Predicting the Conversion Efficiencies of Any Coal Type in CFBCs

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PREDICTING THE CONVERSION EFFICIENCIES OF ANY COAL TYPE IN CFBCS Stephen Niksa, Yasuhiro Sakurai, and Naoki Fujiwara Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.7b00060 • Publication Date (Web): 27 Feb 2017 Downloaded from http://pubs.acs.org on March 14, 2017

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PREDICTING THE CONVERSION EFFICIENCIES OF ANY COAL TYPE IN CFBCS Stephen Niksa* , Yasuhiro Sakurai1, and Naoki Fujiwara2 *Niksa Energy Associates LLC, 1745 Terrace Drive, Belmont, California, 94002 USA 650 654 3182; 650 654 3179 (fax); [email protected] 1

Coal & Environment Research Laboratory, Idemitsu Kosan Co., Ltd., 3-1 Nakasode, Sodegaura,

Chiba, 299 0267 Japan 2

Idemitsu Energy Consulting (Beijing) Co., Ltd., C602 Beijing Lufthansa Center Offices, 50

Liangmaqiao Road, Chaoyang District, Beijing, 100125 China KEYWORDS: CFBC; Burnout; Simulations; Char Oxidation

ABSTRACT. This study uses simulations with detailed chemistry to characterize the conversion of the various fuel components – noncondensable gas mixtures, soot, char fines, and coarse char - and to identify the factors that determine unburned carbon emissions (a.k.a. LOI) from CFBCs. It covers virtually the entire operating domain for commercial CFBCs with 20 test cases in 11 full-scale CFBCs for a range of coal quality from brown coals through anthracite. Soot burnout was complete for all cases except one. LOI was entirely determined by incomplete burnout of coarse char particles, except that an anthracite produced fines that burned slower than coarse

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char. Simulations for a Chinese CFBC accurately depicted how hotter furnace temperatures and load variations affect LOI. The intrinsic char oxidation reactivity determines whether or not char fines and coarse char can effectively compete with noncondensable fuel mixtures for the available O2 at lower elevations, and thereby influence NO production. More reactive chars compete more effectively. Char burning mechanisms determine which segment of the char PSD contributes to LOI, in conjunction with size variations in the transit times of char particles across CFBC risers. Relevant burning mechanisms comprise chemical reaction control, mediation by pore transport and film diffusion, thermal annealing, and ash encapsulation. Depending on the magnitude of the intrinsic char oxidation reactivity, any segment of the char PSD can make the largest contributions to flyash LOI.

Introduction Circulating fluidized bed combustors (CFBCs) have deeply penetrated power generation sectors around the world because they are efficient, consistently meet emissions regulations, and able to process low grade solid fuels with very high ash loadings. Coals across the entire rank spectrum are routinely processed in CFBCs. This widespread utilization has already spawned numerous computer simulations to estimate fuel conversion efficiencies, emissions, heat transfer rates, and pressure profiles around recirculation loops. The simulation strategies range from species mass balances in 1D with empirical hydrodynamics1-3; through population balances for fuel, ash, and sorbent particles4-6; through CFD7-9; to large eddy simulations of particle clusters entrained through risers10. Whereas the treatments for the hydrodynamics and particle dynamics cover an enormous range of mathematical sophistication, the chemistry submodels are consistently rudimentary: Almost exclusively, the ultimate yields and compositions of volatiles


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are specified as input parameters; volatiles combustion is treated with a few global, nth-order reactions; and char burning rates are based on one global nth order reaction in parallel with film diffusion of O2 to the external particle surface, and shrinking core behavior. No analysis has yet included soot formation and combustion. Rudimentary chemistry submodels are probably suitable for simulations of heat transfer rates and pressure profiles, but are bound to degrade the accuracy of estimated fuel conversion efficiencies and emissions. Moreover, the parameters in rudimentary rate expressions must be adjusted for each coal whenever simulations are used in fuel switching scenarios, even though there is neither a rational theoretical basis nor a suitable test protocol at lab-scale for such adjustments. Perhaps the greatest potential liability is that rudimentary chemistry may completely omit phenomena that, in actuality, govern critical stages of the fuel conversion process. This study presents simulations of fuel conversion in CFBCs based on hydrodynamics and particle dynamics that were deliberately simplified to accommodate detailed chemistry: FLASHCHAIN for devolatilization11; a 444-step elementary reaction mechanism for noncondensable fuel oxidation12; an established rate expression for soot oxidation13; and CBK/E for char oxidation11. The main objectives are to identify (i) which fuel components make the largest contributions to unburned carbon (UBC) in flyash, and (ii) the factors responsible for incomplete conversion of diverse coals. The predicted fuel conversion efficiencies are shown to be consistent with reported flyash loss-on-ignition (LOI) from numerous full-scale CFBCs for diverse coals. The validation database covers virtually the entire operating domain for commercial CFBCs with 20 test cases in 11 full-scale CFBCs whose range of coal quality extends from brown coals through anthracite. The predicted fuel burnout histories reveal where


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in a CFBC different fuel components burn, and which ones determine LOI levels, including heretofore unrecognized roles for heterogeneous ignition and thermal annealing of char. 2. Mathematical Analysis 2.1 Phenomenology CFBCs consist of a dense bottom bed, splash zone, riser, exit section, cyclone, and external circulation return. Dense bottom beds are fed by primary air, coal, recirculated bed ash, and, perhaps, limestone to control SO2 emissions. Within the dense bed, the flows of primary air and volatiles partition into bubbles and an emulsion phase that contains nearly all solids. Most of the coal grind releases its volatiles within the bed, which are then converted by secondary volatiles chemistry into noncondensable fuel mixtures (CO, CH4, C2H2, HCN, H2, H2S) plus soot. The residual char is ground by mechanical attrition into fines and a smaller particle size distribution (PSD) of coarse char. So the fuel components with distinctive burning rates are noncondensable gas mixtures, soot, char fines, and char. The splash zone is fed by elutriated particles from the bottom bed, especially the particle clusters ejected by bubbles that happen to rupture through the bed surface. Depending on the distribution of terminal particle velocities, particles either fall back into the bottom bed or form a dispersed suspension that moves upward in a core flow through the riser. The riser flow is usually supplemented with secondary air injection at multiple elevations. It partitions into a dense, downward, annular particle flow along the walls and a dilute upward core flow. The particle concentrations in the core and wall layers decay exponentially with height, so solids must move into the dense wall layer along the riser. Consequently, the area of the core expands with height while that of the particle layer becomes thinner. At the exit zone, a portion of the


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particles recirculate internally into the downward wall layer while the rest are accelerated through the exit duct into cyclones. The internal recirculation rate is a minor contribution in fullscale systems but a major one in lab-scale CFBCs with high aspect ratios and very high particle concentrations at the top of the riser. The cyclone rejects flue gas from the CFBC along with unconverted char and flyash. It returns the larger and denser particles into a downcomer, where they move slowly downward into the particle seal and, ultimately, into the dense bed return. The analysis proceeds through three independent stages. First, particle dynamics calculations determine the partitioning of coal and sorbent into fines, carryover, and ejection streams from the dense bottom bed. Ejection of small portions of the wakes on bubbles is the only transport mechanism that can accommodate reported ash circulation rates. In the splash, particle slip velocities determine the cutoff between particles that fall back into the dense bed, and those that move upward into the riser. Gas velocities diminish along the riser due to thinning of the dense particle wall layer with elevation and to cooling along suspended heat transfer surfaces, but accelerate via injection of secondary air. Secondary air injection more-than-compensates for the decelerating factors, so all combustibles are assumed to remain suspended while they move upward along the riser. Nearly all char passes through the cyclone into the gas cleaning system due to its relatively very low density and diminished particle size. The second calculation pass assigns extents of combustion for noncondensable fuels, soot, char fines, and char, based on finite-rate mixing of secondary air into one continuously stirred tank reactor (CSTR) for the splash zone in series with 40 CSTRs for the riser and exit zone. Bubbles and emulsion gas are completely mixed by the exit of the splash zone, and this core flow accepts secondary air along the riser based on a mixing law for jets in crossflow. Oxygen is apportioned to the various noncondensable and solid fuels based on the competitive chemical reaction


