Investigation of the Gasification Performance of Lignite Feedstock and the Injection Design of an E-Gas Like Gasifier

Proceedings of the 29th Annual International Pittsburgh Coal Conference, Pittsburgh, PA, USA, October 15 - 18, 2012

Investigation of the Gasification Performance of Lignite Feedstock and the Injection Design of an E-Gas like Gasifier

Yan-Tsan Luan and Yau-Pin Chyou* Ting Wang Chemistry Division Energy Conversion & Conservation Center Institute of Nuclear Energy Research University of New Orleans Atomic Energy Council New Orleans, Louisiana 70148, USA Longtan, Taoyuan, Taiwan (R.O.C.)

ABSTRACT CO2 capture among the four reference power plants: coal In the last three years, the Institute of Nuclear Energy combustion, integrated gasification combine cycle (IGCC), Research (INER) has been developing the E-GAS gasification oxy-combustion, and natural-gas combustion. In the second numerical model and analyzing the gasification performance column, coal pre-combustion carbon capture has the lowest by conducting several parametric studies. A preliminary cost of CO2 avoided, which makes it more competitive and a numerical model considering coal particles tracking, two-step welcoming choice when it comes to low-carbon electricity volatiles thermal cracking, and nine chemical reactions has generation. With its inherence of low emissions and cost been established. In last year’s results, the single lateral including CCS, IGCC has the potential to replace traditional injector design in the 2nd stage of E-GAS gasifier is found PC plants. Therefore, conducting research in coal gasification non-ideal for the gasificaiton process. Therefore, one of the and CO2 capture technology to improve its availability, objectives in this study is to modify the 2nd stage injection increase its efficiency, and reduce the cost is essential for and investigate its effect on gasification performance. achieving sustainability of energy utilization. Moreover, because of rising interest in the use of low-rank coals as the feedstock for power plants, the second goal in this paper is to investigate the gasifciation features by using North Dakota (N.D.) Lignite as the feedstock. The result shows that a dual-injector design (either tangential or opposing jets) in the 2nd stage injection of E-GAS gasifier can minimize the non-ideal recirculation zone and improve the gasification performance. Furthermore, a tangential injection design can make the average temperature reach equilibrium more quickly, so the height of the E-GAS gasifier could therefore be shortened. Moreover, by examining the energy needed for the 2nd stage injection, a tangential Fig. 1 Growth in total primary energy demand [1] injection design saves more energy as compared to the opposing-jets design. The result of the assessment of N.D. Lignite shows that under the same condition of O2/Coal and Coal/Slurry ratios, no matter what the feedstock flow rate or the total input heating value, the gasification performance of N.D. Lignite is always lower than that of Illinois #6 Coal. However, the cheaper price of N.D. Lignite makes it more competitive, and if the electricity generating industries can accept a lower gasification performance, the usage of N.D. Lignite can be a choice to be considered.

Key words: gasiﬁcation CFD modeling, low-rank coal, E-GAS entrained-flow gasifier, clean coal technology Fig. 2 The average cost and performance impact of adding CO2 capture in OECD countries [1]

1. INTRODUCTION According to the data from Bureau of Energy of Taiwan 1.1 Background [2], the primary energy supply situation of Taiwan in 2010 is as follows: coal (32.09%), oil (49.04%), natural gas (10.16%), According to the data from IEA (Fig. 1) [1], the and nuclear power (8.28%). Only considering the electricity increasing demand for coal dominates the demand for all other generation in 2010, the gross power generation reached energy sources in the last ten years. The task to meet the 247,045.4 GWh, which has an increase of 7.55% from 2009. energy demand and achieve the low-carbon emission goal has The electricity generation consisted of fossil-fuel fired power become an urgent issue world-wide in the past decade. 79.79%, nuclear power 16.85%, and geothermal, solar and Compared to the other fossil-fuel resources such as crude oil wind power 0.42%. In 2008, the government set the following and natural gas, coal has the largest reserves and least objective and prospect: (1) raise energy efficiency by 2% expensive price for producing electricity. Figure 2 shows the yearly over the next 8 years, (2) develop clean energy and comparison of average cost and performance impact of adding

*Corresponding author (e-mail: [email protected])

