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Pulverized coal flame structures at elevated pressures. Part 1. Detailed operating conditions
Fuel 84 (2005) 1563–1574 www.fuelfirst.com
Pulverized coal flame structures at elevated pressures. Part 1. Detailed operating conditions Gui-su Liu, Stephen Niksa* Niksa Energy Associates, 1745 Terrace Drive, Belmont, CA 94002, USA Received 27 September 2004; received in revised form 1 February 2005; accepted 7 February 2005 Available online 14 March 2005
Abstract The tests and simulations in this study characterize the chemical structure of pressurized pulverized coal flames, particularly (1) how the O2 in simulated near-burner flame zones is apportioned among the various fuel components; and (2) the burner operating conditions and mechanisms that most strongly affect flame structure. CFD simulations resolved the structures of flames of a subbituminous and two hv bituminous coals for stoichiometric ratios (SR) from 0 to 1.8 for pressures from 1.0 to 3.0 MPa. The structures of all flames were largely determined by the accumulation of particles in the turbulent boundary layer on the flow tube wall. Gaseous fuel compounds always ignited first on the wall at the burner inlet, and this flame propagated toward the flow axis to form a 2D parabolic flame surface. Within the core, residual gaseous fuels, soot, and char may have eventually reached their ignition threshold and burned in a premixed mode. Residual CO, H2, and char burned in the near-wall region after the volatiles flame had propagated deeper into the core. Whether or not the flame closed on the centerline was mainly determined by pressure and SR. Inlet conditions that formed closed flames at a lower test pressure eventually sustained open flames at progressively higher pressures. The impact of decreasing SR was qualitatively similar, due to the lower heat release rates for progressively lower SR. As the pressure is increased, flame ignition and, by association, flame stability will become more problematic due to the greater thermal capacitance of air streams at progressively higher pressures. q 2005 Elsevier Ltd. All rights reserved. Keywords: Burner simulations; CFD; Flame structure; Pressure; Pulverized coal
1. Introduction Across the globe developers of coal-fired power generators face imperatives to raise conversion efficiencies to compete better with other fuels, especially where CO2 emissions are being reduced. A multitude of advanced process concepts are under development, as surveyed recently by Beer [1]. All have one thing in common: primary conversion of the coal feed at elevated pressure. Major blocks in the technical foundation supporting development of pressurized coal conversion technology are already in place. The impact of elevated pressure on the rates and yields of the coal conversion kinetics are clearly evident in extensive databases for devolatilization and char oxidation [2]. The database on char gasification needs to be * Corresponding author. Tel.:C1 650 654 3182; fax: C1 650 654 3179. E-mail address: [email protected] (S. Niksa).
0016-2361/$ - see front matter q 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.fuel.2005.02.005
expanded further [3], but only because so many operating conditions and char characteristics are important. Notwithstanding these essential elements of pressurized pulverized coal (p.c.) flames, we have not found any characterization work in the English literature on the structures of actual p.c. flames at elevated pressures. ‘Flame structure’ denotes the spatial sequence of distinctive physicochemical processes in a self-sustaining flame. It can only be ascertained from tests with realistic suspension loadings, because suspension loading determines the proportions of gaseous fuel compounds, soot, and char. The competition among these very different fuel components for the available O2 imposes a wide range of chemical time scales which determine flame structure, in conjunction with transport phenomena. The tests and simulations in this study characterize the chemical structure of pressurized p.c. flames, particularly (1) how the O2 in simulated near-burner flame zones is apportioned among the various fuel components; and (2) the burner operating conditions and mechanisms that most strongly affect flame structure.
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This work builds upon our previous work on the process chemistry in the near-burner region of diverse p.c. flames at atmospheric pressure [4,5]. A similar axisymmetric burner system was developed to operate at pressures from 1.0 to 3.0 MPa, and used to monitor complete product distributions from one subbituminous and two high volatile (hv) bituminous coals for burner stoichiometric ratios (SR) from zero to well above unity [6]. The test results are interpreted in two stages. First, in this paper, CFD simulations identify the mechanisms responsible for several surprising features of these flames, and quantitatively assign the flow and thermal fields. Then, in a companion paper [7], the CFD results are used to specify an equivalent reactor network for the flows. Separate simulations of detailed reaction mechanisms across the reactor networks accurately predict the product distributions with a minimum number of adjustable parameters, and identify the dominant conversion and NOX production mechanisms.
2. Test facility and evaluation data 2.1. Burner configuration The test facility simulated the thermal and chemical environments in the primary zone of a pulverized coal flame without the complications of two-phase turbulent mixing. As explained elsewhere in detail [6], the coal burner was a vertically mounted, cylindrical flow reactor that imposed uniform radial heat fluxes comparable to those in utility burners. An intense external radiant field stabilized
the reaction fronts so that the burner could operate with any inlet O2 level, including none at all in cases that determined the distributions of secondary pyrolysis products. As suspensions moved along the flow tube, they were heated at rates approaching 104 8C/s to the onset temperature for primary devolatilization, released their volatiles, and burned. At any particular operating condition, O2 depletion eventually ‘quenched’ the chemistry at an intermediate stage determined by the proportions of coal and O2 at the inlet. Inlet O2 levels were progressively increased in successive cases to move the process chemistry through oxidative volatiles pyrolysis, volatiles combustion, soot combustion, and char oxidation. The basic reactor configuration, called the ‘radiant coal flow reactor (RCFR)’, was used previously in studies of p.c. combustion at atmospheric pressure [4,5] and coal pyrolysis at elevated pressures [8,9]. The version of the burner for elevated pressures is labeled as the ‘p-RCFR’. The main difference is that the particle flow is downward at atmospheric pressure and upward at elevated pressures, to counteract recirculation induced by buoyancy. The system is sketched in Fig. 1 and described elsewhere [6] in more detail. A feeder dumped classified p.c. particles into an O2/Ar entrainment stream, forming an optically thin suspension that flowed downward into a U-tube. The other end of the U-tube delivered the upward, two-phase flow into the radiant furnace section within the pressure vessel. A 27 cm mullite tube with 1.2 cm i.d. was located on the central axis of an insulated graphite cylinder wrapped by five turns of a water-cooled copper induction coil. The graphite length was 15.8 cm, although the overall hot zone length was
Fig. 1. Schematic of the p-RCFR experimental facility.
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roughly twice as long due to radiation leakage out of the ends of the entrainment tube. The graphite temperature was regulated to 1650 8C. The inlet gas velocities were 29.5 cm/s in all tests, although residence times were significantly variable due to different extents of heat release among the various test conditions. At the inlet, an annular sheath flow surrounded the entrained suspension to help prevent particles from colliding with the tube wall. The coal particles heated and expelled their volatiles which, in the presence of O2, burned along with the char particles. During combustion tests, the heat released during the combustion of volatiles and char supplemented the heating of the entrainment gas by convection from the flow tube and coal suspension. Immediately downstream of the burner, Ar was injected into the flow to quench the process chemistry, nucleate any residual tar into an aerosol, and deliver the char particles, aerosol products, and gases into the product recovery and analysis section. A cylindrical ‘soot filter’ captured soot and tar, which was subsequently extracted from the filters with tetrahydrofuran, and filtered. The solids captured on the membrane were denoted as soot, and the dissolved material was condensed out and denoted as tar. The char yield was assigned as the amount of material in a char collection basket, plus the amount of large particles taken from the first layer of the soot filter, plus any residual solids recovered from the flow system and burner flow tube. All condensible materials were analyzed for elemental composition and ash content. Concentrations of CO, HCN, NH3, and NO were monitored on-line with FTIR spectroscopy; C1–C4 hydrocarbons (CH4, C2H2CC2H4, C2H6, C3H6, C3H8, C4) and H2 were extracted for subsequent chromatography into multiport sampling valves; moisture and CO2 were monitored with separate NDIR detectors. 2.2. Test conditions Properties of the Wyodak subbituminous (PRB) and two hv bituminous coals used in the tests appear in Table 1. Note the extremely high sulfur contents of the Ill. #6 and Pit. #8 samples and the high ash content of the Ill. #6, which are characteristic of raw, as-mined samples. These coals were obtained from the Penn State Database, then aerodynamically classified. The tested samples were a mixture of two Table 1 Coal properties Coal name Pit. #8 Ill. #6 PRB
Proximate analysis (ar wt%) Ultimate analysis (daf wt%) M
Ash
VM
FC
C
H
O*
N
S
0.7 0.2 0.1
12.3 17.3 5.0
37.9 35.8 39.4
49.1 46.7 55.5
80.8 74.1 73.7
5.4 5.5 5.6
5.8 8.2 19.0
1.7 1.4 1.1
6.3 10.8 0.6
*Assigned by difference.