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mechanisms for the different fuel components. The third calculation pass determines extents of Ca-utilization by limestone sorbents and the SO2 capture efficiency, which are beyond the scope of this paper. 2.2 Dense Bed Analysis The sequence of calculations for the dense bubbling bed begins with hydrodynamics, following the recommendations of Pallares and Johnsson14. The hydrodynamic submodel evaluates the velocities of gases through bubbles, vb*, and emulsions, Umf; the bubble volume fraction, δbub; the void fraction in the emulsion, εmf; the expanded bed height, Hbed; and the exchange coefficients between bubbles and emulsion gases, Kbe. The gas velocities through the emulsion and bubble phases determine the transit time through the bubble phase and the mean residence time in the emulsion, and thereby affect gas compositions into the splash zone. The equations that define these variables are compiled in Table 1. The Reynolds number at minimum fluidization is evaluated from the Archimedes number, and then inverted to specify the emulsion gas velocity, and used to evaluate the voidage in the emulsion, εmf. These quantities, in turn, determine the mean bubble size, Dbub, the throughflow fraction, Ψ, and δbub. Hence, the gas velocity through the bubble phase, the voidage of the entire bed, εbed, and the exchange coefficient between bubbles and emulsion gas are fully specified. Commercial CFBCs typically operate with Geldart B particles in the fast bubble regime15. In the fuel conversion submodel, the emulsion is regarded as well mixed, and the bubble phase is in plug flow. When coal is fed into the splash zone instead of directly into the dense bed, only the portion of the feedstream without sufficient buoyancy to move into the riser enters the dense bed. The proximate and ultimate analyses of the coal are initially used to predict the transient


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devolatilization behavior of coal particles in size increments that span the PSD within an isothermal emulsion. The FLASHCHAIN® equations11 are solved simultaneously with an energy balance for individual particles to predict the particle’s thermal history and product distribution. FLASHCHAIN® first predicts the complete distribution of primary devolatilization products, then this distribution is transformed by secondary volatiles pyrolysis into the fuel mixtures that actually burn in the CFBC: soot, char, CO, CH4, C2H2, HCN, H2, and H2S diluted by N2, CO2, and H2O. Then the submodel for volatiles conversion is activated to determine the extents of conversion of volatiles in both the emulsion and bubble phases. FLASHCHAIN also predicts the char composition and the swollen char PSD, based on swelling factor correlations. In the fast bubble regime, volatiles burning rates are limited by the exchange rates of emulsion volatiles into bubbles and of O2 in bubbles into emulsion gas. In dense bottom beds, nominal emulsion gas velocities are about 10 cm/s and expanded bed heights are about 1 m, so the transit time is roughly 10 s. In the fast bubble regime, mean bubble gas velocities are roughly 10 m/s, so the transit time for bubble gas is only a few hundred milliseconds. Such a short transit time would, of itself, raise concerns about whether or not the combustion kinetics would be fast enough to convert noncondensable fuels at the moderate operating temperatures of bottom beds. But the estimated outlet temperature of bubble gas, at only 40 - 50°C, is hundreds of degrees below hydrocarbon ignition temperatures, and therefore much too cool to sustain any combustion at all. Oxidation within emulsion gases is also minimal, because the exchange rates for O2 from bubbles into the emulsion are relatively very slow compared to bubble transit times. Moreover, if the noncondensable fuels are only partially converted within dense beds, then char conversion will be negligible because char burns much slower than noncondensable gas mixtures. So the


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fuel mixtures released by emulsion gases contain noncondensable fuels, partial oxidation products, soot, char fines, and a PSD of char particles. 2.3 Particle Population Balances The char PSD is developed from a population balance that accounts for influx, the density and size changes associated with devolatilization, mechanical attrition, elutriation of fines, ejection in the wakes of bubbles, and withdrawal of bottom ash. Since char oxidation in the dense bed is negligible, char is mostly ground into fines by attrition and ejected by bubble wakes into the splash zone. The population balance for the char PSD is

FC,0 PC,0 (aC ) + FC,S PC,S (aC ) + FC,R PC,R (aC ) − FC,E PC,DB(aC ) − FC,BAPC,DB(aC ) −WC,DBκC (aC )PC,DB − WC , DB

 1 dmC  da  d   = 0  PC , DB (aC ) C  + WC , DB PC , DB  daC  dt  m dt  C 

(1)

where FC,i are mass flowrates of char and the subscripts 0, S, R, E, and BA denote char from the coal feed; settling from the splash zone; injection from the particle seal; ejection into the splash zone; and bottom ash, respectively. The particle size frequency distributions, PC,i, are functions of the particle radius, aC. The char feedrate is based on the coal feedrate adjusted for release of the ultimate volatiles yield. The settling rate is evaluated from the segment of the char PSD in the splash zone with downward slip velocities, where terminal velocities are assigned as 2 . 335 − 1 . 744 φ B   18  Ar 2 / 3 +  Ar 1 / 6   Ut = 1 / 3   ρg2    µ ( ρ S − ρ g ) g 

−1

(2)


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The flowrate for char particle ejection by bubbles at the surface of the bed, FC,E, is evaluated from

F C , E = y C γ C δ bub U br A bed ρ bed

(3)

where yC is the mass fraction of char in bed solids; γC is the volume of solids in wakes per unit bubble volume15; Ubr is the bubble rise component in the bubble gas speed; and is the average density of bed solids. The term for elutriation in eq.1 contains the char holdup in the bed, WC,DB and the elutriation rate constant, κC, as reported by Selcuk et al.16. The term containing daC/dt comprises independent contributions for the rate of size reduction by oxidation and mechanical attrition. The oxidation contribution accounts for simultaneous reductions in density and size during char combustion, and the attrition rate is proportional to U0-Umf and a constant reported by Arena et al.17. In the final term, dmC/dt is the particle burning rate. The PSDs for both ejected solids and bottom ash are assigned as the bed PSD on the premise of representative sampling. Similar analyses for the dynamics of sorbent, circulating ash, and inerts determine their respective PSDs within and out of the dense bed. Particle fluxes into the splash zone are mostly determined by the carryover of larger solids by rupturing bubbles, because it is the only mechanism that can accommodate reported ash circulation rates. 2.4 Splash Zone, Riser, and Exit Zone Within a splash zone, the ejected particle fluxes partition into coarse particle clouds that settle back into the dense bed and suspensions of smaller particles with sufficient buoyancy to move


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upward through the riser, based on terminal velocities for the solids PSDs (cf. eq. 2). Air released from bubbles mixes with the fuel mixtures from the bed emulsion and burns, starting with the most reactive fuel components. The burnout sequence for the various fuel components is determined by their distinctive oxidation kinetics. Accordingly, the simulations incorporate a 444-step elementary reaction mechanism for volatiles combustion including N-species conversion12; Nagle-Strickland-Constable kinetics for soot burnout13; and the carbon burnout kinetics model (CBK/E) for the oxidation of char fines and char11. As evident in the simulation results, only char conversion mechanisms that account for finite-rate surface chemistry, pore transport, film diffusion, transport through an ash encapsulation layer, and the reduction in the intrinsic reactivity due to thermal annealing are suitable for CFBC applications. Unfortunately, there is little, if any, quantitative information on patterns and rates of mixing in splash zones in commercial CFBCs. Mixing of primary air in the splash zone is represented with a single CSTR, while secondary air injections into the splash and along the riser are represented with finite-rate decays for jets in crossflow into series of CSTRs that closely approximate plug flow. The gas velocity, vG,S, is determined from the superficial gas velocity from the top of the dense bed, with adjustments for the diminished flow area and for the appreciable volume fraction of solids, according to

vG , S =

U G , DB ( H )  2tWALL  (1 − ε S , S ) 1 −  D RISER  

2

=

δ BUB vb* + (1 − δ

BUB

) ve

 2 tWALL  (1 − ε S , S ) 1 −  D RISER  

2

(4a)

where

tWALL (h) H H − h 0.73 = 0.55 Ret−0.22 ( 0 )0.21 ( 0 ) DRISER DRISER DRISER

(4b)


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where h is elevation above the top of the dense bed, and H0 is the height where the thickness of the annular particle layer starts to thin out. The nominal transit time across the splash zone for each increment in the solids PSD is given by