1 reduce CO2 emission in 2025 to the amount of emission in mathematical model and assessed the influences of the input 2000, (3) ensure a stable energy supplement. operating parameters. They concluded that: (1) an increase of Considering the guiding principle of the sustainable oxygen/fuel ratio significantly increases the carbon energy policy and the prospect that fossil-fuel fired power conversion, resulting in slightly decreased H2 concentration plants will still be the main source of electricity generation in and increased CO concentration in the product gas; (2) an Taiwan as well as world-wide in the next few decades, the increase in the steam/fuel ratio increases the H2 and CO2 goal for Taiwan’s energy policy is to forge ahead toward a concentration, but it decreases the CO concentration; (3) an high efficiency, environment friendly, affordable, and stable increase in the operating pressure can increase the degree of supplement energy development. carbon conversion particularly at high steam/fuel ratios; and The Institute of Nuclear Energy Research (INER) is one (4) small fuel particles cause higher carbon conversion energy RD&D organization in Taiwan. INER has been efficiencies for a given mass feed rate. devoted to developing clean coal technology and in 2012 Over the past few years, comprehensive research has started to construct a demonstration gasification facility.. been devoted to the gasification process and performance via Since 2009, the Computational Fluid Dynamics (CFD) model CFD modeling [7-19]. Chen et al. [20] implemented a was developed to investigation and gain knowledge of the three-dimensional model to simulate a 200-ton, two-stage air- gasification process in entrained-bed gasifiers. Several blown entrained type gasifier developed for an IGCC process. achievements have been accomplished in the last two years. In The numerical analysis showed turbulent fluctuations 2010, the gasification model was simplified by assuming that affected the volatiles and char-oxygen reactions and the carbon solid was instantaneously gasified; the reaction rate significantly influenced the temperature and gas composition. of each reaction was determined by the turbulence eddies via Furthermore, Chen et al. [21] also found carbon conversion the Eddy-Dissipation Model [3, 4]. In 2011, a more realistic is independent of the devolatilization rate and less sensible to Finite-Rate Model was employed to replace the previous coal particle size, but it is sensitive to the heterogeneous instantaneous gasification model [5]. Moreover, heat and char-oxygen, char-CO2, and char-steam reaction kinetics. momentum transfer between coal particles and the Since 2005, numerous attempts have been made by the stoichiometric tracking method considering the turbulence research team of the Energy Conversion & Conservation effect were investigated as well by the Discrete Phase Model Center (ECCC) at the University of New Orleans [22-28] to (DPM). This comprehensive CFD model has significantly develop a gasification numerical model through contributed to the investigation and understanding of the ANSYS/FLUENT. The model is based on a two-stage, thermal-flow behavior in the unique design of E-GAS gasifier. entrained-flow gasifier following the geometry and input Several operating parameters including O2/Coal ratio, parameters of Bockelie et al. [10] and Chen et al. [20]. They coal slurry concentration ratio, and 1st-2nd stage fuel feeding investigated various operating conditions, and the results are distribution ratio have been studied in [5]. During the previous summarized as follows [23]: (1) coal powder feedstock tends study, it was discovered that the strong lateral single injection to generate more CO and a higher exit temperature than coal design in the 2nd stage of the existing E-Gas gasifier slurries with equivalent coal mass flow rate; (2) the exit gases produced a non-ideal large lopsided flow recirculation of oxygen-blown gasifiers have higher mole concentrations condition that might impede gasification performance in of CO2 and CO than that of air-blown gasifiers; (3) one-stage syngas-production uniformity and flow efficiency. Hence, one operation yields higher H2, CO, and CH4 than if a two-stage of the objectives of this paper is to investigate the potential operation is used but with a lower syngas heating value. benefits of modifying, in the second stage, a single-injection Silaen and Wang [22] found that the water-gas-shift to a twin tangential-injection arrangement. Considering cold (WGS) rate plays an important role in predicting syngas flow study, either experimentally or computationally, has composition, and the reaction rate of the WGS, adopted from often been used as a fast first-step preliminary study of flow Jones and Lindstedt [29], is found to be too fast because the pattern in a gasifier. Hence, the second objective is to compare rate was obtained with the presence of a catalyst. Considering the cold flow pattern with the reactive hot flow pattern and that no catalyst is added in a typical gasiﬁer, the water shift provide an insight into how much the cold flow pattern can be reaction rate is purposely slowed down to make the syngas trusted. In recent years, the usage of low-rank coal has been composition consistent with syngas in the actual production promoted because of its abundant reserves and low cost. of a commercial entrained-flow gasiﬁer with coal slurry feed Because of low-rank coal’s characteristics of high moisture, from the bottom. In a recent paper, Lu and Wang [30] volatiles, and ash content and low heating value, it is not concluded that all of the originally published WGS reaction economic to use as the feedstock of a traditional PC power rates were too fast when they calibrated the WGS reaction rate plant. However, through the continuing development of using the Japanese air-blown, dry-feed CRIEPI gasifier. gasification technology for accommodating flexible fuels, the Adding a backward WGS reaction rate doesn't slow the value of and interest in low-rank coal is increasing. Therefore, reaction rate too much resulting in the same gas composition the third objective of this paper is to investigate the feedstock and temperature at the gasifier exit as the case without adding effect on the output syngas composition and thermal energy the backward WGS reaction rate. And, they also clarified that by employing a low-rank coal, North Dakota (N.D.) Lignite, the calibrated WGS reaction rates were obtained under in the E-GAS gasifier. air-blown and dry-fed operating conditions. These calibrated WGS reaction rates may not be applicable to slurry-fed or 1.2 Literature Review oxygen-blown gasifiers because the higher water vapor concentration in slurry-fed gasifiers and higher operating In 1979, Wen et al. [6] investigated the Texaco temperatures in oxygen-blown gasifiers may affect the global down-draft entrainment pilot plant gasifier through their WGS rate [30]. 2