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sieve sizes, 75–90 mm and from 90 to 105 mm, so the mean size was about 90 mm. The test plan comprised six test series that characterized Pit. #8 at 1.0, 2.0, and 3.0 MPa, Ill. #6 at 1.0 and 2.0 MPa, and PRB at 1.0 MPa only. Each test series contains from seven to ten individual tests with progressively higher inlet O2 levels, hence, SR values. In the tests with Pit. #8 at 1.0 MPa, suspension loadings were nearly uniform at 4.7 wt%, whereas the inlet O2 mass fraction was varied from zero for the secondary pyrolysis case to 9.9%, so that SR was varied from 0 to 0.953 in fairly uniform increments. The suspension loadings were decreased from 4.7% at 1.0 MPa, to 2.3–2.5% at 2.0 MPa, to 1.55% at 3.0 MPa; in other words, coal feedrates were essentially the same at all test pressures. Inlet O2 mass fractions were regulated at the higher pressures to impose similar ranges of SR values in all test series. The maximum SR values were near-unity with Pit. #8; 1.77 with Ill. #6; and 1.27 with PRB. 2.3. Representative results The datasets on PRB coal at 1.0 MPa in Table 2 illustrate the measurement uncertainties and main tendencies in the results. The measured product distributions are arranged in order of increasing SR from left to right, beginning with the case for secondary pyrolysis. Consequently, cases toward the left are dominated by the products of primary devolatilization, secondary volatiles pyrolysis, and partial oxidation of volatiles; whereas those toward the right are governed by water gas shift equilibrium and the conversion of soot and char. When the SR exceeded one-half, all hydrocarbon fuel components had been consumed, but large amounts of CO and H2 persisted. Due to the thermal severity of all tests, tar was not even observed in the case for secondary volatiles pyrolysis alone. Soot yields decayed somewhat erratically from 9.1 daf wt%. The chars were fully ignited for all SR over 0.15, and extents of char burnout increased continuously for progressively higher SR, as expected. Mass balances on individual tests closed within G2% for all cases, while C-balances closed within G4%; H-balances within G5%; and O-balances within G2%. The only adjustments to the product distributions were for the S-species, which were not monitored. We arbitrarily expressed the volatile-S species as H2S under conditions that were sufficiently reducing to retain hydrocarbon fuel compounds, and as SO2 under conditions where essentially all the hydrocarbon fuel compounds had burned. This procedure would be expected to overestimate the volatile-S species under the most reducing conditions, because we have no evidence that all the coal-S had actually been released into the gas phase under the relatively moderate temperatures in these tests. The product distributions indicate that O2 is not apportioned sequentially among the various fuel components during the initial stages of coal combustion; clearly,
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Table 2 PRB, 1.0 MPa datasets (daf wt%) SR
0
Product distribution 6.6 CO2 H2 O 8.1 CO 12.3 CH4 2.9 C2 4.0 C3 0.4 Oils – H2 2.75 HCN 0.69 NH3 0.07 NO 0.00 Tar – H2 S 0.64 – SO2 Soot 9.1 Char 52.4 Mass balances SMass 1.000 SC 1.000 SH 0.975 SO 1.002 Burnout XHC 0 XSoot 0 0 XChar
0.028 6.6 8.1 17.1 2.9 4.0 0.3 1.8 2.75 0.69 0.07 0.00 – 0.64 – 9.1 52.4 1.016 1.021 0.997 1.000 2.0 0 0
0.121 18.9 15.9 30.6 2.4 2.2 0.1 0.9 2.41 .61 0.09 0.00 – 0.64 – 3.8 49.9 1.015 1.028 1.002 0.995 35.6 58.2 4.8
0.241 36.9 23.6 41.5 1.5 1.0 0.0 0.1 2.06 0.57 0.13 0.00 – 0.64 – 1.5 42.8 1.002 0.998 0.961 1.009 64.5 83.5 18.3
0.404 61.1 27.5 50.8 0.6 0.3 0.0 0.0 1.92 0.45 0.17 0.00 – 0.64 – 1.6 34.7 1.001 1.016 0.991 0.992 85.0 82.4 33.8
soot and char burn along with gaseous fuels from the PRB in Table 2 and also from both hv bituminous coals (as seen in Figs. 2 and 3, below). As introduced previously [5], oxygen consumption is evaluated quantitatively in terms of the carbon consumption of three major fuel components: hydrocarbon gases (CH4, C2H2, oils, and HCN), soot, and char. Carbon burnout indices for each component, CXi, are defined as C
X n 100 Xi Z C %Cð0Þ ðII W j K Meas W j ÞC fj fi jZ1
(1)
where IIWj is the yield of component j for secondary pyrolysis only; MeasWj is the measured yield of component j in a combustion test; Cfi is the carbon fraction in component i; n is the number of compounds in the fuel lump under consideration; and %C(0) is the coal-C. As seen in Table 2, all the gaseous fuel compounds were burned out when SR approached 0.6, but only about three-fourths of the soot and just over half of the char were converted at this condition. The same tendencies are evident in the datasets for both hv bituminous coals. The most significant differences are the much higher soot yields for secondary volatiles pyrolysis of 21–23 daf wt% for both coals, and the tendency for fewer gaseous hydrocarbons from coals of progressively higher rank. Char yields were similar at all pressures, because even the lowest test pressure was higher than the threshold for asymptotic total weight loss for primary
0.578 90.4 32.0 49.8 0.1 0.0 0.0 0.0 1.56 0.19 0.11 0.00 – – 1.20 3.1 24.0 0.996 0.985 0.962 1.009 97.3 65.9 54.2
0.751 126.1 36.1 43.3 0.0 0.0 0.0 0.0 1.20 0.05 0.06 0.00 – – 1.20 3.3 18.3 1.003 1.004 0.964 1.008 99.2 63.7 65.1
0.934 166.4 39.4 37.4 0.0 0.0 0.0 0.0 0.82 0.05 0.05 0.00 – – 1.20 4.1 11.5 1.001 1.039 0.954 0.991 99.5 55.0 78.1
1.092 192.6 43.8 31.3 0.0 0.0 0.0 0.0 0.57 0.00 0.02 0.08 – – 1.20 1.7 6.9 0.992 1.007 0.984 0.991 99.9 81.3 86.8
1.272 203.1 45.9 26.4 0.0 0.0 0.0 0.0 0.34 0.00 0.01 0.11 – – 1.20 2.2 4.4 1.003 0.991 0.981 1.012 100.0 75.8 91.6
devolatilization. All datasets have been reported separately [6].
3. CFD simulations CFD simulations were prepared with v. 5.5 of Fluent to assign 2D axisymmetric flow, temperature, and species concentration fields and particle trajectories for each test. Although several essential features became apparent in the CFD results, the original intent was to assign accurate thermal histories for subsequent simulations with detailed chemistry. Accordingly, various model parameters were freely adjusted to match the predicted and reported extents of consumption of volatiles, soot, and char for each test. This procedure ensured that the simulated heat release profiles were accurate, which is important because the heat release profiles are the most uncertain aspect of heat transfer in the burner, by far. The thermal analysis proceeded in two stages, as described previously [10]. First, the radiant energy transfer within the burner was analyzed to assign the temperature profile along the flow tube and the radiant flux onto the flow axis. These profiles were incorporated into the CFD simulation as boundary conditions. Convection from the flow tube through a developing boundary layer in either transitional or turbulent flow was also included. Radiant fluxes onto the centerline of the 15.8 cm radiant section ranged from 61 W/cm2 at 3.0 MPa to 66 W/cm2 at 1.0 MPa,
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Fig. 2. Burnout profiles imposed in the CFD simulations for gaseous fuels (C and solid curve), soot (B and dashed curve), and char (- and dotted curve) for Pit. #8 at 1.0 (upper), 2.0 (middle), and 3.0 MPa (lower).
Fig. 3. Burnout profiles imposed in the CFD simulations for gaseous fuels (C and solid curve), soot (B and dashed curve), and char (- and dotted curve) for Ill. #6 at 1.0 (upper) and 2.0 MPa (middle), and for PRB at 1.0 MPa (lower).
and were within the range of 50–100 W/cm2 estimated for large p. f. flames. All simulations began 6 cm upstream of the heating element, consistent with the calculated radiation leakage into the upstream section of the flow tube. CFD simulations were based on conservation equations for mass, momentum, and energy transfer in the gas phase for steady, 2D, axisymmetric flow. Core and sheath flows had equal flowrates but were distinguished at the inlet with a uniform particle suspension loading in the core and no particles in the sheath. The temperature and flow velocity at the inlet were uniform for the sheath flow, but a fully developed velocity profile was applied to the core flow, based on separate simulations of the inlet section. Transport coefficients for heat and momentum transfer to the tube wall were assigned for transitional boundary layers
at 1.0 MPa and for fully developed turbulent flow at the higher test pressures. Based on our own evaluations of the various options for the turbulence submodels and near-wall treatments, the k–3 turbulence submodel and the two-layer zonal submodel were implemented. This combination was the only one that accurately reproduced measured axial velocity profiles [11] as well as the turbulence intensity profiles across the boundary layers on vertical tubes from large-eddy simulations (LES) [12]. In conjunction with the Stochastic particle dispersion submodel evaluated with 500 particle trajectories, the turbulent flow analysis was also qualitatively consistent with the particle concentration profiles in similar flowfields from LES [12]. Turbulence intensities at the burner inlet were unimportant influences on the downstream flowfield.
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Lagrangian trajectories of individual particles were evaluated in the Stokes limit, consistent with the nearunity Reynolds numbers of small p.c. particles. Particle thermal histories accounted for absorption of the incident radiant heat flux profile balanced against the particle’s thermal capacitance, convection to the gas phase, and the heat and mass release due to devolatilization and char oxidation. Submodels for coal conversion included composition- and temperature-dependent thermophysical properties, a single first-order devolatilization reaction, a swelling factor, and char oxidation according to the classical balance between an intrinsic nth-order surface reaction and film diffusion of O2. The following chemistry submodel was incorporated: Coal/ GasVol C Soot C Char
(2)
GasVol C gO2 O2 / gH2 H2 C gCO CO
(3)
CO C 0:5O2 / CO2
(4)
H2 C 0:5O2 / H2 O
(5)
Soot C 0:5O2 / 2CO
(6)
Char C 0:5O2 / CO
(7)
where GasVol denotes the aggregate lump of CH4, C2H2, HCN, and oils. A single first-order reaction represents devolatilization kinetics, as follows: K
dmp Z kdv ðmp K ð1 K fv0 Þmp0 Þ dt
(8)
where the devolatilization rate is kdv Z Adv expðKEdv =RTp Þ and fv0 is the ultimate volatile fraction on a dry basis. The rate parameters and ultimate yield were assigned from simulations with FLASHCHAINw [13] based on the coal’s proximate and ultimate analyses, and the thermal histories from the CFD simulations. We first estimated the thermal history, then assigned rate parameters for the FLUENT simulation. The updated thermal histories from the FLUENT simulation were used to refine the rate parameters, until the rate parameters did not change. Coal swelling was represented with a swelling factor correlation. Since the primary volatiles are injected into hot gases in the p-RCFR, they are assumed to be instantaneously converted into secondary volatiles pyrolysis products, whose yields were also assigned by FLASHCHAINw. The oxidation of the gaseous fuel components proceeds in two stages: (1) partial oxidiation of hydrocarbons into CO and H2 (Eq. (3)); and (2) CO and H2 oxidation (Eqs. (4) and (5), respectively). One-step schemes are inconsistent with the rapid increase in the CO levels for progressively higher SR values (cf. Table 2). The stoichiometric coefficients in Eq. (3) were evaluated from the complete product distribution of gaseous volatiles, and vary with both coal rank and pressure. The burning rates of volatiles are first
order in both the fuel species and O2. The same global activation energy was applied to each fuel, based on the analysis of Cho and Niksa [14], but the pre-exponential factors were adjusted to match the burnout index for gaseous fuels for each test. The same rate parameters were applied to CO and H2, assuming that CO oxidation kinetics limit the equilibration of the water–gas shift chemistry. At 0.8 mm, soot agglomerates are sufficiently small that soot oxidation can be regarded as a pseudo-gas phase reaction that generates only CO. Based on the soot oxidation data of Park and Appleton [15] below 1700 8C, the soot burning rate is zero-order with respect to O2 with an activation energy of 140.5 kJ/mol. Soot diffusivities were specified by Lau and Niksa [16] as functions of soot size and gas temperature. A typical value of the Brownian diffusivity of 0.12 mm soot at 1730 8C and at atmospheric pressure is 0.57 cm2/s, versus a thermophoretic diffusivity of 0.06 cm2/s. The overall char burning rate reflects the sequential processes of film diffusion of O2 and intrinsic global kinetics, where the order of the surface kinetics is one-half. The activation energy was assigned from correlations in terms of the daf carbon content of the parent coal that were established from a database of flow reactor tests [17], whereas the pre-exponential factors were fit to match the reported char burnout profiles. Both particle density and diameter changed during char combustion, according to the empirical relation between mass and density developed by Hurt and Mitchell [17]. Hence, four pre-exponential factors in the burning rates for gaseous fuels, CO and H2, soot, and char were adjusted to match the burnout profiles of gaseous fuels, soot, and char and the total O2 utilization efficiency for each test.