τ S (a i ) =

HS v S (a i )

(5)

where vS(ai) is the slip velocity of a char particle with radius ai. The height of the splash zone above the dense bed, HS, is estimated as twice the bubble diameter through the top of the dense bed. Since the slip velocities through the splash and riser vary with size, each size increment passes through the CSTR network in distinctive transit times, and these variations are substantial enough to affect extents of char conversion, as illustrated below. A multitude of fuel species – CO, H2, CH4, C2H2, HCN, H2S, soot, char fines, and coarse char – compete for the available O2 along the splash and riser. Realistic chemical kinetics for each distinctive combustion process govern the competition for O2 without any a priori assumptions on this partitioning. The reaction mechanism for chemistry in the gas phase contains 444 elementary reactions among 66 species12. All rate parameters were assigned independently, so there are also no adjustable parameters in the submodel for gas phase chemistry. The soot burnout mechanism also contains no adjustable parameters13. Char burning rates are determined by transitions from chemical kinetic control to mediation by internal pore diffusion and external film diffusion, along with thermal annealing and ash encapsulation (of high-ash chars). CBK/E incorporates all these features, and depicts the impact of variation in gas temperature, O2 level, and char particle size within useful quantitative tolerances11. However, it is not yet possible to


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specify the initial char reactivity within useful tolerances from the standard coal properties. This parameter is assigned in a calibration for LOI for any coal type from commercial CFBCs, as presented below. Hence, the only adjustable parameter in these reaction mechanisms is the initial intrinsic char reactivity. The primary process stream in this reaction system begins as the gases in the bubbles and emulsion from the dense bottom bed into the first CSTR. When a portion of the coal feed remains in the splash zone or penetrates into the riser, succeeding reactors blend secondary volatiles into the process stream, along with any conversion products from soot, char fines, and char. Increments of secondary air are also added to individual CSTRs. All these additions are evaluated from devolatilization kinetics for volatiles; from soot burning rates for soot; from char burning rates for conversion products from char fines and char; and from the mixing submodel for secondary air. Similarly, the temperature of each reactor in the series is based on a measured temperature profile across the riser. Mean gas residence times in each reactor are based on the calculated gas velocities and the elevation increment that each reactor represents. Fuel conversion is analyzed with NEA’s CHEMNET methodology18. In the CFBC chemistry simulations, the splash zone is represented with a single CSTR; the riser by 20 reactors in series; and the exit turn and exit duct by another 20 reactors. Due to the large number of reactors used for the riser, the flowfield is essentially in plug flow. All of these specifications could be refined by CFD simulations. The conservation principle for gas species i across any CSTR in this system is given by


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FPj,+i 1 y Pj +,i1 − FPj,i y Pj ,i − FEj, i y Ej ,i =

ν Ci M i (1 − x ASH )

ν Ci y C , f F f (1 − x ASH ) M i Χ C ( a f ) M C'

M C' +

η C ,0 FC ,0 + η C ,C FC , C + η C , EJ FC , EJ  +

M iν Si FS0 Χ S − ∫ ω i M i dV M S'

(6)

where FP,ij and FP,ij+1 are the mass flowrates of species i into and out of the jth CSTR, respectively, and FE,ij is the increment of secondary air; FC,0, FC,C, and FC,EF are the char flowrates from the coal feedstream, and the carryover and ejected streams from the dense bed; ηC,0 is the extent of char conversion integrated across the entire char PSD; ηC,C and ηC,EJ are analogous contributions for the carryover and ejected char streams; XC(af) is the increment in the extent of char conversion for char fines, of size af; νCi is the stoichiometric coefficient for species i in the char conversion reaction; Mi and MC’ are the molecular weight of species i and the mean element weights in char; XS is incremental soot conversion; νSi is the stoichiometric coefficient for species i in the soot conversion reaction; ωi is the net volumetric generation rate of species i from homogeneous chemistry; and dV is a volume increment. Since the composition within the jth CSTR is uniform, by definition, the last integral is readily evaluated as the product of the mean homogeneous reaction rate and the reactor volume. Note the explicit incorporation of results from the particle dynamics calculations, which specify the partitioning of char into carryover and ejection streams and also determine the flowrate of fines from mechanical elutriation in the dense bed. The overall mass balance has a similar form, except that no homogeneous reaction term appears because chemistry in the gas phase cannot change the mass of the process stream. This system of equations comprises two implicit equations for two unknowns for each species, FP,ij+1 and yP,ij+1. Once specified, these unknowns determine the extents of solids conversion, since char oxidation kinetics may be evaluated once the gas composition has been specified; similarly,


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they determine XS. The specified unknowns also determine the homogeneous reaction rate, ωi, because the homogeneous reaction mechanism specifies all production rates in terms of the species concentrations. Independent balances on the three solids streams for char, char fines, and soot have a similar form, without the contributions for air entrainment and homogeneous chemistry. They determine the flowrate and PSD of the char effluent, and the effluent flowrates of char fines and soot. 3. Validation Database The validation database contains eleven CFBCs with ratings from 1.2 MWth to 235 MW. Ratings, coal types, and literature citations are compiled in Table 2. Replicate tests were performed at five units, usually with different fuels or at different loads, so there are 20 distinct case studies. Commercial CFBCs in the small size class, from 20 to 35 MW, and for power generation, from 100 to 300 MW, are well represented. Four CFBCs are in China and the rest are in Europe. Only one Chinese CFBC did not stage the furnace for NOX control. All the commercial CFBCs have rectangular furnace cross sections that converge into a funnel shape for the dense bottom bed, except for the circular CFBC E furnace. The flow cross sections and furnace heights were specified for all CFBCs, but only two heights of the dense bottom beds were reported. The larger CFBCs have multiple fuel injection ports, and elevations for fuel injectors and solids return ports were usually indicated. Secondary air elevations were reported for all cases that had secondary air injectors, except CFBC K. Dense bed operating temperatures were between 830 and 900°C, except for the 960°C imposed at CFBC C. The CFBCs without limestone injection tend toward hotter operating temperatures. The maximum temperatures were somewhat hotter, and usually recorded at about one-third of the riser height. Temperatures drop across the upper riser elevations, usually by more than 50 degrees. The stoichiometric ratios


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(S.R.) of the dense beds range from 0.50 to 0.85. While these nominal values indicate deeply reducing atmospheres in the dense beds, they do not represent the severe segregation between primary air and emulsion gases in our interpretation, so we expect the actual S.R. of emulsion gases to approach zero. Nineteen different coals cover ranks from brown coals to anthracite with especially good coverage of brown coal and high volatile (hv) bituminous samples. Many samples have excessive ash contents of 40 to 50 %, particularly among the Chinese CFBCs. All required input data that was omitted from the test reports was standardized in the simulations, as follows: (1) The PSD of circulating ash was specified to have 60 % under 375µm and 20 % over 575µm, which gives a mean diameter of 303µm. (2) A circulation ratio of 30 was applied to all cases. (3) Dense bed heights were set to 1.0 m for all cases except one, which reported this value. (4) Bottom ash withdrawal rates were set to 25 % of the coal ash feedrate, based on Yang et al.29. All other specifications were taken from the field test reports and implemented directly in the simulations. Fuel feedrates were assigned to match the reported maximum furnace rating, based on a thermal conversion efficiency of 33.3 %. Excess air levels were either reported or assumed to be 20 %. The primary air split was either reported or assumed to be 60 %. PSDs for coal were either reported or assumed to give a mean size from 2 – 3 mm. PSDs for limestone were either reported or assumed to give a mean size of 150µm. Each simulation took under 10 min on a microprocessor operating at 3.2 GHz.


