 1/Dmicro) 2 + 2εmacroεmeso (1/Dmacro + 1/Dmeso)

dt

=−

νShMcDi A pPi RTmd p

(10)

dTp dt

= Ap

(18)

λNu (T − Tp) + A pεa(Q R − σBTp 4) dp

∑ mc, iQ i i

(19)

where mw, mv, and mc are particle mass transport due to water evaporation, coal pyrolysis, and char gasification, respectively. hw is the latent heat of evaporation, hv is the heat of pyrolysis, and Qi is the reaction heat of char gasification. The Nusselt number, Nu, is evaluated by the Ranz−Marshall method correlation:12

hd p

Nu =

λ

= 2.0 + 0.6Re1/2Pr1/3

(20)

Because the particle phase is dilute in most regions of the OMB gasifier,26 particle interactions are not considered in this simulation. A stochastic particle-tracking model is adopted to consider the effect of turbulent fluctuation on particle dispersion. With the Lagrangian method, particle motion is calculated by

(13)

where Sm and Sui are the source terms of particle mass and momentum, respectively. On the Boussinesq hypothesis, the Reynolds stress is expressed as follows: ρui′uj′ = (2/3)ρkδij + μt((∂ũi/∂xj) + (∂ũj/∂xi)). The turbulent viscosity is given by μt = Cμρk2/ε. According to the better predictions of the cold flow distribution in the OMB gasifier,24 the realizable k−ε model was recommended to solve the turbulent model. The transport equations of turbulent kinetic energy, k, and turbulence dissipation rate, ε, are shown as follows: μ ⎞ ∂k ⎤ ∂ ∂ ⎡⎛ ⎢⎜μ + t ⎟ ⎥ + Gk + G b − ρ̅ ε ( ρ̅ uĩ k) = ∂xi ∂xi ⎢⎣⎝ σk ⎠ ∂xi ⎥⎦

dmc dm w dm v + + dt dt dt

=

+ mwhw + mvhv −

⎡ ∂uj̃ ⎤ ∂ ∂P ̅ ∂ ⎢ ∂uĩ ⎥ + μt + ( ρ̅ uĩ uj̃ ) = − ∂xj ∂xi ∂xj ⎢⎣ ∂xj ∂xi ⎥⎦ + ρgi + Sui

(17)

m pCp

(12)

∂xj

μ ⎞ ∂Y ∂ ∂ ⎛ (ρuiYi) = ⎜ ρ̅ Di ,m + t ⎟ i + SYi + R Yi ∂xi Sc t ⎠ ∂xj ∂xj ⎝

dt

3. GOVERNING EQUATIONS AND SOLUTION METHODS 3.1. Gas-Phase Model. The transport equations of mass and momentum for governing the gas flow are given by

−

(16)

dm p

where ν is the stoichiometric coefficient, Sh is the Sherwood number, and Tm is the arithmetic mean value of particle temperature and gas temperature.

∂(ρui′u′j)

⎛ ⎞ ∂ ∂ ⎜ λ ∂h ⎟ (ρuih) = + Sh ∂xi ∂xj ⎜⎝ Cp ∂xj ⎟⎠

where Yi is the species mass fraction, and the enthalpy is T calculated by the following equation: h = ∑j Yj∫ 298.15 Cp,j dT. λ and Di,m are the thermal conductivity and diffusion coefficient, respectively. Sct is the turbulent Schmidt number equal to 0.7. Sh and SYi are the source terms of particle enthalpy and species, respectively, and RYi is the source term of chemical reactions. 3.2. Particle-Phase Model. The mass and energy transport equations of particles are given by

(11)

∂ ( ρ̅ uĩ ) = Sm ∂xi

(15)

where the standard constants of realizable k−ε model used were given as follows: C1ε = 1.44, C2 = 1.92, σk = 1.0, σε = 1.2, and other constants can be found in ref 6. The conservation of energy and chemical species mass fraction are taken as the following general form:

where Dmacro, Dmeso, and Dmicro denote the effective diffusivities for macro-, meso-, and micropores, respectively . εmacro, εmeso, and εmicro are macro-, meso-, and microvoid fractions, respectively,8 which take the bulk and Knudsen diffusion into account. For bulk diffusion limitation, the reaction rate is

dm c i

ε2 ε + C1ε C3εG b k + νε k

mp

du p dt

=

1 πd p2ρCD|u ̅ − u p|(u ̅ − u p) 8 1 + πd p3(ρp − ρ)g 6

(21)

where CD, the drag coefficient, is a function of relative Reynolds number Re as shown in eq 22:

CD =

(14) 2563

24 (1.0 + 0.15Re 0.687) Re

(22)