4. Results 4.1. Burnout Indices Indices on the burnout of gaseous fuels, soot, and char for the tests with Pit #8 and Ill. #6 and PRB appear in Figs. 2 and 3, respectively. The curves in the figures show the results from the CFD simulations, which closely match the measured values by virtue of adjustments to the four rate parameters discussed in Section 3. The burnout profiles for gaseous volatiles from Pit. #8 at the three test pressures agree within experimental uncertainty. This does not imply that the burning rate and combustion mechanisms were independent of pressure, because the profiles are plotted versus SR rather than a time-based coordinate. Moreover, the gas temperature history reaches progressively cooler maxima at higher pressures, as discussed in Section 4.2.2. At all pressures, the indices approach the asymptotic value of 100% for an SR of 0.6. The soot burnout profiles for Pit. #8 for 2.0 and 3.0 MPa are essentially the same, but have
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higher extents of burnout than the profile for 1.0 MPa at SR values between 0.2 and 0.7. The strongest impact of pressure is seen in the char burnout profiles in Fig. 2. At a SR value of 0.95, the extents of char burnout of Pit. #8 decrease from 57.5 daf wt% at 1.0 MPa to 52.5% at 2.0 MPa to 27% at 3.0 MPa. The impact of pressure is also apparent at all lower SR values in these datasets. In fact, the O2 partial pressures were almost the same for progressively higher pressures because the same coal feed rate was used at all pressures and the O2 flowrates were the same to maintain the same SR. The diminished char burnout is mostly due to the lower particle temperatures at higher pressures, which will be further discussed below. The burnout profiles for Ill. #6 and PRB appear in Fig. 3. Unlike those for Pit. #8, the gas burnout profile for Ill. #6 at 1.0 MPa is higher than at 2.0 MPa. Soot burnout, however, is similar at both pressures. Extents of char burnout at 1.0 MPa are about 10% higher than that at 2.0 MPa across the entire SR range, consistent with the Pit. #8 char burnout profiles. In fact, the soot burnout profiles at both pressures and the gas burnout profile at 2.0 MPa for Ill. #6 are quantitatively consistent with the soot and gas burnout profiles for Pit. #8. Only the gas burnout profile at 1.0 MPa is different. The burnout indices for PRB at 1.0 MPa appear in the bottom panel in Fig. 3. The indices for gaseous fuels are essentially the same as those for Pit. #8 at 1.0 MPa, and those for soot burnout are similar to those from Ill. #6 soot at 1.0 MPa. Soot probably burns at a rate that is independent of coal rank, because it contains mostly carbon (O98.5%) with small amounts of hydrogen and nitrogen. The compositions of soot from all coal types are nearly the same. Unlike the burnout of gaseous volatiles and soot, char burnout is strongly affected by coal rank, as expected. At 1.0 MPa and a SR value of 0.95, the char burnout for Pit. #8 is 57.5 daf wt%; 60% for Ill. #8; and 77% for PRB. PRB has the fastest char burning rate, and the extent of PRB char burnout even overtakes soot burnout at a SR value of 0.85. The impact of coal rank is consistent with extensive databases on char oxidation at both atmospheric pressure [17] and elevated pressures [2]. Regardless of the pressure and coal type, gaseous volatiles combustion consumes most of the available O2 at low SR. Soot effectively competes for O2 at low SR, but burnout of the chars of the lowest rank eventually overtakes soot burnout at SR over 0.8. This char presumably burns faster than soot because chars from low rank coals usually contain substantial levels of alkaline earth and alkali cations that promote char oxidation, whereas soot does not. There are master burnout profiles for the gaseous volatiles and soot from all fuels and all pressures, albeit, each has an exception at one operating condition. But char burnout profiles are strong functions of coal quality and pressure.
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4.2. Baseline flame structure The burnout indices exhibit the expected sequence of fuel consumption from gaseous fuels to soot to char, with well established dependencies on coal rank and pressure. Considering the burner’s highly idealized geometry and inlet conditions, one could anticipate planar, nominally 1D flames across the central region of the flow tube, where the particle concentrations are highest. But such a simple structure is contradicted by the co-existence of substantial amounts of O2, CO, and H2 in the products from every test, except those for secondary volatiles pyrolysis. Typically, only 70% of the inlet O2 was consumed while about 30 daf wt% CO remained in the products, even for SR well below or well above unity. Hindered penetration of O2 from the sheath flow into the core is an obvious explanation that turns out to be incorrect. The key is the characterization of highly nonuniform particle dispersion in Section 4.2.1. The generic aspects of flame structure will be illustrated with the CFD simulations for Pit. #8 at 1.0 MPa with a SR of 0.95, a ‘baseline’ case, before the trends with SR, pressure, and coal quality are presented. 4.2.1. Particle dispersion patterns The stochastic submodel for particle dispersion in the CFD simulations imparts significant fluctuations in the particle motion. Almost immediately after injection, the particles acquire significant radial velocity components due to the turbulence and the wall collisions. All particles eventually penetrate into the sheath, and almost all of them contact the wall at some point. Once the particles move into the boundary layer, they are unlikely to escape back into the core flow, so particles accumulate in the boundary layer and deplete the suspension loading in the core. Fig. 4 shows the radial profiles of normalized particle number concentration at six cross sections. At the inlet, the particle concentration is uniform across the core, which extends to 0.707 of the tube radius (or 0.00424 m). At 1 cm downstream of the inlet, particles are still concentrated in the core, but a concentration gradient extends into the sheath. At 4 cm, the particle concentration profile has been inverted by dispersion into the sheath. Particle concentrations in the near-wall region continuously increase throughout the remainder of the burner, while the concentrations in the core diminish. Hardly any particles remain on the flow axis at the burner outlet. This result bears two important implications for the performance of the p-RCFR that were confirmed in the laboratory. First, particle agglomeration near and on the walls will likely cause operational problems, due to the combination of slow particle velocities, high particle temperatures, and high particle concentrations in the nearwall region. Second, complete O2 utilization will be difficult to achieve, because of the strong propensity for accumulation of all fuels in the near-wall region versus substantial amounts of residual O2 in the core.
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Fig. 4. Radial profiles of particle number concentration of Pit. #8 at 1.0 MPa and SR of 0.95 at axial positions of 1, 4, 6, 12, 18 and 27 cm.
4.2.2. Flow, temperature, and species concentration fields Fig. 5 shows radial profiles of gas velocity and temperature, and of mass fractions of gaseous fuels, soot CO and oxygen at six axial positions. The position labeled as zero is actually 5.8 cm upstream of the inlet to the radiant section, and the radiant section ends at 20.6 cm. The regions before and after the hot zone are affected by the radiant heat flux out the tube ends, as noted previously. The flow accelerates continuously as it moves through the tube, reaching a maximum velocity of 240 cm/s at the outlet. Due to buoyancy effects, the gas velocity reaches local maximum values near the tube wall as long as the flow is within the burner hot zone. The velocity profile becomes parabolic at 24 and 27 cm. A thermal boundary layer propagates into the core from the wall, at about 1650 8C. The centerline gas temperature rises continuously throughout the burner even while, downstream of the hot zone, the temperature profiles are inverted by convective cooling into a cooler tube wall. For this most oxidizing case in the Pit. #8 test series, the maximum gas temperature exceeds 1500 8C.
The rapid gas heating is primarily responsible for the rapid acceleration of the axial gas velocity. Gaseous volatiles and soot ignite at the inlet to the radiant section. Note their very high concentrations near the wall, where the gas temperature is hottest over this portion of the flow reactor. Particles are heated mainly by the radiant flux from the tube, but they are also heated by convection from the hotter gases near the wall. Consequently, the particles dispersed into the near-wall region devolatilize before those in the core, and the gaseous fuel concentrations are correspondingly higher near the walls. By 12 cm, the concentration spikes have been eliminated by combustion, and the highest gaseous fuel concentrations are within the core. The fuel inventory is then depleted by a flame front propagating from the near-wall region toward the flow axis. The flame is fed by diffusion of both fuel compounds and O2 from the core into the sheath flow. Since the gas temperature at 18 cm is below the threshold for ignition, the core fuel concentration is reduced by transport at this axial position. But by 24 cm, the core is fully ignited and the fuel is
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Fig. 5. Simulations of (a) axial gas velocity, (b) gas temperature, (c) gaseous volatiles mass fraction, (d) soot mass fraction, (e) CO mass fraction, and (f) O2 mass fraction for Pit. #8 at 1.0 MPa and a SR of 0.95.