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4. Results 4.1 Calibration With Reported LOI Levels There are no adjustable parameters in the reaction mechanisms for devolatilization, secondary volatiles pyrolysis, volatiles combustion, and soot burnout. A single reactivity parameter for char burnout was specified for each coal to match a correlation of the reported LOI levels from two dozen commercial Chinese CFBCs29, as seen in Fig. 1. The x-axis is a volatility index defined as the ratio of the proximate volatile matter, in daf wt. %, to the lower calorific value, in g/MJ, which could be evaluated for all coals in the database from standard coal properties. This correlation represents coals across the entire rank spectrum whose LOI levels varied from 4 to 28 wt. %. In so far as the Chinese LOI correlation pertains to the CFBCs in the validation database, this calibration procedure ensures that the coal conversion efficiencies from the simulations are reasonably accurate. It was also reassuring that none of the adjustments to the default reactivities from CBK/E were inordinate; among 19 different coals in the database, the median adjustment to the results for default char burnout kinetics was 20 %. Moreover, three cases for a Chinese CFBC demonstrated that hotter furnace temperatures and greater furnace loads reduced predicted LOI, in accord with established trends30. 4.2 Dense Bottom Bed Performance The assigned dense bed operating conditions for the database are in accord with expectations: superficial gas velocities through emulsions and bubbles are 4 – 6 cm/s and 6 – 12 m/s, respectively. Bubble volume fractions vary from 0.25 to 0.4, and the throughflow fractions vary from 0.5 – 0.7. Bed voidages vary from 0.5 – 0.6. Exchange coefficients from bubbles to emulsions vary from 0.05 – 0.10 s-1.


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The fact that the gas throughflow in bubbles exceeds the bubble rise velocity, so that bubbles provide an effective shortcircuit around the emulsion phase to the top of the bed, is widely recognized as one of the most distinctive features of dense bottom beds. However, the disparate flowrates through emulsions and bubbles also carry important implications for fuel conversion in bottom beds. Since all particles are confined to the emulsion, volatiles are necessarily released into the emulsion gases, which raises the question: Can the flowrate of emulsion gases in the simulations accommodate the flowrate of volatiles associated with actual coal feedrates ? Surprisingly, the gas flowrate through the emulsion is too slow to accommodate all the volatiles for coals of even the lowest volatility. With hv bituminous coals, most of the volatiles must spontaneously enter the bubble phase, along with all the primary air from the distributor. With low volatility coals, including anthracites, most of the volatiles remain in the emulsion, but all O2 from the primary air is still excluded from the emulsion at the inlet air distributor plate. Volatiles partitioning is strongly affected by any factors that affect the emulsion velocity, density, and viscosity33. Consequently, it is important to evaluate the emulsion gas density and viscosity from the volatiles composition, rather than a vitiated air composition. The strongest influence is the mean size of circulating ash, due to its impact on the emulsion velocity. Larger circulating ash shifts the gas partitioning to retain much more of the volatiles in the emulsion which, in turn, reduces volatiles conversion because emulsion volatiles are only converted via gas exchange from fast bubbles. Faster gas exchange coefficients and hotter bed temperatures also promote volatiles conversion. However, the S. R. of the dense bed, the expanded bed height, and the circulation ratio are unimportant. Across the database, the hydrodynamic parameters exhibit relatively little variation, probably because the mean size of circulating ash is the same in all cases, so the variation in the emulsion


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velocity is relatively moderate. The most important variations are collected in Table 3 which, from left to right, shows the test label; CFBC rating; ultimate volatiles yield; volatiles percentage accommodated by the emulsion flow; predicted extent of volatiles conversion in the dense bed; and flowrates of char fines and large char ejected into the splash zone. The ultimate devolatilization yields range from almost 8 daf wt. % to 65 %, and these extreme values coincide with the extremes in fuel rank of anthracite and lignite. In turn, this wide range is responsible for the acute variability in the volatiles split between the phases, from 8 to 82 % into the emulsion. Most important, the emulsions cannot accommodate all the volatiles from even those coals with the lowest volatility, including several medium and low volatile bituminous coals and an anthracite. Hence, the premise that emulsion gases are composed of volatiles with small amounts of their partial oxidation products has now been demonstrated across the complete domain of fuel quality in commercial applications. Due to the relatively slow exchange of O2 into the emulsion, extents of volatiles conversion rarely exceeds a few percent. Considering that O2 exchange into the emulsion is always insufficient to burnout all volatiles and soot, it is inconceivable that appreciable char can be converted in dense beds because char burning rates are much slower than the noncondensable fuels’. According to these simulations, the primary functions of dense bottom beds are to partition the fuel into fines and large particles, and the sorbent into fines, flyash, and circulating sorbent; and to eject circulating ash into the splash zone. The analysis predicts that dense beds generate an abundance of char fines that are elutriated through the splash zone. As seen in Table 3, most of the CFBCs passed more char fines than char particles into the risers, by factors of two to six; only six cases had more char particles than fines. The coarse char particles had much smaller sizes than the mean size of the coal grind, and the size cutoff for fines is only 1.5 µm. Hence,


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dense bottom beds should be recognized as highly effective grinding devices that micronize most of the coal feed before it burns. In three CFBCs, E, H2, and H3, the dense beds generate comparable streams of fines and particles; and in three units, A4, H1, and I, the beds generate more coarse particles than fines. The latter three CFBCs are different units at the same power station, and the relatively stronger air flows through these dense beds and splash zones elutriate char particles before they could be micronized in the dense bed. Case A4 processed anthracite that was deliberately pulverized into a relatively finer grind, to compensate for the low intrinsic burning rate of this low volatility fuel; so here too, smaller char particles were elutriated before they could be micronized in the dense bed. Notwithstanding these exceptions, most dense bottom beds convert the fuel stream into mostly fines plus a particulate suspension whose mean size is much smaller than the nominal coal size. 4.2 Burnout of Fuel Mixtures Throughout this section, generic results are based on Case E which is a commercial furnace rated at 105 MWth that fires a blend of hv bituminous coals. The elevation of the bottom of the exit duct is 20 m above the distributor plate, and the riser includes two secondary air injectors at 2.3 and 6.9 m. The splash zone extends from 1 to 2.2 m. Profiles of transit time, temperature, and flowrate of the process stream appear in Fig. 2. The gas flowrate is normalized by the value at the cyclone exit. The increases in the flowrate reflect both secondary air additions as well as the conversion of solid fuels to gaseous products. Due primarily to the secondary air injections, the transit time per unit elevation diminishes across the splash zone and the first several meters of the riser. These additions plus the production of gaseous products more-than-doubles the gas flow across this region. Thereafter, the gas flowrate is almost steady because only larger char remains to be converted, and transit time increases in direct proportion to elevation. Midway up


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the riser, the gas temperature passes through a weak maximum, and starts to diminish when the flow contacts a pendant superheater from 9.2 to 12.7 m. Due to the assumption in the simulations that emulsion gases and bubbles are fully mixed within the splash zone (which covers elevations from 1.0 to 2.2 m in Fig. 2), most of the O2 from the primary air stream and all noncondensable fuels and intermediates were completely consumed in the splash zone in Fig. 2 and in all other CFBCs in the database, which is too fast. This flaw reflects the absence of any objective means to estimate the mixing rates in this section of the CFBC, and was adopted to avoid unbounded tuning of mixing parameters. It should eventually be replaced by a finite-rate mixing process tuned to CFD simulations. As seen in Fig. 2, char fines burn fastest among the solid fuels, by far, and are completely consumed in the splash zone with most coal types. The relatively very fast burning rate of fines reflects the high intrinsic reactivity in conjunction with the very small size, which eliminates all transport resistances. Conversely, hardly any soot or large char burns in the splash in this particular case. Soot burns out through the first three-fourths of the CFBC, while coarse char particles cannot completely burn out in the available transit time. The important implication is that char fines can effectively compete for O2 with the noncondensable fuels. It is the only particulate form that can mediate the burning rates of noncondensable fuels, and thereby influence N-species conversion (although soot will catalyze radical recombinations and heterogeneous NO reduction). However, this implication cannot universally apply to all coal types, because intrinsic char oxidation rates strongly diminish for coals of progressively higher rank11. In most cases LOI was entirely determined by incomplete burnout of large char particles, with no contribution from char fines. But in six cases, incomplete conversion of fines was a