Article

3.3. Heat-Transfer Model. Mechanisms of the heattransfer process in the gasifier are presented in Figure 4. The

4. APPLICATION TO INDUSTRIAL OMB CWS ENTRAINED FLOW GASIFIER The numerical simulation for industrial OMB CWS entrained flow gasifier as presented in Figure 2 was performed in this study. The gasifier operates at an elevated pressure of 5.9 MPaG, and the feed loading of CWS and oxygen are listed in Table 3. The Shenfu coal is used in the industrial operation, Table 3. Operating Conditions of Industrial OMB Gasifier CWS concentration, wt %

60.5

CWS feed rate per burner, m3/h

18

oxygen feed rate per burner, N m3/h oxygen concentration, vol %

8490 99.8

and the properties are shown in Table 4. The random pore model is adopted to evaluate the reactivity data of Shenfu char Table 4. Properties of Shenfu Coal Proximate Analysis, wt % volatile matter, daf fixed carbon, daf ash, dry Ultimate Analysis, wt %

Figure 4. Mechanisms of the heat-transfer process in the gasifier.

heat exchange through convection and radiation between gases and particles are evaluated by the first and second terms on the right side of eq 19. According to reasonable performance for combustion applications,27 the P-1 model is used for solving the radiative heat-transfer equation, and the change in radiative intensity over the path length ds is

σ dI = − (εa + ε p + εs)I + B (εaT 4 + ε pTp 4) π ds 4π εs I dΩ + 4π 0

∫

C, daf H, daf N, daf S, daf O, daf HHVd, MJ/kg ash fluid temperature (FT), °C

83.30 4.65 0.97 0.22 10.86 30.14 1170

under chemical reaction control, and model parameters are listed in Table 5. The macro-, meso-, and microvoid fractions are 0.07, 0.72, and 0.21, respectively.

(23)

where the radiative absorption coefficient of gas εa is a function of local species concentrations, path length, and total pressure, which is evaluated by the weighted-sum-of-gray-gases model (WSGGM).28 εp and εs represent the absorption and scattering coefficients of particle, respectively. The heat flux from fluid cell to wall boundary is computed as

⎛ δ⎞ Q wall = (Tg − Tw )/⎜ ∑ i ⎟ λi ⎠ ⎝

36.65 63.35 8.85

Table 5. Heterogeneous Char Reaction Kinetics (Shenfu Coal) gasification agent −1

−n

A (s ·MPa ) E (kJ/mol) n Ψ

O25

CO2

H2O

1.36 × 106 130 0.68 14

3.78 × 104 178 0.53 2

1.33 × 107 226 0.60 2

The boundary conditions and computational grids for simulation are shown in Figure 5. Taking the symmetry of gasifier geometry and computational cost into account, the onequarter of gasifier is adopted as the calculation domain. The number of grids is determined as 260 000 through the griddependent test. The CWS particles are represented by 27 600 particle tracks. The liquid slag layer is thick at the wall below −4 m, where the particles are trapped, and reflected at other wall regions. The heat flux at the wall is calculated by eq 24; the wall temperature Tw is measured about 190 °C. The thickness and thermal conductivity for the metal shell are 100 mm and 17 W/(m·K), respectively, and for refractory are 555 mm and 1.6 W/(m·K). The thickness and thermal conductivity for liquid slag layer is 7 mm and 0.6 W/(m·K), and the thickness of the solid slag layer is neglected.29 4.1. Model Validation. For industrial gasifier operates at high temperature and high pressure, the syngas compositions are measured by gas chromatography at the outlet of scrubber tower. The temperature is measured by thermocouples near the

(24)

where the wall temperature T w is measured by the thermocouple on the surface of steel metal. δi and λi represent the thickness and thermal conductivity, respectively, of steel metal, refractory, solid slag layer, and liquid slag layer. 3.4. Numerical Solution Methods. The governing equations for the conservation of mass, momentum, energy, and turbulence are solved sequentially using the finite-volume method (FVM), on Fluent 6.2. The velocity correction is realized to satisfy continuity through the SIMPLE algorithm which couples velocity and pressure. To evaluate the convective terms and species, the second-order upwind scheme is used. The standard scheme is used for pressure discretization, and the first-order upwind scheme is used for other terms (turbulent kinetic energy, dissipation rate, and P-1 model). The current char gasification reaction model is added to Fluent via userdefined functions (UDFs). 2564



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Figure 5. Boundary conditions and computational grids for simulation.