where the second term on the RHS is the source term for 2. NUMERICAL SCHEME particle-gas heat transfer, evaporation energy (latent heat), the radiation energy, and reaction heat. The CFD solver used in this study is the commercial CFD code ANSYS FLUENT V.12.0. FLUENT is a 2.2 Turbulence model finite-volume-based CFD solver written in C language and has the ability to solve fluid flow, heat transfer, and chemical The turbulence kinetic energy, k, and its rate of reactions in complex geometries, and it supports both dissipation, ε, are obtained from the following standard k-ε structured and unstructured meshes. The second-order transport equations: discretization scheme is applied. The SIMPLE algorithm is used in this study as the algorithm for pressure-velocity ∂ ∂  µ  ∂k  coupling.  t  (ρkui ) =  µ +   + Gk − ρε (5) The 3-D, steady-state Navier-Stokes equations are ∂xi ∂x j  σ k  ∂x j  solved in an Eulerian-Lagrangian frame of reference. All the and coal particles are treated as a discrete, secondary phase ∂ ∂  µ  ∂ε  ε ε 2 dispersed in the continuous phase via the Discrete Phase  t  (ρεui ) =  µ +   + C1ε Gk − C2ε ρ (6) Model (DPM) with the stochastic tracking to consider the ∂xi ∂x j  σ ε  ∂x j  k k turbulent dispersion effect. The particles are tracked for every 50th continuous phase flow iterations. The P1 radiation model In these equations, Gk represents the generation of is used, and the gravitational force is considered in the turbulence kinetic energy due to the mean velocity gradients. modeling. The standard k-ε model is used to capture the σ and σ are the turbulent Prandtl numbers for k and ε. The turbulence flow. The species transport equations are solved k ε turbulent (or eddy) viscosity, μt, is computed by combining k through the Finite-Rate/Eddy-Dissipation Model. In this and ε as follows: model, both the finite rate and the eddy-dissipation rate are used and compared, and the slower rate is selected to compute 2 the continuous phase reactions. k μ = ρC (7) t μ ε 2.1 Governing equations the empirical model constants are C = 1.44, C = 1.92, C = 1ε 2ε μ The equation for conservation of mass is: 0.09, σk = 1.0, σε= 1.3 [31].  ∇ ⋅ (ρυ) = Sm (1) 2.3 Discrete phase

The source term Sm is the additional mass including the In the Largrangian approach, each particle is tracked. vaporization of water droplets and the devolatilization of coal The force balance on the discrete phase particles can be particles added to the continuous phase from the dispersed written as second phase flow. The momentum conservation equation is: du p gx (ρ p − ρ) = F (u − u ) + (8) dt D p ρ     p ∇ ⋅(ρυυ) = −∇p + ∇ ⋅(τ ) + ρg + F (2) where FD(u-up) is the drag force per unit particle mass: where p is the static pressure, τ is the stress tensor, and   18µ CD Re ρg and F are the gravitational body force and external FD = (9) ρ d 2 24 body forces (e.g. that arise from interaction with the dispersed p p phase), respectively. ρd p u p − u Re = (10) The stress tensor τ is given by µ

  T 2   where u is the fluid phase velocity, up is the particle velocity, μ τ = µ(∇υ + ∇υ )− ∇ ⋅υI  (3)  3  is the molecular viscosity of the fluid, ρ is the fluid density, ρp is the particle density, and d is the diameter of the particle p which is treated as a spherical shape. Re is the relative where µ is the molecular viscosity, I is the unit tensor, and the Reynolds number calculated from the slip velocity between second term on the RHS is the e ect of volume dilation. the fluid phase velocity and the particle velocity. The energy conservation equation is: The track of the particle is evaluated by the ﬀ instantaneous velocity, which is related to the mean velocity ∂ ∂  ∂T  and the fluctuating velocity as: (ρu h) =  K  + S (4) i   ph ∂xi ∂xi  ∂xi  u* = u + u' (11)

3 where the fluctuating velocity in Eq. (11) is defined by the dTp 4 4 stochastic discrete random walk model. mpcp = hAp (T∞ −TP )+ e p Apσ (θR −Tp ) dt The relationships for calculating the burning rate of the (18) char particles are presented and discussed in detail by Smith dmp dmp + h fg − fh Hreac [32]. The reaction rate R (kg/s) of particle surface species dt dt depletion as a particle undergoing an endothermic reaction in the gas phase is given as: where mp and Tp are the particle mass and particle temperature, respectively. θ is the radiating temperature, R R = A ηYR (12) defined as: p N  R  θ = (G / 4σ )1/ 4 (19) R = Rkin  pn −  (13) R  D 

2 and G is the incident radiation, which is related to the where Ap is the particle surface area (m ), Y is the mass radiative intensity as: fraction of surface species in the particle, η is the effectiveness factor (dimensionless), R is the rate of particle surface species 4π 2 G = ∫0 IdΩ (20) reaction per unit area (kg/m -s), pn is the bulk concentration of the gas phase species (kg/m3), D is the diffusion rate coefficient for the reaction, Rkin is the kinetic rate of reaction The radiation in the gasifier is described by the P1 (units vary), and N is the apparent order of reaction (In this model [36]. The convection coefficient (Nusselt number, Nu) case, N is 0.5). The kinetic rate of reaction r is defined as: between the particle and the gaseous flow field is evaluated by the empirical relation proposed by Ranz and Marshall [33] as: n −(E / RT ) Rkin = AT e (14) hd Nu = p = 2.0 + 0.6Re1/ 2 Pr1/ 3 (21) When the coal slurry particle is injected into the k d gasifier, the water is assumed to be atomized to small droplets and undergoes evaporation. The rate of vaporization is 2.4 Devolatilization controlled by concentration difference between the surface and gas streams, and the corresponding mass change rate of The devolatilization reaction of the coal particle is the droplet can be given by: described by the two-competing-rates model proposed by Kobayashi [37] as:

dmp 2 =pd kc (Cs − C∞ ) (15) − dt = (E1 / RTp ) (22) R1 A1e −(E2 / RTp ) R2 = A2e (23) where kc is the mass transfer coefficient and Cs is the concentration of the vapor at the particle’s surface, which is evaluated by assuming that the flow over the surface is These two kinetic rates are used as a weight function by saturated. C∞ is the vapor concentration of the bulk flow, an expression for devolatilization as: obtained by solving the transport equations. The value of kc m (t) t can be calculated from empirical correlations by [33, 34]: v = a +a ∫ ( 1R1 2R2 ) (1− fω,0 )mp,0 − ma 0 (24) kcd 0.5 0.33 t Shd = =2.0 + 0.6 Red Sc (16)   D exp− (R1 + R2 )dt dt  ∫0  where Sh is the Sherwood number, Sc is the Schmidt number (defined as ν/D), D is the diffusion coefficient of vapor in the where α1 and α2 are yield factors, fω is the mass fraction of moisture, mp is the mass of particle, and ma is the mass of ash. bulk flow. Red is the Reynolds number, defined as uν/D, u is 5 7 The value of the constants are A1 = 2×10 , A2 = 1.3×10 , E1 = the slip velocity between the droplet and the gas and D is the 8 8 droplet diameter. 1.046×10 J/kg mol, and E2 = 1.67×10 J/kg mol. When the droplet temperature reaches the boiling point, As the volatiles diffuse out of the coal particle, the the following equation can be used to evaluate its evaporation heterogeneous reactions on the particle surface occur. Those rate [35]: particle surface reactions are described by the implicit relations of Smith et al. [32] as:

dmd 2 l  0.5 =pd  (2.0 + 0.46Red )ln(1+ C p (T∞ −T )/ hfg )/ C p (17) R = A η Y R (25) dt  d  k ,r p r k k ,r N  R  r =  − k,r  (26) The energy equation for the particle accounts for Rk,r Rkin,r  pn  D0,r  convection, radiation, devolatilization, and surface reactions,  −( ) R = A T n,re Er / RT (27) is given as: kin,r r

2.5 Chemical reaction Table 1 Global reaction rate constants in this study

Reaction Ar Er (J/kmol) The conservation equation of the species is in the C + 1/2O → CO 0.052 6.1E+07 following general form: (s) 2 C(s) + CO2 → 2CO 0.0732 1.125E+08   ∇⋅ ρυ = −∇⋅ + + C(s) + H2O → CO + H2 0.0782 1.15E+08 ( Yi ) Ji Ri Si (28) CO + 1/2O2 → CO2 2.2E+12 1.67E+08 where Ri is the net rate of production of species i by chemical H2 + 1/2O2 → H2O 6.8E+15 1.68E+08 reaction and Si is the rate of mass creation by addition from  CH2.761O0.264 → 0.479H2 + 0.264CO the dispersed phase sources. Ji is the diffusion flux of + 0.356CH4 + 0.19C2H2 Eddy-dissipation rate only species i, which arises due to the gradients in concentration C2H2 + O2 → 2CO + H2 and temperature. CH4 + 1/2O2 → CO + 2H2 The source term for species i due to all reactions is CO + H O → CO + H 2.75E+02 8.38E+07 computed as the sum of the Arrhenius reaction sources over 2 2 2 the NR reactions that the species participate in: 3. MODEL DESCRIPTION

N R = ˆ 3.1 Computational model Ri M w,i ∑ Ri,r (29) r=1 Most geometric dimensions of the E-GAS-like gasifier ˆ in this study are referred to in the NETL’s published document where Mw,i is the molecular weight of species i and R is i,r [38]. Part of the dimensions, including injector size and throat the Arrhenius molar rate of production/consumption of species height and diameter, is estimated from the published diagrams i in reaction r. in the public domain. The sketch of the computational model The Finite-rate/Eddy-dissipation-rate model computes is shown in Fig. 3. The E-GAS-like gasifier is divided into both Arrhenius rates and turbulent mixing rates, and the two sections, consisting of a horizontal vessel (width: 8m; smaller one is chosen for the homogeneous reactions. While diameter: 2m) and a vertical cylinder (height: 10m, diameter: only finite rates are used for the heterogeneous reactions. 1.6m). In the horizontal section (1st stage section), there are two injectors located at the height of 1m and opposed to each 2.6 Finite-rate/Eddy-dissipation-rate model other. However, there is only one injector located in the vertical section (2nd stage section) at the height of 3.6m. All 2.6.1 Arrhenius rate: of the oxidant consisting of 95% O2 and 5% N2 is fed into the 1st stage section. The coal slurry is injected in a two-stage N η′ ˆ ′′ ′ ∏ j ,r Ri ,r = (n i ,r −n i ,r )(k f ,r [C j ] ) arrangement with 78% of the coal slurry being fed into the 1st j=1 (30) stage section, and the remaining coal slurry is supplied to the = n − k f ,r ArT exp( Er / RT ) 2nd stage section. where ν ′ ,ν ′′ are the reactant and product stoichiometric i,r i,r Syngas coefficients, respectively, are the rate exponents for η′j ,r Geometry reactant j, in reaction r, n is the temperature exponent of •D1 = 2m reaction r, E is the activation energy, R is the universal gas r •D = 1.6m constant, Ar is the pre-exponential factor, and Cj is the molar 2 concentration of species j. •H = 12m D 2 2.6.2 Eddy-dissipation rate: •L = 8m • 3 H ( R ) ( P ) V = 42.24m Ri,r = min( Ri,r ,Ri,r ) ε ( R ) ′ YR Ri,r =n i,r M i Ar ( ) k n R′ ,r M R (31)