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contributing factor, albeit a small one. Only an anthracite produced fines that burned slower than coarse char to give burnout levels of 82 % for fines vs. 96 % for coarse char. In pulverized coal combustors, extents of char burnout monotonically decrease for progressively greater char sizes, and the UBC in flyash LOI exclusively comes from the largest particles in the coal grind. But with the much coarser coal grinds fed into CFBCs, the size dependence is more complex, as seen in Fig. 3 (and also in Figs 6 and 7, below). Extents of burnout do diminish with size for the smallest particles in the entrained char PSD, but the burnout then grows for even larger sizes. The minimum reflects the much longer transit times for the largest particles, which compensate for the stiffer transport resistances in their burning rates, and explains why the burnout for the largest particles exceeds that for moderately smaller ones. Consequently, the largest sizes in the char PSD do not make the largest contributions to LOI. Rather, LOI comprises a fairly broad segment of the larger char particles whose contributions by size are variable. Among entrained char particles (excluding fines), only the smallest sizes effectively compete for O2 in the splash zone. This tendency is illustrated in Fig. 4 by histories of burnout and particle temperatures for three sizes that span the PSD of entrained char. The smallest size is rapidly consumed in the splash and lowest riser elevations, even though this size class did not achieve a fully ignited burning state. It burned under chemical control, although the intrinsic oxidation reactivity was still fast enough to complete burnout. In contrast, the two larger sizes ignite to temperatures that are much hotter than local gas temperatures. At face value, the abrupt and simultaneous surges in both burnout and temperature look like the signatures for ignition delays. In fact, the times before the particle ignite are the estimated transit times through the splash zone, which vary from 400 ms to 4.5 s for this char PSD. Since only primary air moves


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through the splash in this case, and since nearly all the noncondensable fuels and char fines are converted in the splash, the estimated O2 concentration is less than 0.2 % in this case, which is too low to ignite coarse char. Only particles smaller than 100µm give appreciable burnout in the splash. Larger char particles ignite as soon as they contact the much greater O2 levels above the secondary air injectors, as evident in Fig. 4 by much faster burning rates and abrupt surges of 300 to 400°C in the particle temperatures. At these hotter temperatures, most char particles burn out within their total transit times, at which point the particle temperatures relax back to the gas temperature. Since the particle temperatures after ignition exceed the local gas temperatures by hundreds of degrees, coarse char will be subject to thermal annealing. A complete burnout history and the impact of thermal annealing on the intrinsic reactivity appear in Fig. 5. This combustion history includes the decay in the normalized char particle diameter, which indicates continuous shrinkage during char burnout, although the particle density also diminished throughout this burnout history. Consistency with shrinking core behavior is gauged by the Chi-parameter in the right panel, which is the ratio of the instantaneous burning rate to the burning rate determined by film diffusion of O2 onto the external particle surface. Chi ranges from zero to unity, and thereby indicates the transition from chemical kinetic control to film diffusion control. It approaches unity as the char moves across the splash zone, because burning rates become limited by film diffusion for progressively lower O2 levels, and O2 levels are very low in this splash zone. After ignition, this parameter tracks the variations in particle temperature, because the surface oxidation kinetics have a much stronger temperature dependence than the O2 film diffusion rate. Accordingly, Chi exceeds 0.8 soon after ignition and then varies from 0.8 to 0.9 during char burnout. Such large values indicate substantial mediation of the char burning rate by


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film transport, with an appreciable contribution from the oxidation kinetics. Strictly speaking, shrinking core behavior would be associated with a Chi-value of unity throughout burnout. The effectiveness factor, η, in the left panel of Fig. 5 also does not corroborate shrinking core behavior. This parameter is a gauge for the penetration of O2 toward the centers of char particles. It is very low at ignition, but rises continuously during burnout to a value of twothirds, reflecting the expansion of the internal voidage throughout char conversion. The final important aspect of char conversion in this case is the impact of thermal annealing on the intrinsic char oxidation reactivity. This feature is illustrated in Fig. 5 by the ratio of the instantaneous rate constant for oxidation to its initial value. Even while the particle traverses the splash zone at 880°C, the reactivity diminishes by a factor of fifty. Since the annealing rate surges after ignition due to the much hotter temperatures, the reactivity decays at a faster rate by almost another order of magnitude. The simulation results only account for annealing after char has been expelled from the dense bed, so the initial impact would have been slightly greater if soaking in the dense bed had been included. This strong impact of annealing affects the ignition characteristics and, in principle, might also reduce the potential to manage LOI by recirculating the largest unconverted char particles through multiple passes through the CFBC. To investigate this possibility, char particles of the density and size at the end of a first pass through the CFBC were artificially annealed to the same reduction in burning rate before they completed another identical cycle through the CFBC. Due to their much lower particle density and smaller size, these particles ignited and burned slightly hotter on the second pass, despite their much slower reactivity at the ignition point, and achieved complete burnout. This calculation did not account for fragmentation during the second exposure to the dense bed or for the shorter transit time of a lighter, smaller particle through the


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second pass, which would reduce the predicted extent of burnout at the CFBC exit. Even so, it gives no evidence that thermal annealing would retard ignition during subsequent passes through a CFBC, because most annealing occurs upstream of the ignition point in the first pass. 4.3 Coal Quality Impacts Two cases in this section illustrate the burnout characteristics for coals at the extremes of the rank spectrum. The CFBC in Case J burns brown coal with 30 as rec’d wt. % ash, which generates a char whose ash level exceeds 60 %. The initial oxidation reactivity is 50 times faster than the char from the bituminous blend considered in Figs. 2 – 5. Secondary air is injected into the dense bottom bed and also above the splash zone, so this splash zone has much greater O2 levels than many other CFBCs. Burnout as a function of the initial coal size is plotted for several elevations in Fig. 6, along with histories of burnout and particle temperature for two size classes. This case abides by the expected tendency for p. c. combustors for monotonically decreasing burnout for progressively larger particle sizes. But the multiple reaction and transport mechanisms in this case expose complexity across the entire PSD. The initial reactivity is fast enough for the entire PSD to compete with noncondensable fuels and char fines for the available O2 in the splash zone. Indeed, burnout in the splash varies from 40 to 97 %, with a very weak minimum in the middle of the PSD. The important implication of these high conversion levels is that transit times through the riser are much less sensitive to size, and therefore much more uniform than the situation in Fig. 3, because particle velocities across the PSD relax to the gas velocity. Even though the entire PSD ignites and reaches particle temperatures well above the local gas temperature in the splash zone, two factors immediately extinguish the combustion into a much


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cooler slow-burn state: (1) The large diffusional resistance across the ash encapsulation layer surrounding the combustibles; and (2) Thermal annealing. Thereafter, the particles burn isothermally at a slower rate, and achieve ultimate extents of burnout determined by their relatively uniform transit times through the riser. This is why the curves of burnout vs. size in Fig. 6 are similar for all elevations. Hence, with low rank coals, thermal annealing and, for highash coals, the resistance of an ash encapsulation layer, are the determining factors on LOI levels. The opposite extreme is based on the CFBC in Case A4, which burns an anthracite with 8 % ash and one-fourth the oxidation reactivity of the bituminous blend in Figs. 2 – 5. This furnace uses a much lower staging level than the others, so the estimated O2 concentration in the splash zone is much greater than the other cases. As seen in Fig. 7, this case gives the most complex size dependence of all in the predicted extents of burnout; at each elevation burnout as a function of size first passes through a local maximum before it diminishes through a local minimum at intermediate sizes, and then rises again for the largest size classes. The segments from the local maximum through the largest sizes reiterate the impact of longer transit times for progressively larger particles (cf. Fig 3). The segment from the smallest size to the local maximum reflects ignition problems for the smaller sizes. As seen in the burnout histories for the extreme size classes in Fig. 7, smaller sizes never achieve a fully ignited state, due to their relatively faster heat loss rates to the entrainment stream, whereas larger sizes ignite to a particle temperature that remains much hotter than the local gas temperature throughout most of the transit time. Consequently, the smaller particles burn slower than the larger particles, and thereby have lower extents of burnout. In fact, they burn with Chi-values approaching zero and effectiveness factors approaching unity, which is the opposite limit from shrinking core behavior. This behavior also explains why char fines only contribute to LOI for coals with the slowest oxidation reactivity.