depleted much faster. Almost all the gaseous fuels and soot were consumed before the reactor outlet in this particular run. The CO profiles are considerably more complex than those for the volatiles-derived fuels, because CO is generated as a product of primary devolatilization and secondary volatiles pyrolysis, as well as by partial oxidation of gaseous fuels, soot, and char. At 6 cm, there is no CO in the core, where the gases remain below the onset temperature for devolatilization. But in the near-wall region, the CO profile exhibits the spike seen in the gaseous fuel profile. Further downstream, the CO level in the core grows mostly through contributions from char oxidation through 24 cm, before it finally diminishes at the reactor outlet. But within the near-wall region, CO accumulates at roughly 1.2 wt% along the entire flow tube, due to depletion of nearwall O2 and diffusion of CO from the higher concentration in the core beyond 18 cm. The final oxidation of CO and H2 is relatively slow to begin with, and decelerated further by O2 depletion. The O2 profiles fall continuously from the inlet concentration, beginning at 6 cm in the near-wall region. Diffusion of O2 from the core into the sheath reduces the core O2 level by 12 cm. By 18 cm, the near-wall O2 level is below the threshold for gaseous volatiles combustion. Further downstream, the core O2 levels diminish more rapidly once the char suspension has ignited. Overall, 81.4%
of the O2 was consumed during this run. Not all the O2 was consumed in the near-wall region, but nearly all the residual O2 in the products came from the core flow. Hence, CO and O2 (and H2) co-exist in the products because the delayed ignition and relatively lean mixtures in the core preclude complete combustion. 4.2.3. Particle combustion histories Particle thermal histories were assigned as massweighted average values for the full population of particle trajectories in the CFD simulation. They begin to heat rapidly 150 ms from the inlet, at a nominal rate of 7300 8C/s, and were driven to 1410 8C at 380 ms. Thereafter, the particles cool at nearly the same nominal rate as they were heated. In light of the steep gradients in temperature, O2 concentration, and particle concentration in the CFD simulation, it is not surprising that the particle residence time distribution (RTD) is broad and non-normal. Almost 80% of the particles have residence times between 330 and 480 ms, but the maximum time is 790 ms. The RTD has the form of a gamma distribution and resembles the RTDs for one or a few stirred tanks in series. Whereas there are no short-circuits in this flowfield, the relatively few particles that become trapped in the near-wall region have significantly longer residence times than those remaining in
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the bulk flow. The mean residence time is 450 ms with a standard deviation of 70 ms. The particle RTD is only one of the factors responsible for the broad distributions on the extents of char burnout from this burner. For SR values less than 0.5, the burnout distributions are normal distributions with higher dispersions for progressively higher SR values. The mean values also shift toward higher values, as expected. But at SR values of 0.72 and 0.95, the burnout distributions became much more dispersed, and lost their Gaussian form, for two reasons: First, the particles in the core fully ignite and contribute to the highest bins in the burnout distribution only at the higher SR. Second, burning particles in the sheath were extinguished by O2 depletion, which freezes about half the population. 4.3. Impact of operating conditions The CFD flame structures for all other tests are qualitatively similar to the baseline case primarily because the accumulation of particles in a boundary layer on the reactor wall is insensitive to SR, pressure, and coal quality. Progressively increasing SR increases the maximum gas temperatures near the burner outlet, especially in the core. At the outlet, centerline gas temperatures were 550 8C hotter for a SR of 0.95 than for 0.15. But the maximum mean particle temperature increased by only 100 8C over this same range, because most particles are confined to the wall layer, not the core. These large variations in gas temperature with SR affect the flame structure, as seen in Fig. 6. For purposes of illustration, these fronts were arbitrarily located on the locus of positions where gas temperature is 1050 8C, which is hot enough to ignite all fuel components. As the flow moves through the tube, a flame front propagates toward the flow axis, driven by convective heat transfer from the wall and by
Fig. 6. Flame structures for Pit. #8 combustion at 1.0 MPa and at SR of (a) 0.95; (b) 0.72; (c); 0.51; (d) 0.37; (e); 0.25; and (f) 0.15. The position of the flame front is the locus of positions where the gas temperature is 1050 8C.
the heat release from combustion of gaseous volatiles and soot. Hence, the flame is sustained by outward diffusion of volatiles and O2 toward the wall, and by inward heat transfer toward the center. For some conditions, the annular flame front closes to a point on the flow axis; for others, closure cannot occur before the downstream edge of the radiant section. The sketch of this flame structure shares elements in common with both premixed Bunsen flames and laminar (Burke–Schumann) diffusion flames. But confined pressurized p.c. flames differ from both archetypes, because fuel consumption is not restricted to the volatiles flame. This flame segregates the flow according to the following three stages of combustion: (1) within the core, residual gaseous fuels, soot, and char may eventually reach their ignition threshold and burn in a premixed mode; (2) gaseous volatile fuels and soot sustain the volatiles flame as it propagates from the near-wall region toward the flow axis; and (3) residual CO, H2, and char burns in the nearwall region after the volatiles flame has propagated deeper into the core. Whether or not the flame closes on the centerline in the available residence time will be mainly determined by SR and pressure. Conditions that sustain a closed flame at a higher SR will eventually sustain open flames at progressively lower SR, which occurs at a SR of 0.15 with Pit. #8 at 1.0 MPa in Fig. 6. Comparable SR thresholds were 0.5 with PRB and 0.8 with Ill. #6. The impact of increasing pressure is similar. Inlet conditions that form closed flames at a lower test pressure will eventually sustain open flames at progressively higher pressures. In fact, the calculated maximum core gas temperatures at the burner outlet at 2.0 and 3.0 MPa with Pit. #8 were only 875 and 630 8C, which are 650 and 850 8C cooler than at 1.0 MPa. The maximum mean particle temperatures fell by 320 8C over the same pressure range. Consequently, neither flame at the higher test pressures closed across the centerline. None of the flames were closed for any coal at 2.0 and 3.0 MPa for SR values near unity. The impact of coal quality is somewhat more complex. Ill. #6 generates the coolest gas temperature profile, and PRB and Pit. #8 generate hotter and very similar profiles. At the burner outlet, the Pit. #8 core burns at 1500 8C; the PRB core burns at 1310 8C; and the Ill. #6 core burns at only at 1080 8C. But the maximum mean particle temperatures are within 100 8C, although PRB devolatilizes and ignites faster than both other coals. PRB char also has the fastest intrinsic oxidation reactivity, so it ignites at the lowest temperature and burns the fastest among these three coals. Consequently, the ultimate extents of burnout at 1.0 MPa for comparable SR were 57.9 daf wt%, 66.8 and 78.1% for Pit. #8, Ill. #6 and PRB, respectively. It may seem surprising that a coal with an intermediate char oxidation reactivity generates the most open flame structure in this comparison. The reason is the lost heat release associated with a large portion of unburned CO and
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H2 in the post-flame gases from Ill. #6. Only 50% of total weight loss from the Ill. #6 was converted into CO2 and H2O, and most of the residual fuel components were present as unburnt CO and H2 in the products. The comparable percentages were 67% for Pit. #8 and 73% for PRB. Consequently, the total heat release within the flow tube was lowest with Ill. #6, which hindered this coal’s flame stabilization.
5. Discussion Across the wide domain of test conditions in this study, the burnout indices of the major fuel groups show the expected order of consumption from gaseous volatiles to soot to char. All the indices increase for progressively greater SR, as expected. Those for both gaseous fuels and soot were insensitive to variations in both pressure and coal rank, albeit with exceptions. Soot effectively competes for the available O2 at low SR but gases win the competition under more oxidizing conditions. The extents of char burnout uniformly diminished for progressively higher pressures, due to the cooler gas temperatures that inhibited char ignition at elevated test pressures. The chars from coals of lower rank also burned faster, as expected. Based on such behavior, one could reasonably have expected that the upward flowing, sheathed cylindrical p.c. suspensions in the p-RCFR closely approached the idealized structures of planar, 1D flames. But this simple view is contradicted by the co-existence of O2, CO, and H2 in the products of every test that had O2 at the burner inlet. It was also superceded by the complex flame structures in CFD simulations, which are largely determined by the accumulation of particles in the turbulent boundary layer on the flow tube wall. At another superficial level, these flame structures resemble both premixed Bunsen flames and laminar diffusion flames, primarily because gaseous volatiles always ignite first on the wall at the burner inlet, and this flame propagates toward the core to form a 2D parabolic flame surface. This flame only marks the penetration of the thermal layer toward the centerline. It does not indicate the major fuel consumption patterns, because major portions of the combustibles burn both within the volatiles flame and beyond it. Within the core, residual gaseous fuels, soot, and char may eventually reach their ignition threshold and burn in a premixed mode. Residual CO, H2, and char burn in the near-wall region after the volatiles flame has propagated deeper into the core as long as O2 is available. Whether or not the flame closes on the centerline in the available residence time will be mainly determined by pressure and SR, although there are also coal quality effects. The thermal capacitance of the gas flow is proportional to the gas density and, therefore, increases
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for progressively higher pressures. Since the radiant heat flux to the suspension is insensitive to pressure, the core gas temperature diminishes at higher pressures. Consequently, inlet conditions that form closed flames at a lower test pressure will eventually sustain open flames at progressively higher pressures. The impact of decreasing SR is qualitatively similar. For lower SR, the volatiles flame in the near-wall region releases less heat, because its burning rate is slower at the lower O2 level. The slower heat release in the near-wall region slows the convection rate into the core, which delays ignition, and the lower O2 level diminishes the heat release after the core finally ignites. These joint effects lower core gas temperatures. These characteristics have important implications for the near-burner performance of p.c. burners at elevated pressures. As the pressure is increased, flame ignition and, by association, flame stability will become more problematic due to the greater thermal capacitance of air streams at progressively higher pressures. Increasing the suspension loadings to compensate for the higher thermal capacitance with proportionate increases in the heat release rates is infeasible, because of the severe agglomeration characteristics of p.c. at elevated pressures. For example, to re-scale the suspension loading for operation at 3.0 MPa, the loading for atmospheric pressure must be increased by a factor of 30, which is bound to exacerbate burner deposits, abrasion, erosion, agglomeration, and plugging.
Acknowledgements Financial support was provided through a subcontract from Fluent, Inc. under the High Pressure Coal Combustion Kinetics Program sponsored by the US Department of Energy.
References [1] Beer JM. Combustion technology developments in power generation in response to environmental challenges. Prog Energy Combust Sci 2000;26(4–6):301–27. [2] Niksa S, Liu G-S, Hurt H. Coal conversion submodels for design applications at elevated pressures. Part I. Coal devolatilization and char oxidation. Prog Energy Combust Sci 2003;29(5):425–77. [3] Liu GS, Niksa S. Coal conversion submodels for design applications at elevated pressures. Part II. Coal gasification. Prog Energy Combust Sci, 2004;30(6):697–717. [4] Niksa S, Cho S. Conversion of fuel-nitrogen in the primary zones of pulverized coal flames. Energy Fuels 1996;10(2):463–73. [5] Niksa S, Cho S. Assigning meaningful stoichiometric ratios for pulverized coal flames. Proc Combust Inst 1998;27:2905–13. [6] Eckstrom DJ, Hirschon A, Malhotra R. Products of pulverized coal combustion at elevated pressures. Energy Fuels, submitted for publication.