Table 6. Model Validation for the Industrial OMB Gasifier dry basis, mol % industrial operating data simulation data

CO

H2

CO2

CO + H2

carbon conversion, %

T, °C

46.77 46.64

35.10 35.52

17.62 17.13

81.87 82.16

98.0 98.1

1220 1230

Figure 6. Velocity and temperature distributions in OMB gasifier.

flame of about 2100 °C. Near the outlet of the burner, the droplets are heated and evaporated to form a low-temperature region, which could protect the burner from ablation by hightemperature gas. In the impinging flow region, most volatiles generate and combust with oxygen leading to higher temperature at 2360 °C. According to Figure 7, in the impinging−jet flow region the main reactions are char−CO2 reaction, char−steam

wall inside the gasifier, and its specific position is presented in Figure 2. With the proposed model, the predicted performances for the OMB gasifier are in good agreement with the industrial operating data (as shown in Table 6). 4.2. Flow Field, Temperature, and Composition Distributions. The velocity distribution of the X−Y plane is presented in Figure 6a; the flow field in the OMB gasifier can be categorized into five regions: jet flow region, impinging flow region, impinging−jet flow region, reflux region, and plug flow region. The pure oxygen injects into the gasifier at high speed of about 120 m/s in order to atomize CWS fully into small droplets. The turbulent jet flow is fully developed and then decayed. In the impinging flow region, due to higher turbulence shearing, the gas radial velocity decreases less than 10 m/s, which is near the stagnation point, and then the gas reflects to the axial direction. Therefore, the average residence time of coal particles in the impinging flow region will increase due to the decay of gas radial velocity. The temperature distribution of the X−Y plane is presented in Figure 6b. In the jet flow region, the high-temperature hydrogen and carbon monoxide are entrained into the oxygen-slurry jet and combust with oxygen rapidly to form a high-temperature

Figure 7. Reaction rate versus axial position of central line. 2565



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reaction, and the reverse water-shift reaction. These endothermic reactions reduce the gas temperature dramatically, and at the top of the dome the gas temperature reaches 1300 °C, which prolongs refractory life. As shown in Figure 8, the volume fractions of CO and H2 increase along the central line at the axial direction, while CO2 and H2O decrease. 4.3. Performance of OMB Gasifier. In this work, the Nusselt number (Nu) and Sherwood number (Sh) are adopted to describe the heat- and mass-transfer of particles in the OMB gasifier. As shown in Figure 9, for particle diameter equal to 200 μm, Nu and Sh are 40 and 16 in the impinging flow region, respectively, whereas those are lower in the reflux region at about

10 and 6. The reason is that four feeds are collided at the impinging flow region to produce a high relative velocity and intensify the mass and heat transfer between gas−solid phases. Consequently, high-transfer rate favors higher carbon conversion. The carbon conversion is an important parameter to characterize the gasifier performance, which is affected by the motion and reaction condition of particles. The carbon conversion and residence time of different particle size are shown in Table 7, which reveals that smaller particles follow the gas phase closely and have a longer residence time in the gasifier. The char fraction of representative particles in the OMB gasifier is shown in Figure 10. For particle diameter

Figure 8. Gas mole fraction distributions of X−Y plane in OMB gasifier.

Figure 9. Nusselt and Sherwood number distributions of axial section through the burner plane. 2566



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Table 7. Effects of the Particle Size on Carbon Conversion and Residence Time in the OMB Gasifier

Table 8. Effects of Oxygen/Coal Ratio on the OMB Gasifier Performance at Fixed Slurry Concentration of 60.5%

dp, μm

carbon conversion, % average residence time, s

0 < dp < 45

45 < dp < 70

70 < dp < 380

380 < dp < 680

99.9 9.22

99.9 9.35

99.3 7.88

87.6 7.11

60 μm, the carbon is almost consumed at the jet flow region and impinging flow region. While the particle diameter is 600 μm, few particles are entrained into the reflux region and carbon is partially consumed at the outlet of the gasifier. 4.4. Effects of the Oxygen/Coal Ratio and Slurry Concentration on OMB Gasifier Performance. Effects of the oxygen/coal ratio on the OMB gasifier performance are shown in Table 8. As the oxygen/coal ratio increases, the cold gas efficiency decreases, while the carbon conversion slightly increases. When more oxygen is added, the CO and H2 are combusting competitively to produce CO2 and steam, and the gas temperature increases in the jet flow region and impinging flow region. The above two factors contribute to a higher rate of the char−CO2 and char−steam reactions and promote the reverse water-shift reaction rate in the impinging−jet flow region and plug flow region. Therefore, the mole fraction of CO at the outlet of the gasifier slightly changes whereas those of CO2 and H2 decrease. On the assumption that CWS is atomized fully under the various slurry concentrations, the effects of slurry concentration on OMB gasifier performance are studied (Table 9). The results reveal that higher slurry concentration leads to higher cold gas efficiency, and the carbon conversion decreases slightly. Higher slurry concentration means that less water is added and the exit gas temperature increases. Additionally, the gas composition at the outlet of the gasifier changes slightly.