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Beyond the local maximum, extents of burnout diminish for larger particles because their film diffusion rates are slower, and also because thermal annealing partially extinguishes the fully ignited state sooner for progressively larger particles. Eventually, this tendency is compensated by the longer transit times for larger particles, which enhances burnout. Consequently, across the PSD for low reactivity chars, burning rates shift from chemical reaction control (Chi = 0; η = 1) toward film diffusion-limited burning mediated by internal pore diffusion (Chi > 0.5; η < 0.2). No size class exhibits shrinking core behavior. 5. Discussion In dense bottom beds, fast bubbles exchange very little O2 with emulsions and remain much cooler than ignition temperatures for noncondensable fuel mixtures. Consequently, no fuel components are converted in the bubble phase, even though most volatiles must move through bubbles with all high volatility coals. Emulsion gases are extremely reducing atmospheres because emulsions cannot accommodate all the volatiles from even those coals with the lowest volatility, so no primary air enters emulsions, and O2 exchange from bubbles is relatively slow. Consequently, only minor portions of noncondensable fuel mixtures are converted in dense bottom beds, although the exact amounts are subject to the large uncertainties on exchange coefficients from bubbles to emulsions. Notwithstanding, there seems to be little doubt that extents of char burnout in bottom beds are negligible. Since no O2 accumulated in the emulsions of any of the CFBCs in our database, none of these dense beds sustained any char conversion at all. The primary functions of dense bottom beds are to partition the fuel into fines and large particles, and the sorbent into fines, flyash, and circulating sorbent; and to eject circulating ash


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into the riser. Most of the CFBCs passed more char fines than char particles into the risers, and the larger char particles had much smaller sizes than the mean sizes of the coal grinds. Dense bottom beds should therefore be recognized as highly effective grinding devices that micronize most of the coal feed before it is subjected to char burnout. Due to the assumption in the simulations that emulsion gases and bubbles are fully mixed before they burn in the splash zone, most of the O2 from the primary air stream and all noncondensable fuels (CH4, C2H2, H2, HCN, H2S) and intermediates (CO, H2) burned out in the splash zone. Char fines burn fastest, by far, among the solid fuels and are completely consumed in the splash zone with most coals. Conversely, hardly any soot burns in the splash due to its low inherent reactivity. Extents of char burnout in the splash are also determined by the intrinsic oxidation reactivity, which spans two orders of magnitude for coals across the rank spectrum11. With reactive low rank coals, the entire char PSD can be at least partially converted, whereas with much less reactive bituminous coals, only the smallest char sizes can compete for the available O2. With the least reactive low volatility coals, ignition may become problematic. Most important, the rank dependence for char oxidation reactivity determines the mode of burning which, in turn, determines which portions of the entrained char PSD make the greatest contribution to flyash LOI. With the most reactive coals, all sizes ignite in quasi-steady combustion at temperatures several hundred degrees above local gas temperatures. These burning rates are mostly governed by film diffusion with appreciable mediation by internal pore diffusion. The main mitigating factors are (i) thermal annealing, which can partially extinguish the fully ignited state; and (ii) for chars with high ash loadings, the diffusional resistance through ash encapsulation layers. When both mitigating factors come into play, the largest char particles make the greatest contributions to LOI. With coals of intermediate reactivity, the smallest sizes


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in the char PSD never ignite to a fully developed state, but their chemically limited burning rates are still sufficient to consume all their combustibles. Whereas the larger sizes will fully ignite, the strong size variation in transit times across a riser compensates for the size dependence in the burning rates, so that intermediate sizes in the PSD make disproportionate contributions to LOI. With the least reactive chars, the smallest sizes never ignite and their intrinsic burning rates are too slow to completely burn out the combustibles. So the smallest char and even char fines contribute more to LOI than the largest coarse char. In fact, except for the largest particles, the entire char PSD from the least reactive coals contributes to LOI. The unifying phenomenon for LOI is the potential for ignition to a near-diffusion-limited burning rate across the char PSD. LOI levels increase whenever the ignition is suppressed in any portion of the char PSD, and when fully ignited states are extinguished as char moves through the riser. Ignition is suppressed by low intrinsic oxidation reactivities, starting from the smallest char sizes. Fully ignited particles may be extinguished by thermal annealing with any coal, and by ash encapsulation with coals with excessive ash levels. Thermal soaking at 850°C for a few seconds diminishes char reactivity by nearly two orders of magnitude due to thermal annealing, and the annealing rate surges after ignition due to the much hotter char particle temperatures. But thermal annealing does not seem to undermine the potential to manage LOI by recirculating the largest unconverted char particles through multiple passes through the CFBC, because the annealing occurs upstream of the riser even on the first pass. Obviously, char oxidation mechanisms that impose shrinking core behavior across the char PSD cannot possibly depict the required range of behavior to accurately predict LOI levels from diverse coals. Rather, the burning mechanism must automatically shift from reaction control


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through mediation by pore diffusion to film diffusion limited burning, and also factor in thermal annealing and the transport resistance for ash encapsulation. 6. Conclusions Our simulation results support the following conclusions: 1)

Emulsion gases in dense bottom beds cannot accommodate all the volatiles from even

those coals with the lowest volatility, so they admit no primary air and are composed of volatiles with small amounts of their partial oxidation products. 2)

Due to the relatively slow exchange of O2 into emulsions, extents of volatiles conversion

rarely exceed a few percent and extents of char conversion are even smaller because their burning rates are much slower than the noncondensable fuels’. 3)

Most dense bottom beds convert the coal feed into mostly char fines plus a char

suspension whose mean size is much smaller than the nominal coal size, and should be recognized as highly effective grinding devices that micronize most of the coal feed before it burns. 4)

The intrinsic char oxidation reactivity determines whether or not char fines and coarse

char can effectively compete with noncondensable fuel mixtures for the available O2 at lower elevations, and thereby influence NO production. More reactive chars compete more effectively. 5)

Char burning mechanisms determine which segment of the char PSD contributes to LOI,

in conjunction with size variations in the transit times of char particles across CFBC risers.


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Relevant burning mechanisms comprise chemical reaction control, mediation by pore transport and film diffusion, thermal annealing, and ash encapsulation. Acknowledgement This study was sponsored by the Coal & Environment Laboratory of Idemitsu Kosan Co. Ltd. Nomenclature ai

Particle radius of char fines, coarse char, ash, inerts, and sorbent, cm

Abed Cross sectional area of the dense bottom bed, m2 Anoz Aggregate area of all air injection nozzles, m2 Ar

Archimedes number defined in Table 1

Dbub Mean bubble diameter along the bottom bed, m Deff Effective diffusivity for transport out of bubbles, cm2/s dP

mean size of bed particles, cm

DRISER

Characteristic dimension of a riser cross section, m

FC,j Mass flowrate of char in stream j, kg/s FE,i Mass flowrate of gaseous species i in any stream entrained into a process stream, kg/s Ff

Mass flowrate of char fines, kg/s

FP,i Mass flowrate of gaseous species i in the process stream, kg/s


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FS0 Mass flowrate of soot from coal devolatilization and secondary volatiles pyrolysis, kg/s g

Acceleration due to gravity, cm/s2

h

Elevation along the bottom bed, m

Hi

Elevation above the top of a dense bottom bed, m

H0

Height above a bottom bed where the annular particle layer starts to thin, m

Hbed Height of the dense bottom bed, m Kbe Exchange coefficient from bubbles to emulsion, s-1 mC Mass of a char particle, g MC’ Molecular weight based on a char’s elemental composition, g/mol Mi

Molecular weight of species i, g/mol

PC,j Frequency distribution function for char size Remf Reynolds number for minimum fluidization Ret Reynolds number based on a particles’s terminal velocity t

Time, s

tWALL

Thickness of the annular layer of particles along a riser wall, m

Ubr Superficial rise velocity of bubbles, cm/s UBUB Superficial gas velocity through the bubble phase, cm/s


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Umf Superficial gas velocity for minimum fluidization, cm/s Ut

Terminal velocity of a particle, cm/s

U0

Superficial gas velocity through the bottom bed, cm/s

vb* Superficial speed of gases through bubbles, cm/s vi

Actual gas velocity across a splash zone, riser, or exit duct, cm/s

WC,DB

Holdup of char in the dense bottom bed, kg

XASH Weight fraction of ash in char XC(af)

Extent of burnout of char fines

XS

Extent of soot burnout

yC

Mass fraction of char in bed solids

yE,i Mass fraction of gaseous species i in any stream entrained into a process stream yP,i Mass fraction of gaseous species i in the process stream Greek Symbols δBUB Volume fraction of the bubble phase εmf Bed voidage at minimum fluidization εbed Voidage in the dense bottom bed