( P ) ε ∑P YP Ri,r =n i′′,r M i ABr Coal slurry k ∑Nn ′′ M j j,r j (No oxidant) Coal where YR and YP are the mass fractions of reactant and Coal produce species, respectively. A is the Magnussen constant for slurry slurry reactants (4.0), B is the Magnussen constant for products D1 (0.5), M is the molecular weight, and the R and P subscripts Oxidant are the reactants and products. Oxidant The global gasification reaction rates used in this study L are listed in Table 1 [24]. Fig. 3 Schematic of the computational model 5

The grid sensitivity study was conducted with f grids treated as gas phase fed together with the oxidant in the (0.12, 0.31, 0.49, and 1.05 million cells). The differences of model. The treatment of the ash transformation (i.e. the species between the cases of 0.31, 0.49M and 1.05M are transformation to slag) is neglected; instead, SiO2 is used as a less than 1%, and the difference of outlet temperature between gas species to represent ash in the model. For a real E-GAS the two cases is less than 3%. The detailed information was gasifier, there is a hopper at the bottom of the gasifier to documented in [5] and is not repeated here. To save discharge the slag. However, because of the assumption stated computational time, the grid containing 318,848 tetrahedral above, in this study, both the physical behavior of slag meshes is used with denser meshes being established near the formation and the hopper design at the bottom are included in region of the three injectors. The fluid residence time is the simulation. about 2~8 seconds, but most of the coal particles have reacted The total slurry mass flow rate is 39.7 kg/s, and the total within about 0.4 seconds, which seems shorter than the actual oxidant mass flow rate is 22.9 kg/s. The operating pressure is operating experience. 28 atm. Refractory brick is used in the real case of the E-Gas gasifier, but the properties of the material are not known. Therefore, in this study, the wall is assumed to be adiabatic. 3.2 Boundary Condition The mass ratio of O2/Coal (DAF) is 0.92, and the coal slurry concentration (defined as Mass of Coal (MF)/Mass of Coal Illinois #6 coal was used for the base case, and N.D. Slurry) is 0.67. The detailed inlet and boundary conditions and Lignite was chosen as the low-rank coal in this study for the corresponding parameter ratios of the base case are comparison. The properties of both are shown in Table 2 and summarized in Table 4. Table 3, respectively. It is worth notice that there is a “pre- dry” sub-column in the N.D. Lignite column. The inherent water content of lignite is usually too high to be gasified after Table 4 The feedstock conditions of base case it is mixed with water to make coal slurry. Usually the inherent water content in the coal does not help the coal slurry Flow rate (tons/day) making process because a certain amount of external water is - Coal (as received): 2545 needed for the coal particles to mix well with the water bath. - Slurry water: 831 Thus, the lignite needs to be dried to a level of 12-15%, so it - Coal slurry: 3376 can be mixed well and used in the gasifier. The “after dry” fractions in Tables 3 and 4 reflect this adjustment. First Stage Flow rate (tons/day) - Coal Slurry Feed: 2633 Table 2 Proximate analysis of feedstock coal - Oxidant: 1947

Second Stage Flow rate (tons/day) Proximate Illinois #6 Coal N.D. Lignite analysis - Coal Slurry Feed: 743 wet wet pre-dry (wt. %) Operating Condition: Moisture 11.12 33.3 12.00 - Pressure (atm): 28 Ash 9.7 7.4 9.76 - Inlet Oxidant Temp. (K): 411 Volatiles 34.99 29.08 38.37 - Inlet Feedstock Temp. (K): 450 Fixed Carbon 44.19 30.22 39.87 Operating Parameters: sum 100 100 100 - O2/Coal (DAF): 0.92 HHV (MJ/kg) 27.1 16.5 21.8 - Coal (MF)/Coal Slurry: 0.67 - O2/C: 1.14 Table 3 Ultimate analysis of feedstock coal - H2O/C: 0.69

Ultimate Illinois #6 Coal N.D. Lignite analysis (wt. %) wet wet pre-dry 4. RESULTS AND DISCUSSION Moisture 11.12 33.3 12.00 4.1 Base case C 63.75 42.22 55.70

H 4.5 3 3.96 The outlet syngas composition and temperature results N 1.25 0.67 0.88 are compared with the reference data from the NETL’s CFD Cl 0.29 - - modeling results [38], as shown in Table 5. Except H2O, which has a 5.85 percentage point difference, the differences O 6.88 12.675 16.72 of all other exit gas species are within one percentage point. S 2.51 0.735 0.97 The predicted exit temperature is about 200 K higher than the Ash 9.7 7.4 9.76 value from the NETL's report. sum 100 100 100.00