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[7] Niksa S, Liu GS. Pulverized coal flame structures at elevated pressures Part 2. Interpreting NOX production with detailed reaction mechanisms. Fuel 2005. [8] Cor J, Manton N, Mul G, Eckstrom DJ, Olson W, Malhotra R, et al. An experimental facility for the study of coal pyrolysis at 10 atmospheres. Energy Fuels 2000;14(3):692–700. [9] Manton N, Cor J, Eckstrom DK, Malhotra R, Niksa S. The impact of pressure variations on coal devolatilization products. Part 2. Detailed product distributions at 1.0 MPa. Energy Fuels 2004;18(2): 520–30. [10] Niksa S, Liu G-S. Incorporating detailed reaction mechanisms into simulations of coal-nitrogen conversion in p.f. flames. Fuel 2002;81: 2371–85. [11] Schlichting H. Boundary layer theory. 7th ed. New York: McGrawHill Book Co.; 1979.
[12] Uijttewaal WSJ, Oliemans RVA. Particle dispersion and deposition in direct numercal and large eddy simulations of vertical pipe flows. Phys Fluids 1996;8(10):2590–604. [13] Niksa S. Predicting the devolatilization behavior of any coal from its ultimate analysis. Combust Flame 1995;100:384–94. [14] Cho S, Niksa S. Elementary reaction models and correlations for burning velocities of multicomponent organic fuel mixtures. Combust Flame 1995;101:411–27. [15] Park C, Appleton JP. Shock-tube measurements of soot oxidation rates. Combust Flame 1973;20:369–79. [16] Lau C-W, Niksa S. The impact of soot on the combustion characteristics of particles of different coal types. Combust Flame 1993;95:1–21. [17] Hurt RH, Mitchell RE. Unified high-temperature char combustion kinetics for a suite of coals of various rank. Proc Combust Inst 1992; 24:1243–50.
Pulverized coal flame structures at elevated pressures. Part 1. Detailed operating conditions Gui-su Liu, Stephen Niksa* Niksa Energy Associates, 1745 Terrace Drive, Belmont, CA 94002, USA Received 27 September 2004; received in revised form 1 February 2005; accepted 7 February 2005 Available online 14 March 2005
Abstract The tests and simulations in this study characterize the chemical structure of pressurized pulverized coal flames, particularly (1) how the O2 in simulated near-burner flame zones is apportioned among the various fuel components; and (2) the burner operating conditions and mechanisms that most strongly affect flame structure. CFD simulations resolved the structures of flames of a subbituminous and two hv bituminous coals for stoichiometric ratios (SR) from 0 to 1.8 for pressures from 1.0 to 3.0 MPa. The structures of all flames were largely determined by the accumulation of particles in the turbulent boundary layer on the flow tube wall. Gaseous fuel compounds always ignited first on the wall at the burner inlet, and this flame propagated toward the flow axis to form a 2D parabolic flame surface. Within the core, residual gaseous fuels, soot, and char may have eventually reached their ignition threshold and burned in a premixed mode. Residual CO, H2, and char burned in the near-wall region after the volatiles flame had propagated deeper into the core. Whether or not the flame closed on the centerline was mainly determined by pressure and SR. Inlet conditions that formed closed flames at a lower test pressure eventually sustained open flames at progressively higher pressures. The impact of decreasing SR was qualitatively similar, due to the lower heat release rates for progressively lower SR. As the pressure is increased, flame ignition and, by association, flame stability will become more problematic due to the greater thermal capacitance of air streams at progressively higher pressures. q 2005 Elsevier Ltd. All rights reserved. Keywords: Burner simulations; CFD; Flame structure; Pressure; Pulverized coal
1. Introduction Across the globe developers of coal-fired power generators face imperatives to raise conversion efficiencies to compete better with other fuels, especially where CO2 emissions are being reduced. A multitude of advanced process concepts are under development, as surveyed recently by Beer [1]. All have one thing in common: primary conversion of the coal feed at elevated pressure. Major blocks in the technical foundation supporting development of pressurized coal conversion technology are already in place. The impact of elevated pressure on the rates and yields of the coal conversion kinetics are clearly evident in extensive databases for devolatilization and char oxidation [2]. The database on char gasification needs to be * Corresponding author. Tel.:C1 650 654 3182; fax: C1 650 654 3179. E-mail address: [email protected] (S. Niksa).
0016-2361/$ - see front matter q 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.fuel.2005.02.005
expanded further [3], but only because so many operating conditions and char characteristics are important. Notwithstanding these essential elements of pressurized pulverized coal (p.c.) flames, we have not found any characterization work in the English literature on the structures of actual p.c. flames at elevated pressures. ‘Flame structure’ denotes the spatial sequence of distinctive physicochemical processes in a self-sustaining flame. It can only be ascertained from tests with realistic suspension loadings, because suspension loading determines the proportions of gaseous fuel compounds, soot, and char. The competition among these very different fuel components for the available O2 imposes a wide range of chemical time scales which determine flame structure, in conjunction with transport phenomena. The tests and simulations in this study characterize the chemical structure of pressurized p.c. flames, particularly (1) how the O2 in simulated near-burner flame zones is apportioned among the various fuel components; and (2) the burner operating conditions and mechanisms that most strongly affect flame structure.
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This work builds upon our previous work on the process chemistry in the near-burner region of diverse p.c. flames at atmospheric pressure [4,5]. A similar axisymmetric burner system was developed to operate at pressures from 1.0 to 3.0 MPa, and used to monitor complete product distributions from one subbituminous and two high volatile (hv) bituminous coals for burner stoichiometric ratios (SR) from zero to well above unity [6]. The test results are interpreted in two stages. First, in this paper, CFD simulations identify the mechanisms responsible for several surprising features of these flames, and quantitatively assign the flow and thermal fields. Then, in a companion paper [7], the CFD results are used to specify an equivalent reactor network for the flows. Separate simulations of detailed reaction mechanisms across the reactor networks accurately predict the product distributions with a minimum number of adjustable parameters, and identify the dominant conversion and NOX production mechanisms.
2. Test facility and evaluation data 2.1. Burner configuration The test facility simulated the thermal and chemical environments in the primary zone of a pulverized coal flame without the complications of two-phase turbulent mixing. As explained elsewhere in detail [6], the coal burner was a vertically mounted, cylindrical flow reactor that imposed uniform radial heat fluxes comparable to those in utility burners. An intense external radiant field stabilized
the reaction fronts so that the burner could operate with any inlet O2 level, including none at all in cases that determined the distributions of secondary pyrolysis products. As suspensions moved along the flow tube, they were heated at rates approaching 104 8C/s to the onset temperature for primary devolatilization, released their volatiles, and burned. At any particular operating condition, O2 depletion eventually ‘quenched’ the chemistry at an intermediate stage determined by the proportions of coal and O2 at the inlet. Inlet O2 levels were progressively increased in successive cases to move the process chemistry through oxidative volatiles pyrolysis, volatiles combustion, soot combustion, and char oxidation. The basic reactor configuration, called the ‘radiant coal flow reactor (RCFR)’, was used previously in studies of p.c. combustion at atmospheric pressure [4,5] and coal pyrolysis at elevated pressures [8,9]. The version of the burner for elevated pressures is labeled as the ‘p-RCFR’. The main difference is that the particle flow is downward at atmospheric pressure and upward at elevated pressures, to counteract recirculation induced by buoyancy. The system is sketched in Fig. 1 and described elsewhere [6] in more detail. A feeder dumped classified p.c. particles into an O2/Ar entrainment stream, forming an optically thin suspension that flowed downward into a U-tube. The other end of the U-tube delivered the upward, two-phase flow into the radiant furnace section within the pressure vessel. A 27 cm mullite tube with 1.2 cm i.d. was located on the central axis of an insulated graphite cylinder wrapped by five turns of a water-cooled copper induction coil. The graphite length was 15.8 cm, although the overall hot zone length was
Fig. 1. Schematic of the p-RCFR experimental facility.
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roughly twice as long due to radiation leakage out of the ends of the entrainment tube. The graphite temperature was regulated to 1650 8C. The inlet gas velocities were 29.5 cm/s in all tests, although residence times were significantly variable due to different extents of heat release among the various test conditions. At the inlet, an annular sheath flow surrounded the entrained suspension to help prevent particles from colliding with the tube wall. The coal particles heated and expelled their volatiles which, in the presence of O2, burned along with the char particles. During combustion tests, the heat released during the combustion of volatiles and char supplemented the heating of the entrainment gas by convection from the flow tube and coal suspension. Immediately downstream of the burner, Ar was injected into the flow to quench the process chemistry, nucleate any residual tar into an aerosol, and deliver the char particles, aerosol products, and gases into the product recovery and analysis section. A cylindrical ‘soot filter’ captured soot and tar, which was subsequently extracted from the filters with tetrahydrofuran, and filtered. The solids captured on the membrane were denoted as soot, and the dissolved material was condensed out and denoted as tar. The char yield was assigned as the amount of material in a char collection basket, plus the amount of large particles taken from the first layer of the soot filter, plus any residual solids recovered from the flow system and burner flow tube. All condensible materials were analyzed for elemental composition and ash content. Concentrations of CO, HCN, NH3, and NO were monitored on-line with FTIR spectroscopy; C1–C4 hydrocarbons (CH4, C2H2CC2H4, C2H6, C3H6, C3H8, C4) and H2 were extracted for subsequent chromatography into multiport sampling valves; moisture and CO2 were monitored with separate NDIR detectors. 2.2. Test conditions Properties of the Wyodak subbituminous (PRB) and two hv bituminous coals used in the tests appear in Table 1. Note the extremely high sulfur contents of the Ill. #6 and Pit. #8 samples and the high ash content of the Ill. #6, which are characteristic of raw, as-mined samples. These coals were obtained from the Penn State Database, then aerodynamically classified. The tested samples were a mixture of two Table 1 Coal properties Coal name Pit. #8 Ill. #6 PRB
Proximate analysis (ar wt%) Ultimate analysis (daf wt%) M
Ash
VM
FC
C
H
O*
N
S
0.7 0.2 0.1
12.3 17.3 5.0
37.9 35.8 39.4
49.1 46.7 55.5
80.8 74.1 73.7
5.4 5.5 5.6
5.8 8.2 19.0
1.7 1.4 1.1
6.3 10.8 0.6
*Assigned by difference.