O2/dry coal (N m3/kg)

0.651

0.659

0.667

0.674

0.686

CO, mol % H2, mol % CO2, mol % H2O, mol % T, °C carbon conversion, %

37.15 28.30 13.65 20.33 1231 98.1

37.38 27.55 13.50 20.89 1261 99.0

37.58 27.12 13.33 21.29 1277 99.4

37.75 26.63 13.13 21.81 1301 99.6

37.92 25.90 13.01 22.49 1324 99.7

Table 9. Effects of Slurry Concentrations on the OMB Gasifier Performance at Fixed Oxygen/Coal Ratio of 0.659 N m3/kg slurry concentration, wt %

59.5

61.0

62.0

CO, mol % H2, mol % CO2, mol % H2O, mol % T, °C carbon conversion, % cold gas efficiency, %

37.01 27.60 13.66 21.05 1245 98.8 73.59

37.57 27.73 13.46 20.55 1248 98.6 73.84

37.96 27.86 13.32 20.17 1251 98.4 74.11

5. CONCLUSIONS A comprehensive three-dimensional numerical model for simulation on the industrial OMB CWS gasifier is established. The “effectiveness factor” method is used to extrapolate the intrinsic char reactivity data to industrial CWS gasifier conditions. With the proposed models, the predicted temperature and gas compositions at the outlet of the gasifier are in good agreement with industrial operating data. The flow field in the OMB gasifier can be categorized into five regions: jet flow region, impinging flow region, impinging− jet flow region, reflux region, and plug flow region. In the impinging−jet flow region and plug flow region the main

Figure 10. Char fraction of representative particles in OMB gasifier. 2567



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reactions are char−CO2 reaction, char−steam reaction, and the reverse water-shift reaction. In addition, the Nusselt number and Sherwood number of particles are investigated, which imply that the heat and mass transfers between gas−solid phases are intensified in the impinging flow region. As the oxygen/coal ratio increases, the cold gas efficiency decreases, while the carbon conversion slightly increases. The mole fraction of CO at the outlet of the gasifier slightly changes, whereas those of CO2 and H2 decrease. The higher slurry concentration leads to higher temperature and cold gas efficiency, while the carbon conversion decreases slightly.

■

u and up xi and xj Yi

gas and particle velocity, m s−1 coordinate of directions, m mass fraction of species i

δij ε εa and εp

1 (i = j), =0 (i ≠ j) turbulence dissipation rate, m−2 s−3 radiative absorption coefficients of gas and particle, m−1 macro-, meso-, and microvoid fraction, respectively particle radiative scatter coefficient, m−1 thermal conductivity of the gas phase, W m−1 K−1 effectiveness factor dynamic viscosity, kg m−1 s−1 turbulent viscosity, kg m−1 s−1 gas and particle densities, kg m−3 Stefan−Bolzmann constant, 5.67 × 10−8 W m−2 K−4 Thiele modulus random pore model constant solid angle

Greek Letters

εmacro, εmeso, and εmicro εs λ

AUTHOR INFORMATION

η μ μt ρ and ρp σB

Corresponding Author

*Tel.: +86-21-6425 2974. Fax: +86-21-6425 1312. E-mail: [email protected].

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ACKNOWLEDGMENTS This work is financially supported by the National Key State Basic Research Development Program of China (973Program, Grant 2010CB227006), the National High Technology Research and Development of China (863 Program, Grant 2009AA04Z159), and the National Nature Science Foundation of China (Grant 20876048).

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REFERENCES

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NOMENCLATURE frequency factor surface area of the particle, m2 drag coefficient specific heat capacity of particle, J kg−1 K−1 specific heat capacity of gas species j, J kg−1 K−1 Cw swelling coefficient dp particle diameter, m Di diffusivity of species i in the gas mixture, m2 s−1 Deff effective diffusivity, m2 s−1 Dmacro, Dmeso, Dmicro effective diffusivity for macro-, meso-, and micropores, respectively, m2 s−1 E activation energy h enthalpy, J kg−1 hw, hv latent heat of evaporation and heat of pyrolysis, respectively, J kg−1 I radiative intensity, W m−2 s−1 k turbulent kinetic energy, m2 s−2 Mw and Mc molecule of water and carbon, g mol−1 mw, mv, and mc particle mass transport due to water evaporation, coal pyrolysis, and char gasification, kg mp particle mass, kg Nu Nusselt Number P pressure, Pa QR radiative heat flux, J m−2 s−1 R molar gas constant, 8314 J kmol−1 K−1 Re relative Reynolds number Sh Sherwood number the source terms of particle mass, Sm, Sui, Sh, SYi momentum, enthalpy and species, respectively Sct turbulent Schmidt number s path length, m T and Tp gas and particle temperature, K A Ap CD Cp Cp,j

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