To simplify the computational model, the masses of N, Cl and S shown in the proximate analysis are lumped as ‘N2’ and 6

Table 5 Exit syngas temperature and compositions of base results in a large recirculation zone above the 2nd stage case injector. This slow-moving recirculation zone occupies a Parameters Base case Reference [38] portion of the gasifier like a blockage that reduces the effective area for the core flow to pass through. This leads to a CO 36.24 35.90 longer residence time for the flow trapped in the recirculation Mole H 22.23 22.90 fraction 2 zones but accelerates the core flow due to reduced effective (%) CO2 12.69 12.20 through-flow area. The gas species trapped in the recirculation H2O 18.05 23.90 zone seems to extend the mixing time of gas species, but the reduced area separates the accelerated upward-stream from Exit temperature (K) 1803 ~ 1600 the recirculation zone, and reduces the opportunity for species mixing. This issue of large recirculation will be rectified in the 4.2 Cold Flow versus Hot Flow next section by modifying the configuration of the 2nd stage injection. In summary, the cold flow pattern can provide a Considering cold flow study, either experimentally or coarse view of the scaled-down hot flow field by catching the computationally, has often been used as a fast first-step flow impingement in the first-stage of horizontal cylinder and preliminary study of flow pattern in a gasifier; it will be the large recirculation zone but with significantly reduced interesting to compare the cold flow pattern with the reactive flow velocity. hot flow pattern and provide an insight into how much the cold flow pattern can be trusted. This section provides such 4.3 Effect of Configuration of 2nd Stage Injection comparison and discussion. Figure 4 shows the comparison of velocity vectors in To eliminate the recirculation zone that resulted from the midplane cutting through the injectors and on several the single lateral injection, two dual-jets arrangements are cross-sectional planes. Generally speaking, the cold flow has a considered: one is two opposing injectors (or opposing-jet), similar overall flow pattern as the hot flow field. For example, and the other is two tangential injectors (or tangential-jet). The both have an accelerating flow field in the throat region, and a nd result in Fig. 5 indeed shows that the large recirculation region recirculation zone located above the 2 stage injector in the second stage of the original single jet arrangement is position. However, the flow speed of the cold flow is much significantly reduced in the case of opposing jets and almost slower than that of the hot flow. The difference in flow speed fully vanishes for the case of tangential jets is mainly caused by the change of fluid volume. For the Figure 6 shows the mass averaged cross-sectional surface reacting flow, the fluid volume in the 1st stage is significantly temperature along the gasifier height. It shows that both the increased due to the large amount of released water vapor and dual-jet arrangements produce lower average wall temperature volatiles. Furthermore, the flow volume also expands due to in the 2nd stage than the single jet arrangement, with the reduced density caused by increasing temperature. The effect opposing jets achieving the lowest wall temperature. Lower of gas volume expansion on local velocity is significantly temperature near the wall region is more favorable for the more pronounced in the throat region - the maximum speed gasifier because a lower wall temperature could reduce for the hot flow is about 31.6 m/s, which is an order damage to the refractory, which can reduce maintenance cost magnitude faster than 2.58 m/s of the cold flow. In addition, st as well as extend the life span of the refractory material. the velocity fields are somewhat different in the 1 stage section. For the cold flow, all un-reacted particles are included Table 6 Comparison of exit syngas and temperature results and considered in the computational domain until the iteration of particles tracking exceeds the assigned upper limit of Single jet Opposing Tangential iteration step. However, for the hot flow, the particles will be Parameters (Base case) jets jets reduced or even disappear because of the solid-gas reactions. 36.2 Therefore, the main reason leading to the difference between CO 36.24 36.4 the cold flow and hot flow will be the fluid temperature. The Mole 21.93 H2 22.23 22.23 cold flow appears as a bell shaped profile in the mid-plane in Fraction 12.77 the 1st stage section with a maximum flow speed in the (%) CO2 12.69 12.53 17.94 centerline. However, for the hot flow, the flow profile appears H2O 18.05 18.04 more uniform around the centerline. It is deduced that when 1793 the coal particles are burned and turned into high-temperature Exit Temp. (K) 1803 1791 Syngas gas; the gas expands fast, and the flow rushes upward from 10.54 10.59 10.54 the central region. Hence it leads to a more uniform velocity HHV (MJ/kg) in the central region in the horizontal section. For the CGE (% 81.39 81.72 81.38 cross-sectional view of velocity profile in the vertical uprising Normalized 1 0.91 0.69 region, both of the cold flow and hot flow show small vortices Injection Power at the height of 3.6 meters, and these vortices disappear at the height of 7 meters. Figure 7 shows the cross-sectional average of species Both velocity vectors show a large circulation above mole fraction along the gasifier height. The distribution of the 2nd stage injector. As previously explained in [5], this species mole fraction of the three configurations is close to large recirculation is not an ideal condition. Due to the nd each other. It seems that the different arrangement of the 2 strongly lateral injection, the upward stream passes through stage injection doesn’t affect much of the syngas production the throat area, deflects away from the 2nd stage injection, and even though the velocity fields are different. Because in the 7 real experimental case [40], there exists some unburned char power needed for the 2nd stage injection, it shows the at the exit, the char reaction rate in our present numerical advantage of employing the dual-jet design during the 2nd model seems too fast to adequately predict this phenomenon stage injection. About 9% of the power can be saved with of unburned char. To better capture the unburned char, the opposing-jets; almost 31% of the power can be saved with char reaction rate needs to be calibrated against the tangential-jets. The formula for the calculation of the needed experimental data in the future before the quench section. energy is: Energy needed (W) = Volume flow rate (m3/s) x Figure 8 shows the results of fluid pathlines and pressure difference between inlets and exit of gasifier (N/m2) cross-sectional temperature contours. Generally speaking, the The power of consumption is normalized with the opposing-jet arrangement produces a more uniform and original single-jet condition. The subtle difference in exit axisymmetric cross-sectional temperature distribution than syngas composition and heating values for three different that of a single-jet. cases could suggest that the current CFD model may Table 6 shows the results of the exit temperature and over-predict the reaction rates. Hence, experimental data species concentration and indicates that the differences are immediately at the exit of gasifier before cooling and subtle between these three cases. However, by examining the gas-cleaning are needed for calibrating this CFD model.