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sieve sizes, 75–90 mm and from 90 to 105 mm, so the mean size was about 90 mm. The test plan comprised six test series that characterized Pit. #8 at 1.0, 2.0, and 3.0 MPa, Ill. #6 at 1.0 and 2.0 MPa, and PRB at 1.0 MPa only. Each test series contains from seven to ten individual tests with progressively higher inlet O2 levels, hence, SR values. In the tests with Pit. #8 at 1.0 MPa, suspension loadings were nearly uniform at 4.7 wt%, whereas the inlet O2 mass fraction was varied from zero for the secondary pyrolysis case to 9.9%, so that SR was varied from 0 to 0.953 in fairly uniform increments. The suspension loadings were decreased from 4.7% at 1.0 MPa, to 2.3–2.5% at 2.0 MPa, to 1.55% at 3.0 MPa; in other words, coal feedrates were essentially the same at all test pressures. Inlet O2 mass fractions were regulated at the higher pressures to impose similar ranges of SR values in all test series. The maximum SR values were near-unity with Pit. #8; 1.77 with Ill. #6; and 1.27 with PRB. 2.3. Representative results The datasets on PRB coal at 1.0 MPa in Table 2 illustrate the measurement uncertainties and main tendencies in the results. The measured product distributions are arranged in order of increasing SR from left to right, beginning with the case for secondary pyrolysis. Consequently, cases toward the left are dominated by the products of primary devolatilization, secondary volatiles pyrolysis, and partial oxidation of volatiles; whereas those toward the right are governed by water gas shift equilibrium and the conversion of soot and char. When the SR exceeded one-half, all hydrocarbon fuel components had been consumed, but large amounts of CO and H2 persisted. Due to the thermal severity of all tests, tar was not even observed in the case for secondary volatiles pyrolysis alone. Soot yields decayed somewhat erratically from 9.1 daf wt%. The chars were fully ignited for all SR over 0.15, and extents of char burnout increased continuously for progressively higher SR, as expected. Mass balances on individual tests closed within G2% for all cases, while C-balances closed within G4%; H-balances within G5%; and O-balances within G2%. The only adjustments to the product distributions were for the S-species, which were not monitored. We arbitrarily expressed the volatile-S species as H2S under conditions that were sufficiently reducing to retain hydrocarbon fuel compounds, and as SO2 under conditions where essentially all the hydrocarbon fuel compounds had burned. This procedure would be expected to overestimate the volatile-S species under the most reducing conditions, because we have no evidence that all the coal-S had actually been released into the gas phase under the relatively moderate temperatures in these tests. The product distributions indicate that O2 is not apportioned sequentially among the various fuel components during the initial stages of coal combustion; clearly,
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Table 2 PRB, 1.0 MPa datasets (daf wt%) SR
0
Product distribution 6.6 CO2 H2 O 8.1 CO 12.3 CH4 2.9 C2 4.0 C3 0.4 Oils – H2 2.75 HCN 0.69 NH3 0.07 NO 0.00 Tar – H2 S 0.64 – SO2 Soot 9.1 Char 52.4 Mass balances SMass 1.000 SC 1.000 SH 0.975 SO 1.002 Burnout XHC 0 XSoot 0 0 XChar
0.028 6.6 8.1 17.1 2.9 4.0 0.3 1.8 2.75 0.69 0.07 0.00 – 0.64 – 9.1 52.4 1.016 1.021 0.997 1.000 2.0 0 0
0.121 18.9 15.9 30.6 2.4 2.2 0.1 0.9 2.41 .61 0.09 0.00 – 0.64 – 3.8 49.9 1.015 1.028 1.002 0.995 35.6 58.2 4.8
0.241 36.9 23.6 41.5 1.5 1.0 0.0 0.1 2.06 0.57 0.13 0.00 – 0.64 – 1.5 42.8 1.002 0.998 0.961 1.009 64.5 83.5 18.3
0.404 61.1 27.5 50.8 0.6 0.3 0.0 0.0 1.92 0.45 0.17 0.00 – 0.64 – 1.6 34.7 1.001 1.016 0.991 0.992 85.0 82.4 33.8
soot and char burn along with gaseous fuels from the PRB in Table 2 and also from both hv bituminous coals (as seen in Figs. 2 and 3, below). As introduced previously [5], oxygen consumption is evaluated quantitatively in terms of the carbon consumption of three major fuel components: hydrocarbon gases (CH4, C2H2, oils, and HCN), soot, and char. Carbon burnout indices for each component, CXi, are defined as C
X n 100 Xi Z C %Cð0Þ ðII W j K Meas W j ÞC fj fi jZ1
(1)
where IIWj is the yield of component j for secondary pyrolysis only; MeasWj is the measured yield of component j in a combustion test; Cfi is the carbon fraction in component i; n is the number of compounds in the fuel lump under consideration; and %C(0) is the coal-C. As seen in Table 2, all the gaseous fuel compounds were burned out when SR approached 0.6, but only about three-fourths of the soot and just over half of the char were converted at this condition. The same tendencies are evident in the datasets for both hv bituminous coals. The most significant differences are the much higher soot yields for secondary volatiles pyrolysis of 21–23 daf wt% for both coals, and the tendency for fewer gaseous hydrocarbons from coals of progressively higher rank. Char yields were similar at all pressures, because even the lowest test pressure was higher than the threshold for asymptotic total weight loss for primary
0.578 90.4 32.0 49.8 0.1 0.0 0.0 0.0 1.56 0.19 0.11 0.00 – – 1.20 3.1 24.0 0.996 0.985 0.962 1.009 97.3 65.9 54.2
0.751 126.1 36.1 43.3 0.0 0.0 0.0 0.0 1.20 0.05 0.06 0.00 – – 1.20 3.3 18.3 1.003 1.004 0.964 1.008 99.2 63.7 65.1
0.934 166.4 39.4 37.4 0.0 0.0 0.0 0.0 0.82 0.05 0.05 0.00 – – 1.20 4.1 11.5 1.001 1.039 0.954 0.991 99.5 55.0 78.1
1.092 192.6 43.8 31.3 0.0 0.0 0.0 0.0 0.57 0.00 0.02 0.08 – – 1.20 1.7 6.9 0.992 1.007 0.984 0.991 99.9 81.3 86.8
1.272 203.1 45.9 26.4 0.0 0.0 0.0 0.0 0.34 0.00 0.01 0.11 – – 1.20 2.2 4.4 1.003 0.991 0.981 1.012 100.0 75.8 91.6
devolatilization. All datasets have been reported separately [6].
3. CFD simulations CFD simulations were prepared with v. 5.5 of Fluent to assign 2D axisymmetric flow, temperature, and species concentration fields and particle trajectories for each test. Although several essential features became apparent in the CFD results, the original intent was to assign accurate thermal histories for subsequent simulations with detailed chemistry. Accordingly, various model parameters were freely adjusted to match the predicted and reported extents of consumption of volatiles, soot, and char for each test. This procedure ensured that the simulated heat release profiles were accurate, which is important because the heat release profiles are the most uncertain aspect of heat transfer in the burner, by far. The thermal analysis proceeded in two stages, as described previously [10]. First, the radiant energy transfer within the burner was analyzed to assign the temperature profile along the flow tube and the radiant flux onto the flow axis. These profiles were incorporated into the CFD simulation as boundary conditions. Convection from the flow tube through a developing boundary layer in either transitional or turbulent flow was also included. Radiant fluxes onto the centerline of the 15.8 cm radiant section ranged from 61 W/cm2 at 3.0 MPa to 66 W/cm2 at 1.0 MPa,
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Fig. 2. Burnout profiles imposed in the CFD simulations for gaseous fuels (C and solid curve), soot (B and dashed curve), and char (- and dotted curve) for Pit. #8 at 1.0 (upper), 2.0 (middle), and 3.0 MPa (lower).
Fig. 3. Burnout profiles imposed in the CFD simulations for gaseous fuels (C and solid curve), soot (B and dashed curve), and char (- and dotted curve) for Ill. #6 at 1.0 (upper) and 2.0 MPa (middle), and for PRB at 1.0 MPa (lower).
and were within the range of 50–100 W/cm2 estimated for large p. f. flames. All simulations began 6 cm upstream of the heating element, consistent with the calculated radiation leakage into the upstream section of the flow tube. CFD simulations were based on conservation equations for mass, momentum, and energy transfer in the gas phase for steady, 2D, axisymmetric flow. Core and sheath flows had equal flowrates but were distinguished at the inlet with a uniform particle suspension loading in the core and no particles in the sheath. The temperature and flow velocity at the inlet were uniform for the sheath flow, but a fully developed velocity profile was applied to the core flow, based on separate simulations of the inlet section. Transport coefficients for heat and momentum transfer to the tube wall were assigned for transitional boundary layers
at 1.0 MPa and for fully developed turbulent flow at the higher test pressures. Based on our own evaluations of the various options for the turbulence submodels and near-wall treatments, the k–3 turbulence submodel and the two-layer zonal submodel were implemented. This combination was the only one that accurately reproduced measured axial velocity profiles [11] as well as the turbulence intensity profiles across the boundary layers on vertical tubes from large-eddy simulations (LES) [12]. In conjunction with the Stochastic particle dispersion submodel evaluated with 500 particle trajectories, the turbulent flow analysis was also qualitatively consistent with the particle concentration profiles in similar flowfields from LES [12]. Turbulence intensities at the burner inlet were unimportant influences on the downstream flowfield.