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εS,S Volume fraction of solids in the dense bottom bed φbed Mean sphericity of bed particles γC

Volume fraction of solids in wakes per unit bubble volume, cm-3

ηC,j Mean extent of burnout across a char PSD in stream j κC

Elutriation constant

µ

Mean gas viscosity

νCi

Stoichiometric coefficient for species i for char oxidation

νSi

Stoichiometric coefficient for species i for soot oxidation

ωi

Molar production rate of species i per unit volume by homogenous chemistry, mol/cm3

ρi

Density of particles or gases in a bed, g/cm3

Mean bulk density of solids in a dense bottom bed, g/cm3

τS

Transit time through a splash zone, s

Ψ

Throughflow fraction of gases through the bubble phase

Subscripts 0

Coal feedstream

BA Bottom ash


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bed Dense bottom bed BUB Bubble phase C

Coarse char

F

Char fines

S

Splash zone

R

Circulating ash return

E

Ejection stream from dense bottom bed

Superscripts j

Index on a CSTR in the reactor analysis

5. References 1. Adanez J, Gayan P, Grasa G, de Diego L F, Armesto L, Cabanillas A. Circulating fluidized bed combustion in the turbulent regime: modelling of carbon combustion efficiency and sulphur retention. Fuel 2001, 80 1405-14. 2. Huilin L, Guangbo Z, Rushan B, Yongjin C, Gidaspow D. A coal combustion model for circulating fluidized bed boilers. Fuel 2000, 79 165-72.


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3. Krzywanski J, Czakiert T, Muskala W, Sekret R, Nowak W. Modeling of solid fuels combustion in oxygen-enriched atmosphere in circulating fluidized bed boiler. Fuel Process. Technol. 2010, 91 290-95. 4. Hannes J, Renz U, van den Bleek C M. The IEA model for circulating fluidized bed combustion. Fluidized Bed Combust, ASME 1, 1995. 5. Scala F, Salatino P. Modelling fluidized bed combustion of high-volatile solid fuels. Chem. Eng. Sci. 2002, 57 1175-96. 6. Selcuk N, Ozkan M. Simulations of circulating fluidized bed combustors firing indigenous lignite. Int. J. Thermal Sci. 2011, 50 1109-15. 7. Gungor A. Analysis of combustion efficiency in CFB coal combustors. Fuel 2008, 87 108395. 8. Atsonios K, Nikolopoulos A, Karellas S, Nikolopoulos N, Grammelis P, Kakaras E. Numberical investigation of the grid spation resolution and the anisotropic character of EMMS in CFB multiphase flow. Chem. Eng. Sci. 2011, 66 3979-90. 9. Zhou W, Zhao C S, Duan L B, Qu C R, Chen X P. Two-dimensional computational fluid dynamics simulations of coal combustion in a circulating fluidized bed combustor. Chem. Eng. J. 2011, 166 306-14. 10. Li J, Cheng C, Zhang Z, Yuan J, Nemet A, Fett F N. The EMMS model – its application, development and updated concepts. Chem. Eng. Sci. 1999 54 5409-25.


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Page 36 of 49

11. Niksa S, Liu G-S, Hurt R H. Coal conversion submodels for design applications at elevated pressures. Part I. Devolatilization and char oxidation, Prog. Energy Combust. Sci. 2003, 29(5) 425-77. 12. Glarborg P, Alzueta M U, Dam-Johansen K, Miller J A Kinetic modeling of hydrocarbon/nitric oxide interactions in a flow reactor. Combust. Flame 1998, 115 1-27. 13. Nagle J, Strickland-Constable R F. Oxidation of carbon between 1000 - 2000°C, Fifth Carbon Conf. 1962; Pergamon, Oxford, UK, Vol. 1, pp. 154-64. 14. Pallares, D and Johnsson, F. Macroscopic modeling of fluid dynamics in large-scale circulating fluidized beds. Prog. Energy Combust. Sci. 2006, 32 539-69. 15. Kunii K, Levenspiel O. Fluidization Engineering. Butterworth-Heinemann, Stoneham, MA, 1991. 16. Selcuk N, Deirmenci E et al. Evaluation of an improved code for the performance of AFBCs. J. institute of Energy 1997, 70 31-50. 17. Arena U, D'Amore M, et al. Carbon attrition during the fluidized combustion of a coal. AIChE. J. 1983, 29(1) 40-49. 18. Niksa S, Liu G-S. Advanced CFD post-processing for p. f. flame structure and emissions, 28th Int. Technical Conf. on Coal Utilization and Fuel Systems, Coal Technology Assoc. 2013; Clearwater, Fl.


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Energy & Fuels

19. Luecke, K, Hrtge, E-U, and Werther, J. Modeling of solids and gas mixing effects in large-scale CFB combustors. Proc. 17th Int. Conf. on Fluidized Bed Combust., Paper No., FBC2003-028, ASME 2003; New York, p. 245. 20. Kilpinen, P, Kallio, S, and Hupa, M. Advanced modeling of nitrogen oxide emissions in CFBCs. Proc. 15th Int. Conf. on Fluidized Bed Combust., Paper No. FBC99-0155, ASME 1999; New York. 21. Hannes, J P, Renz, U, and van den Bleek, C M. Mathematical modeling of CFBC in industrial scale power plants. Proc. 14th Int. Conf. on Fluidized Bed Combust., ASME 1997; New York, p. 1151. 22. Amand, L-E and Leckner, B. Influence of air supply on the emissions of NO and N2O from a CFB boiler. Proc. Int. Symp. On Combust. 1992, 24 1407-14. 23. Barletta, D, Marzocchella, A, Salatino, P, Kang, S G, and Stromberg, P T. Modelling fuel and sorbent attrition during CFB combustion of coal. Proc. 17th Int. Conf. on Fluidized Bed Combust., Paper No., FBC2003-065, ASME 2003; New York, p. 341. 24. Johnsson, F and Leckner, B. Vertical distribution of solids in a CFB furnace. Proc. 15th Int. Conf. on Fluidized Bed Combust., Vol. 1, ASME 1995; New York, p. 671. 25. Knobig, T, Werther, J, Amand, L E, and Leckner, B. Comparison of large- and smallscale CFBCs with respect to pollutant formation and reduction for different fuels. Fuel 1998, 77 1635-42.


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26. Lu, J, Zhang, J, Yue, G, Liu, Q, Lu, X, Lu, J, Zhao, Z, Liu, Y, Yu, L, Lin, X, and Li, Z. The progress of the water cooled separator CFB boiler in China. Proc. 15th Int. Conf. on Fluidized Bed Combust., Paper No. FBC99-0038, ASME 1999; New York. 27. Fang, M, Luo, Z, Li, X, Wang, Q, Shi, Z, Ni, M, and Cen, K. Development of the first demonstration CFB boiler for gas and steam cogeneration. Proc. 14th Int. Conf. on Fluidized Bed Combust., Vol. 2, ASME 1997; New York, p. 663. 28. Boemer, A, Braun, A, and Renz, U. Emission of N2O from four different large scale CFBCs. Proc. 12th Int. Conf. on Fluidized Bed Combust., Vol. 1, ASME 1993; New York, p. 585. 29. Yang H, Yue G, Xiao X, Lu J, Liu Q. 1D modeling on the material balance in FCB boiler. Chem. Engr. Sci. 2005, 60 5603-11. 30. Zhao X, Lu J, Yang J, Zhang Q, Dong F, Yu L, Yang Z, Yue G. Operational performance and optimization of a 465t/h CFB boiler in China, Proc. 18th Int. Conf. on Fluidized Bed Combust., ASME 2005; New York, p. 791. 31. Nowak, W, Bis, Z, Laskawiec, J, Krzywoszynski, W, and Walkowiak, R. Design and operation experience of 230 MW CFB boilers at Turow Power Plant in Poland. Proc. 15th Int. Conf. on Fluidized Bed Combust., Paper No. FBC99-0122, ASME 1999; New York. 32. Mirek, P, Czakiert, T, and Nowak, W. NOX emission reduction by the optimization of the primary air distribution in the 235 MW CFB boiler. Proc. 20th Int. Conf. on Fluidized Bed Combust., Eds. G. Yue, H. Zhang, C. Zhao, Z. Luo, Tsinghua Univ. Press 2009; Beijing.