Velocity vector Temperature (K)

y= 7m

y=7m

y= 5m y=5m

y= 3.6m y=3.6m

x=3m x=2m x=1m x= 3m x= 2m x= 1m

(a) Cold flow (b) Reacting flow

Fig. 4 Comparison between cold flow and hot gasified gas flow fields.

(a) (b) Fig. 5 Gases velocity vector and temperature contour for the case of (a) opposing jets and (b) tangential jets.

12.6 12.6 11.6 11.6 Opposite jets 12.6 Tangent jets 10.6 10.6 Single jet 9.6 9.6 11.6 8.6 8.6

7.6 Opposite jets 7.6 Tangent jets 10.6 6.6 6.6 Single jet Gasifier height (m) Gasifier height (m) 5.6 5.6 9.6 4.6 4.6 3.6 3.6 Opposing jets 0.25 0.3 0.35 0.4 0.15 0.2 0.25 8.6 CO H2 Tangent jets 12.6 12.6 7.6 Single jet 11.6 11.6 Opposite jets 10.6 10.6 Tangent jets

Gasifier height (m) Single jet 9.6 9.6 6.6 8.6 8.6 Opposite jets 7.6 7.6 Tangent jets 5.6 6.6 Single jet 6.6 Gasifier height (m) Gasifier height (m) 5.6 5.6 4.6 4.6 4.6 1780 1800 1820 3.6 3.6 0.11 0.16 0.21 0.15 0.2 0.25 0.3 Temperature (K) CO2 H2O

Fig. 7 Comparison of the mass average cross-sectional Fig. 6 Comparison of the mass flow average cross-sectional species mole fraction along the gasifier height surface temperature distribution along the gasifier height

Temperature (K)

(a) single jet (b) twin tangential-jet (c) twin opposing-jet

Fig. 8 Close-up view of the cross section temperature and the pathline of the cases with different 2nd stage injection arrangements

4.4 Comparison of a Low-Rank Coal (North Dakota by the higher H2O/C ratio. In general, higher O2/C ratio leads Lignite) and a Bituminous Coal (Illinois #6) to higher exit temperature while higher H2O/C ratio tends to lower the exit temperature. The N.D Lignite case, which has a The comparison of a low-rank coal and a bituminous higher O2/C ratio but results in a lower exit temperature, could coal is done in this section. The Illinois #6 coal discussed in be due to higher water evaporation and more endothermic the base case is used in this section to represent the thermal cracking of higher volatiles content in the Lignite. bituminous coal; the tested low-rank coal is the North Dakota The modeled volatiles form of the Illinois #6 Coal is Lignite. This section is divided into two parts, and both have CH2.761O0.264, and its standard state enthalpy of formation, different reference conditions; the first part employs the same derived from the coal heating value and the proximate feedstock mass flow rate as used in the base case E-GAS analysis, is -1.53E+07 J/kmole. On the other hand, the gasifier; and the second part employs a fixed input heating modeled volatiles form of the N.D. Lignite is CH3O0.792, and value to the gasifier. Other than the differences stated above, its standard state enthalpy formation is -3.71E+08 J/kmole. all other operating conditions are the same; namely, the Because of the different values of enthalpy of formation and O2/Coal (DAF) ratio is 0.92 and the Coal (MF)/Coal Slurry thermal cracking products, the reaction heat that N.D. ratio is 0.67. The corresponding O2/C and H2O/C ratios are Lignite’s volatiles absorb during the thermal cracking process presented in Table 7. is higher than that of the Illinois #6 Coal, and therefore, by combining more water evaporation, it causes a temperature 4.4.1 Under the same feedstock flow rate condition drop at the exit for the Lignite case. The low-rank coal possesses a lower heating value. Table 7 shows the operating conditions, operating Therefore, under the same feedstock mass flow rate condition, parameters, exit temperature, and species concentration under the case using N.D. Lignite generates syngas with a lower the condition of fixed feedstock mass flow rate. To operate in heating value. Moreover, because of the low gasification the same O2/Coal ratio, the mass flow rate of oxidant for the temperature, there is some unburned carbon left at the exit, lignite case is reduced to 17.9 kg/s, and the corresponding which results in a slight decrease in carbon conversion (CC) O2/C and H2O/C ratios are also adjusted as presented in the and cold gas efficiency (CGE) as well. It will be interesting to table. As Table 7 shows, N.D. Lignite, with the same investigate the result with equal amount of total input heating feedstock flow rate, predicts lower CO and H2 mole fractions value by increasing lignite’s input flow rate and reducing the and higher CO2 and H2O mole fractions at the exit. The result O2/C ratio to the same value as the Illinois #6 Coal case, as of the higher CO2 concentration can be explained by a higher shown in next section. O2/C ratio, which contributes to more carbon oxidation reaction; the higher exit H2O concentration is clearly caused 10