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Lagrangian trajectories of individual particles were evaluated in the Stokes limit, consistent with the nearunity Reynolds numbers of small p.c. particles. Particle thermal histories accounted for absorption of the incident radiant heat flux profile balanced against the particle’s thermal capacitance, convection to the gas phase, and the heat and mass release due to devolatilization and char oxidation. Submodels for coal conversion included composition- and temperature-dependent thermophysical properties, a single first-order devolatilization reaction, a swelling factor, and char oxidation according to the classical balance between an intrinsic nth-order surface reaction and film diffusion of O2. The following chemistry submodel was incorporated: Coal/ GasVol C Soot C Char
(2)
GasVol C gO2 O2 / gH2 H2 C gCO CO
(3)
CO C 0:5O2 / CO2
(4)
H2 C 0:5O2 / H2 O
(5)
Soot C 0:5O2 / 2CO
(6)
Char C 0:5O2 / CO
(7)
where GasVol denotes the aggregate lump of CH4, C2H2, HCN, and oils. A single first-order reaction represents devolatilization kinetics, as follows: K
dmp Z kdv ðmp K ð1 K fv0 Þmp0 Þ dt
(8)
where the devolatilization rate is kdv Z Adv expðKEdv =RTp Þ and fv0 is the ultimate volatile fraction on a dry basis. The rate parameters and ultimate yield were assigned from simulations with FLASHCHAINw [13] based on the coal’s proximate and ultimate analyses, and the thermal histories from the CFD simulations. We first estimated the thermal history, then assigned rate parameters for the FLUENT simulation. The updated thermal histories from the FLUENT simulation were used to refine the rate parameters, until the rate parameters did not change. Coal swelling was represented with a swelling factor correlation. Since the primary volatiles are injected into hot gases in the p-RCFR, they are assumed to be instantaneously converted into secondary volatiles pyrolysis products, whose yields were also assigned by FLASHCHAINw. The oxidation of the gaseous fuel components proceeds in two stages: (1) partial oxidiation of hydrocarbons into CO and H2 (Eq. (3)); and (2) CO and H2 oxidation (Eqs. (4) and (5), respectively). One-step schemes are inconsistent with the rapid increase in the CO levels for progressively higher SR values (cf. Table 2). The stoichiometric coefficients in Eq. (3) were evaluated from the complete product distribution of gaseous volatiles, and vary with both coal rank and pressure. The burning rates of volatiles are first
order in both the fuel species and O2. The same global activation energy was applied to each fuel, based on the analysis of Cho and Niksa [14], but the pre-exponential factors were adjusted to match the burnout index for gaseous fuels for each test. The same rate parameters were applied to CO and H2, assuming that CO oxidation kinetics limit the equilibration of the water–gas shift chemistry. At 0.8 mm, soot agglomerates are sufficiently small that soot oxidation can be regarded as a pseudo-gas phase reaction that generates only CO. Based on the soot oxidation data of Park and Appleton [15] below 1700 8C, the soot burning rate is zero-order with respect to O2 with an activation energy of 140.5 kJ/mol. Soot diffusivities were specified by Lau and Niksa [16] as functions of soot size and gas temperature. A typical value of the Brownian diffusivity of 0.12 mm soot at 1730 8C and at atmospheric pressure is 0.57 cm2/s, versus a thermophoretic diffusivity of 0.06 cm2/s. The overall char burning rate reflects the sequential processes of film diffusion of O2 and intrinsic global kinetics, where the order of the surface kinetics is one-half. The activation energy was assigned from correlations in terms of the daf carbon content of the parent coal that were established from a database of flow reactor tests [17], whereas the pre-exponential factors were fit to match the reported char burnout profiles. Both particle density and diameter changed during char combustion, according to the empirical relation between mass and density developed by Hurt and Mitchell [17]. Hence, four pre-exponential factors in the burning rates for gaseous fuels, CO and H2, soot, and char were adjusted to match the burnout profiles of gaseous fuels, soot, and char and the total O2 utilization efficiency for each test.
4. Results 4.1. Burnout Indices Indices on the burnout of gaseous fuels, soot, and char for the tests with Pit #8 and Ill. #6 and PRB appear in Figs. 2 and 3, respectively. The curves in the figures show the results from the CFD simulations, which closely match the measured values by virtue of adjustments to the four rate parameters discussed in Section 3. The burnout profiles for gaseous volatiles from Pit. #8 at the three test pressures agree within experimental uncertainty. This does not imply that the burning rate and combustion mechanisms were independent of pressure, because the profiles are plotted versus SR rather than a time-based coordinate. Moreover, the gas temperature history reaches progressively cooler maxima at higher pressures, as discussed in Section 4.2.2. At all pressures, the indices approach the asymptotic value of 100% for an SR of 0.6. The soot burnout profiles for Pit. #8 for 2.0 and 3.0 MPa are essentially the same, but have
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higher extents of burnout than the profile for 1.0 MPa at SR values between 0.2 and 0.7. The strongest impact of pressure is seen in the char burnout profiles in Fig. 2. At a SR value of 0.95, the extents of char burnout of Pit. #8 decrease from 57.5 daf wt% at 1.0 MPa to 52.5% at 2.0 MPa to 27% at 3.0 MPa. The impact of pressure is also apparent at all lower SR values in these datasets. In fact, the O2 partial pressures were almost the same for progressively higher pressures because the same coal feed rate was used at all pressures and the O2 flowrates were the same to maintain the same SR. The diminished char burnout is mostly due to the lower particle temperatures at higher pressures, which will be further discussed below. The burnout profiles for Ill. #6 and PRB appear in Fig. 3. Unlike those for Pit. #8, the gas burnout profile for Ill. #6 at 1.0 MPa is higher than at 2.0 MPa. Soot burnout, however, is similar at both pressures. Extents of char burnout at 1.0 MPa are about 10% higher than that at 2.0 MPa across the entire SR range, consistent with the Pit. #8 char burnout profiles. In fact, the soot burnout profiles at both pressures and the gas burnout profile at 2.0 MPa for Ill. #6 are quantitatively consistent with the soot and gas burnout profiles for Pit. #8. Only the gas burnout profile at 1.0 MPa is different. The burnout indices for PRB at 1.0 MPa appear in the bottom panel in Fig. 3. The indices for gaseous fuels are essentially the same as those for Pit. #8 at 1.0 MPa, and those for soot burnout are similar to those from Ill. #6 soot at 1.0 MPa. Soot probably burns at a rate that is independent of coal rank, because it contains mostly carbon (O98.5%) with small amounts of hydrogen and nitrogen. The compositions of soot from all coal types are nearly the same. Unlike the burnout of gaseous volatiles and soot, char burnout is strongly affected by coal rank, as expected. At 1.0 MPa and a SR value of 0.95, the char burnout for Pit. #8 is 57.5 daf wt%; 60% for Ill. #8; and 77% for PRB. PRB has the fastest char burning rate, and the extent of PRB char burnout even overtakes soot burnout at a SR value of 0.85. The impact of coal rank is consistent with extensive databases on char oxidation at both atmospheric pressure [17] and elevated pressures [2]. Regardless of the pressure and coal type, gaseous volatiles combustion consumes most of the available O2 at low SR. Soot effectively competes for O2 at low SR, but burnout of the chars of the lowest rank eventually overtakes soot burnout at SR over 0.8. This char presumably burns faster than soot because chars from low rank coals usually contain substantial levels of alkaline earth and alkali cations that promote char oxidation, whereas soot does not. There are master burnout profiles for the gaseous volatiles and soot from all fuels and all pressures, albeit, each has an exception at one operating condition. But char burnout profiles are strong functions of coal quality and pressure.
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4.2. Baseline flame structure The burnout indices exhibit the expected sequence of fuel consumption from gaseous fuels to soot to char, with well established dependencies on coal rank and pressure. Considering the burner’s highly idealized geometry and inlet conditions, one could anticipate planar, nominally 1D flames across the central region of the flow tube, where the particle concentrations are highest. But such a simple structure is contradicted by the co-existence of substantial amounts of O2, CO, and H2 in the products from every test, except those for secondary volatiles pyrolysis. Typically, only 70% of the inlet O2 was consumed while about 30 daf wt% CO remained in the products, even for SR well below or well above unity. Hindered penetration of O2 from the sheath flow into the core is an obvious explanation that turns out to be incorrect. The key is the characterization of highly nonuniform particle dispersion in Section 4.2.1. The generic aspects of flame structure will be illustrated with the CFD simulations for Pit. #8 at 1.0 MPa with a SR of 0.95, a ‘baseline’ case, before the trends with SR, pressure, and coal quality are presented. 4.2.1. Particle dispersion patterns The stochastic submodel for particle dispersion in the CFD simulations imparts significant fluctuations in the particle motion. Almost immediately after injection, the particles acquire significant radial velocity components due to the turbulence and the wall collisions. All particles eventually penetrate into the sheath, and almost all of them contact the wall at some point. Once the particles move into the boundary layer, they are unlikely to escape back into the core flow, so particles accumulate in the boundary layer and deplete the suspension loading in the core. Fig. 4 shows the radial profiles of normalized particle number concentration at six cross sections. At the inlet, the particle concentration is uniform across the core, which extends to 0.707 of the tube radius (or 0.00424 m). At 1 cm downstream of the inlet, particles are still concentrated in the core, but a concentration gradient extends into the sheath. At 4 cm, the particle concentration profile has been inverted by dispersion into the sheath. Particle concentrations in the near-wall region continuously increase throughout the remainder of the burner, while the concentrations in the core diminish. Hardly any particles remain on the flow axis at the burner outlet. This result bears two important implications for the performance of the p-RCFR that were confirmed in the laboratory. First, particle agglomeration near and on the walls will likely cause operational problems, due to the combination of slow particle velocities, high particle temperatures, and high particle concentrations in the nearwall region. Second, complete O2 utilization will be difficult to achieve, because of the strong propensity for accumulation of all fuels in the near-wall region versus substantial amounts of residual O2 in the core.
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Fig. 4. Radial profiles of particle number concentration of Pit. #8 at 1.0 MPa and SR of 0.95 at axial positions of 1, 4, 6, 12, 18 and 27 cm.
4.2.2. Flow, temperature, and species concentration fields Fig. 5 shows radial profiles of gas velocity and temperature, and of mass fractions of gaseous fuels, soot CO and oxygen at six axial positions. The position labeled as zero is actually 5.8 cm upstream of the inlet to the radiant section, and the radiant section ends at 20.6 cm. The regions before and after the hot zone are affected by the radiant heat flux out the tube ends, as noted previously. The flow accelerates continuously as it moves through the tube, reaching a maximum velocity of 240 cm/s at the outlet. Due to buoyancy effects, the gas velocity reaches local maximum values near the tube wall as long as the flow is within the burner hot zone. The velocity profile becomes parabolic at 24 and 27 cm. A thermal boundary layer propagates into the core from the wall, at about 1650 8C. The centerline gas temperature rises continuously throughout the burner even while, downstream of the hot zone, the temperature profiles are inverted by convective cooling into a cooler tube wall. For this most oxidizing case in the Pit. #8 test series, the maximum gas temperature exceeds 1500 8C.