Table 7 Comparison of low-rank coal and bituminous coal inherent water content, and the next case considers this coal (Illinois #6) under the condition of fixed feedstock flow rate pre-treatment condition. This investigation is conducted, firstly, by calculating Illinois the ultimate and proximate analysis of coal and the N. D. Lignite #6 Coal corresponding coal heating value after drying (Table 2 and Table 3). After obtaining the properties of the dried lignite, the Fixed Base Fixed feedstock and oxidant flow rate are modified to the same Operating Condition heating Pre-dry case flow rate O2/Coal and Coal/Coal Slurry ratios. The right column value (pre-dry) in Table 7 shows the results of this pre-treated lignite Feedstock rate (kg/s) 29.93 29.93 49.15 37.26 case under the fixed input heating value. Both CO and H2 Oxidant rate (kg/s) 22.9 17.9 29.40 28.65 mole fractions decease slightly; in addition, the values of Price of feedstock per CGE, CC, and syngas heating value reduce as well. The result 49.46 14.57 14.57 14.57 ton (USD/ton) indicates that modeling the drying of the water content inside Price of feedstock per the lignite leads to a lower gasification performance. 127,901 37,677 61,873 46,905 day (USD) Operating Parameters The present price of Illinois #6 Coal is not available, so the average price of bituminous coal from Illinois in 2010 is O2/C 1.14 1.35 1.35 1.31 used instead to represent its price. The price of N.D. Lignite is H2O/C 0.69 0.79 0.79 0.79 obtained in a similar way by referring to the average lignite Exit Results price in 2010 from North Dakota. Both of these references are from the report by the U.S. Energy Information CO 36.24 33.56 35.25 34.58 Administration [39] and are listed in Table 7. Although fixing Mole H2 22.23 20.90 19.89 19.69 the lignite flow rate as the bituminous coal flow rate leads to a Fraction (%) CO2 12.69 15.65 14.06 13.97 lower syngas HHV, the price per syngas HHV is much lower. Furthermore, the input feedstock flow rate would increase H2O 18.05 24.40 25.40 26.34 when fixing the lignite’s input heating value as the Illinois #6 Temperature (K) 1803 1258 1257 1190 coal; it means that the injectors’ configuration should be CGE (%) 81.39 73.26 74.37 72.30 modified to meet the raised flow rate, and the cost will CC (%) 100 99.93 99.51 96.06 therefore be increased. The feedstock price per unit syngas Syngas HHV (MJ/kg) 10.54 7.56 7.69 7.56 HHV for ignite is about US$1.19/GJ versus US$2.66/GJ for Illinois #6 coal. Since the feedstock price per unit syngas US$/GJ 2.66 1.21 1.19 1.09 energy is cheaper for lignites, gasifying lignites is attractive for electricity production. 4.4.2 Under the same input heating value condition 5. CONCLUSION The column under the “fixed heating value” in Table 7 shows the exit results of the N.D. Lignite case under a fixed The results predicted by the present CFD model are (same) input heating value as in the Illinois #6 coal case. To comparable to the NETL’s CFD prediction—all exit gas reach the same input heating value, the lignite's flow rate is species mole fractions are within 1 percentage point increased. The injectors are assumed to be suitable for the difference, except the H2O result which has a 5.85 percentage raised feedstock flow rate without modifying their point difference. The other results in this study are configuration. Under the same O2/Coal (0.97) and Coal/Coal summarized as follows: Slurry ratio (0.67), the exit syngas heating value and the overall gasification performance are not affected by the  The cold flow pattern can provide a coarse view of feedstock flow rate, even though the feedstock heating value scaled-down hot flow field by catching the flow is already raised to the same level as that of the Illinois #6 impingement in the first-stage of horizontal cylinder and Coal case. the large recirculation zone but with significantly Because the current CFD model cannot tell the reduced flow velocity. difference between these two types of water, the inherent water content inside lignite is treated as the water content in  The opposing and tangential dual-jet designs for 2nd the coal slurry. However, in the real case, the gasification stage injection eliminate the non-ideal large process involved with inherent water content inside the coal is recirculation above the 2nd stage injector in the original different from the additional water that is used to blend coal to single jet design. generate the coal slurry. In other words, the current CFD model has not included the diffusion process and the energy  The opposing-jet arrangement results in the lowest wall needed to expel water from the coal/lignite structure. So far, surface temperature while the tangential-jet has the only the latent heat (evaporation heat) of the H2O phase second lowest wall temperature and can save the most change is included in the model. In a real operation, the energy in fuel injection power. low-rank-coal is often practically dried to about 12% (wt.)

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