The rapid gas heating is primarily responsible for the rapid acceleration of the axial gas velocity. Gaseous volatiles and soot ignite at the inlet to the radiant section. Note their very high concentrations near the wall, where the gas temperature is hottest over this portion of the flow reactor. Particles are heated mainly by the radiant flux from the tube, but they are also heated by convection from the hotter gases near the wall. Consequently, the particles dispersed into the near-wall region devolatilize before those in the core, and the gaseous fuel concentrations are correspondingly higher near the walls. By 12 cm, the concentration spikes have been eliminated by combustion, and the highest gaseous fuel concentrations are within the core. The fuel inventory is then depleted by a flame front propagating from the near-wall region toward the flow axis. The flame is fed by diffusion of both fuel compounds and O2 from the core into the sheath flow. Since the gas temperature at 18 cm is below the threshold for ignition, the core fuel concentration is reduced by transport at this axial position. But by 24 cm, the core is fully ignited and the fuel is
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Fig. 5. Simulations of (a) axial gas velocity, (b) gas temperature, (c) gaseous volatiles mass fraction, (d) soot mass fraction, (e) CO mass fraction, and (f) O2 mass fraction for Pit. #8 at 1.0 MPa and a SR of 0.95.
depleted much faster. Almost all the gaseous fuels and soot were consumed before the reactor outlet in this particular run. The CO profiles are considerably more complex than those for the volatiles-derived fuels, because CO is generated as a product of primary devolatilization and secondary volatiles pyrolysis, as well as by partial oxidation of gaseous fuels, soot, and char. At 6 cm, there is no CO in the core, where the gases remain below the onset temperature for devolatilization. But in the near-wall region, the CO profile exhibits the spike seen in the gaseous fuel profile. Further downstream, the CO level in the core grows mostly through contributions from char oxidation through 24 cm, before it finally diminishes at the reactor outlet. But within the near-wall region, CO accumulates at roughly 1.2 wt% along the entire flow tube, due to depletion of nearwall O2 and diffusion of CO from the higher concentration in the core beyond 18 cm. The final oxidation of CO and H2 is relatively slow to begin with, and decelerated further by O2 depletion. The O2 profiles fall continuously from the inlet concentration, beginning at 6 cm in the near-wall region. Diffusion of O2 from the core into the sheath reduces the core O2 level by 12 cm. By 18 cm, the near-wall O2 level is below the threshold for gaseous volatiles combustion. Further downstream, the core O2 levels diminish more rapidly once the char suspension has ignited. Overall, 81.4%
of the O2 was consumed during this run. Not all the O2 was consumed in the near-wall region, but nearly all the residual O2 in the products came from the core flow. Hence, CO and O2 (and H2) co-exist in the products because the delayed ignition and relatively lean mixtures in the core preclude complete combustion. 4.2.3. Particle combustion histories Particle thermal histories were assigned as massweighted average values for the full population of particle trajectories in the CFD simulation. They begin to heat rapidly 150 ms from the inlet, at a nominal rate of 7300 8C/s, and were driven to 1410 8C at 380 ms. Thereafter, the particles cool at nearly the same nominal rate as they were heated. In light of the steep gradients in temperature, O2 concentration, and particle concentration in the CFD simulation, it is not surprising that the particle residence time distribution (RTD) is broad and non-normal. Almost 80% of the particles have residence times between 330 and 480 ms, but the maximum time is 790 ms. The RTD has the form of a gamma distribution and resembles the RTDs for one or a few stirred tanks in series. Whereas there are no short-circuits in this flowfield, the relatively few particles that become trapped in the near-wall region have significantly longer residence times than those remaining in
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the bulk flow. The mean residence time is 450 ms with a standard deviation of 70 ms. The particle RTD is only one of the factors responsible for the broad distributions on the extents of char burnout from this burner. For SR values less than 0.5, the burnout distributions are normal distributions with higher dispersions for progressively higher SR values. The mean values also shift toward higher values, as expected. But at SR values of 0.72 and 0.95, the burnout distributions became much more dispersed, and lost their Gaussian form, for two reasons: First, the particles in the core fully ignite and contribute to the highest bins in the burnout distribution only at the higher SR. Second, burning particles in the sheath were extinguished by O2 depletion, which freezes about half the population. 4.3. Impact of operating conditions The CFD flame structures for all other tests are qualitatively similar to the baseline case primarily because the accumulation of particles in a boundary layer on the reactor wall is insensitive to SR, pressure, and coal quality. Progressively increasing SR increases the maximum gas temperatures near the burner outlet, especially in the core. At the outlet, centerline gas temperatures were 550 8C hotter for a SR of 0.95 than for 0.15. But the maximum mean particle temperature increased by only 100 8C over this same range, because most particles are confined to the wall layer, not the core. These large variations in gas temperature with SR affect the flame structure, as seen in Fig. 6. For purposes of illustration, these fronts were arbitrarily located on the locus of positions where gas temperature is 1050 8C, which is hot enough to ignite all fuel components. As the flow moves through the tube, a flame front propagates toward the flow axis, driven by convective heat transfer from the wall and by
Fig. 6. Flame structures for Pit. #8 combustion at 1.0 MPa and at SR of (a) 0.95; (b) 0.72; (c); 0.51; (d) 0.37; (e); 0.25; and (f) 0.15. The position of the flame front is the locus of positions where the gas temperature is 1050 8C.
the heat release from combustion of gaseous volatiles and soot. Hence, the flame is sustained by outward diffusion of volatiles and O2 toward the wall, and by inward heat transfer toward the center. For some conditions, the annular flame front closes to a point on the flow axis; for others, closure cannot occur before the downstream edge of the radiant section. The sketch of this flame structure shares elements in common with both premixed Bunsen flames and laminar (Burke–Schumann) diffusion flames. But confined pressurized p.c. flames differ from both archetypes, because fuel consumption is not restricted to the volatiles flame. This flame segregates the flow according to the following three stages of combustion: (1) within the core, residual gaseous fuels, soot, and char may eventually reach their ignition threshold and burn in a premixed mode; (2) gaseous volatile fuels and soot sustain the volatiles flame as it propagates from the near-wall region toward the flow axis; and (3) residual CO, H2, and char burns in the nearwall region after the volatiles flame has propagated deeper into the core. Whether or not the flame closes on the centerline in the available residence time will be mainly determined by SR and pressure. Conditions that sustain a closed flame at a higher SR will eventually sustain open flames at progressively lower SR, which occurs at a SR of 0.15 with Pit. #8 at 1.0 MPa in Fig. 6. Comparable SR thresholds were 0.5 with PRB and 0.8 with Ill. #6. The impact of increasing pressure is similar. Inlet conditions that form closed flames at a lower test pressure will eventually sustain open flames at progressively higher pressures. In fact, the calculated maximum core gas temperatures at the burner outlet at 2.0 and 3.0 MPa with Pit. #8 were only 875 and 630 8C, which are 650 and 850 8C cooler than at 1.0 MPa. The maximum mean particle temperatures fell by 320 8C over the same pressure range. Consequently, neither flame at the higher test pressures closed across the centerline. None of the flames were closed for any coal at 2.0 and 3.0 MPa for SR values near unity. The impact of coal quality is somewhat more complex. Ill. #6 generates the coolest gas temperature profile, and PRB and Pit. #8 generate hotter and very similar profiles. At the burner outlet, the Pit. #8 core burns at 1500 8C; the PRB core burns at 1310 8C; and the Ill. #6 core burns at only at 1080 8C. But the maximum mean particle temperatures are within 100 8C, although PRB devolatilizes and ignites faster than both other coals. PRB char also has the fastest intrinsic oxidation reactivity, so it ignites at the lowest temperature and burns the fastest among these three coals. Consequently, the ultimate extents of burnout at 1.0 MPa for comparable SR were 57.9 daf wt%, 66.8 and 78.1% for Pit. #8, Ill. #6 and PRB, respectively. It may seem surprising that a coal with an intermediate char oxidation reactivity generates the most open flame structure in this comparison. The reason is the lost heat release associated with a large portion of unburned CO and
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H2 in the post-flame gases from Ill. #6. Only 50% of total weight loss from the Ill. #6 was converted into CO2 and H2O, and most of the residual fuel components were present as unburnt CO and H2 in the products. The comparable percentages were 67% for Pit. #8 and 73% for PRB. Consequently, the total heat release within the flow tube was lowest with Ill. #6, which hindered this coal’s flame stabilization.
5. Discussion Across the wide domain of test conditions in this study, the burnout indices of the major fuel groups show the expected order of consumption from gaseous volatiles to soot to char. All the indices increase for progressively greater SR, as expected. Those for both gaseous fuels and soot were insensitive to variations in both pressure and coal rank, albeit with exceptions. Soot effectively competes for the available O2 at low SR but gases win the competition under more oxidizing conditions. The extents of char burnout uniformly diminished for progressively higher pressures, due to the cooler gas temperatures that inhibited char ignition at elevated test pressures. The chars from coals of lower rank also burned faster, as expected. Based on such behavior, one could reasonably have expected that the upward flowing, sheathed cylindrical p.c. suspensions in the p-RCFR closely approached the idealized structures of planar, 1D flames. But this simple view is contradicted by the co-existence of O2, CO, and H2 in the products of every test that had O2 at the burner inlet. It was also superceded by the complex flame structures in CFD simulations, which are largely determined by the accumulation of particles in the turbulent boundary layer on the flow tube wall. At another superficial level, these flame structures resemble both premixed Bunsen flames and laminar diffusion flames, primarily because gaseous volatiles always ignite first on the wall at the burner inlet, and this flame propagates toward the core to form a 2D parabolic flame surface. This flame only marks the penetration of the thermal layer toward the centerline. It does not indicate the major fuel consumption patterns, because major portions of the combustibles burn both within the volatiles flame and beyond it. Within the core, residual gaseous fuels, soot, and char may eventually reach their ignition threshold and burn in a premixed mode. Residual CO, H2, and char burn in the near-wall region after the volatiles flame has propagated deeper into the core as long as O2 is available. Whether or not the flame closes on the centerline in the available residence time will be mainly determined by pressure and SR, although there are also coal quality effects. The thermal capacitance of the gas flow is proportional to the gas density and, therefore, increases
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for progressively higher pressures. Since the radiant heat flux to the suspension is insensitive to pressure, the core gas temperature diminishes at higher pressures. Consequently, inlet conditions that form closed flames at a lower test pressure will eventually sustain open flames at progressively higher pressures. The impact of decreasing SR is qualitatively similar. For lower SR, the volatiles flame in the near-wall region releases less heat, because its burning rate is slower at the lower O2 level. The slower heat release in the near-wall region slows the convection rate into the core, which delays ignition, and the lower O2 level diminishes the heat release after the core finally ignites. These joint effects lower core gas temperatures. These characteristics have important implications for the near-burner performance of p.c. burners at elevated pressures. As the pressure is increased, flame ignition and, by association, flame stability will become more problematic due to the greater thermal capacitance of air streams at progressively higher pressures. Increasing the suspension loadings to compensate for the higher thermal capacitance with proportionate increases in the heat release rates is infeasible, because of the severe agglomeration characteristics of p.c. at elevated pressures. For example, to re-scale the suspension loading for operation at 3.0 MPa, the loading for atmospheric pressure must be increased by a factor of 30, which is bound to exacerbate burner deposits, abrasion, erosion, agglomeration, and plugging.
Acknowledgements Financial support was provided through a subcontract from Fluent, Inc. under the High Pressure Coal Combustion Kinetics Program sponsored by the US Department of Energy.
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