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Energy & Fuels

33. Niksa S. The chemical structure of dense bottom beds in coal-fired CFBCs. Proc. 2013 Int. Conf. on Coal Sci. and Technol., IEA 2013; University Park, PA.


39

Energy & Fuels

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Page 40 of 49

Table 1. Definitions for the hydrodynamic variables in dense bottom beds. Variable Ar

Defining Equation

 ρ g (ρ S − ρ g )g  Ar = d P3   µ2   U mf d p ρ g

[

= (27.2) + 4.08 × 10 −2 Ar

]

1/ 2

Remf

Re mf =

Umf

U mf =

εmf

ε3 −

Dbub

H  Dbub = 0.54(U o − U mf )  bed + 4 A noz  g −0.2 , in m/s  2 

µ

2

− 27.2

µ Remf d p ρg

150 Re mf

φ B Ar 2

(1 − ε ) −

1.75 Re mf

φ B Ar

=0 0.8

Ψ

0.4

Ψ ( h) = f CFB ( h + 4 Anoz ) 0.4

[

]

Ubub

1 .35 , in m/s U bub = 1 .6 (U O − U mf ) + 1 .13 D bub D bed

δbub

δ bub =

vb*

vb* =

εbed

SAT ε bed = δ bub + (1 − δ bub )ε mf ≤ ε bed

Kbe

K be

Ψ(U o − U mf ) U bub

U o − U mf

δ bub

      1 1 = +  1/ 2  D eff 1 / 2 g 1 / 4  U mf   D eff ε mf U br   + 5.85  6.77   4. 5   5/ 4 3 D bub  D bub   D bub   

−1


40

Page 41 of 49

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

Table 2. Cases in the validation database. Label A1 A2 A3 A4 A5 B C D1 D2 D3 E F1 F2 G H1 H2 H3 I J K

Rating 12 MWth 12 MWth 12 MWth 12 MWth 12 MWth 22 MW 22 MW 72 MWth 72 MWth 72 MWth 105 MWth 120 MWth 120 MWth 229 MWth 135 MW 135 MW 135 MW 135 MW 230 MW 235 MW

Coal hv bit hv bit hv bit anth lv bit hv bit hv bit hv bit hv bit hv bit hv bit hv bit hv bit brown hv bit hv bit hv bit lv bit brown brown

Citation Luecke et al.19; Kilpinen et al.20 Hannes et al.21; Amand and Leckner22 Barletta et al.23; Johnsson and Leckner24 Luecke et al.19 Knobig et al.25 Lu et al.26 Fang et al.27 Boemer et al.28 Boemer et al.28 Boemer et al.28 Boemer et al.28 Boemer et al.28 Boemer et al.28 Boemer et al.28 Yang et al.29 Yang et al.29 Yang et al.29 Zhao et al.30 Nowak et al.31 Mirek et al.32


41

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Page 42 of 49

Table 3. Predicted partitioning of volatiles and solids in dense beds and splash zones. Label A1 A2 A3 A4 A5 B C D1 D2 D3 E F1 F2 G H1 H2 H3 I J K

V∞ VEMUL daf wt.% % 39.6 22.6 43.8 18.8 39.3 23.0 7.8 81.8 41.6 16.1 49.9 22.1 37.8 10.3 41.7 24.6 42.3 23.2 44.1 22.3 28.4 22.7 44.3 21.1 45.7 21.5 62.9 7.1 35.1 20.2 35.4 19.2 35.4 24.7 28.8 19.7 60.8 10.7 64.8 10.0

XVOL % 1.1 1.0 1.1 5.7 10.4 2.1 1.4 2.3 2.0 2.0 1.4 1.7 1.8 1.1 1.3 1.2 1.6 1.4 1.4 1.0

FSP,fin kg/h 492 491 417 90 517 2482 4751 4428 4188 7034 4827 11714 9522 10300 15887 21014 16808 22569 54659 44206

FSP,C kg/h 271 167 260 994 209 8950 1584 1254 1038 1957 3872 3253 4333 2426 31421 26970 20930 36393 12586 7300


42

Page 43 of 49

40 35 30

LOI, wt. %

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

25 20 15 10 5 0 0

10

20

30

40

50

Coal Index I, g/MJ

Figure 1. Assigned flyash LOI () for all CFBCs in the validation database based on the curve reported for commercial Chinese CFBCs29.


43

Temperature

Case E

Transit Time

2.5 890 2.0 880 1.5

Gas Flow

1.0

Page 44 of 49

100 900

870

BO for Char, Fines & Soot, %

3.0

Gas Temperature, C

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Transit Time, s & Fractional Gas Flow

Energy & Fuels

Char Fines

Char Soot

80

60

40

20

0.5 860 0.0

Case E 0

0

5

10

15

20

0

Elev. Above Distributor, m

5

10

15

20

Elev. Above Distributor, m

Figure 2. (left) Operating conditions for CFBC E and (right) predicted burnout profiles for soot, char fines, and char.


44

Page 45 of 49

100

80

20.0m

8

15.1

7

60

9.2 40

5.7

Transit Time, s

Exit

Char Burnout, %

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

Case E

Exit

20.0m 15.1 9.2

6

5.7

5

Splash 4 3 2

20

1

Case E

Splash 0

0 0

100

200

300

400

500

600

0

100

Initial Coal Diameter,µm

200

300

400

500

600

Initial Coal Diameter,m

Figure 3. (left) Extent of char burnout vs. initial char particle size from the splash zone through the exit duct; and (right) Char transit times to the indicated elevations along CFBC E.


45

Energy & Fuels

100

34 µm

1300

80

Particle Temperature, C

300 µm Char Burnout, %

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 46 of 49

560 µm

60

40

1200

300 µm 1100

1000

560 µm 20

900

34 µm

Case E

Case E

800

0 0

2

4

6

8

0

2

4

6

8

Transit Time, s

Transit Time, s

Figure 4. (left) Fractional char burnout vs. transit time for three sizes of char particles from the outlet of the splash zone through the exit duct; and (right) Particle temperatures during these combustion histories along CFBC E.


46

Page 47 of 49

1

1.0

1.0 1300

1100

TP 0.4

1000

η 0.2

0.1

Chi 0.6

Chi

0.6

1200

k/k0

d/d0


0.8

0.8

d/d0 & η

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

0.4

k/k0

0.01

0.2

900

Case E

Case E 0.0 0.0

0.5

1.0

1.5

2.0

2.5

800 3.0

1E-3 0.0

0.5

1.0

1.5

2.0

2.5

0.0 3.0

Transit Time, s

Transit Time, s

Figure 5. (left) Combustion history for char from 297µm coal particles including particle temperature, scaled particle size, and effectiveness factor; and (right) the associated decay in the char oxidation reactivity and the Chi parameter.


47

Energy & Fuels

1150

100 100 1100

60

Splash

40

80

Char Burnout, %

Exit 20 m 15.9 9.2 5.0

1050

60

1000

950

40

340 µm 900

20

Particle Temperatures, C

115 µm 80

Char Burnout, %

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 48 of 49

20 850

Case J

0

Case J

0

800

0

100

200

300

400

0

1

2

Initial Coal Diameter, µm

3

4

5

6

7

8

Transit Time, s

Figure 6. (left) Extent of char burnout vs. initial char particle size from the splash zone through the exit duct; and (right) Char burnout and particle temperature histories for 115 and 340µm char along CFBC J which fires brown coal.


48

Page 49 of 49

100

100 1200

Exit

80

60

Char Burnout, %

10.3 m 7.6 5.1 40

3.0

760 µm 1100

60 1000 40

20

900

60 µm

1.8 20

Splash Case A4

0

Case A4

800

0 0

150

300

450

600

750

0

1

Initial Coal Diameter, µm

2

3

4

5

6

Transit Time, s

Figure 7. (left) Extent of char burnout vs. initial char particle size from the splash zone through the exit duct; and (right) Char burnout and particle temperature histories for 115 and 340µm char along CFBC J which fires anthracite.

AUTHOR INFORMATION

Corresponding Author * Stephen Niksa* Niksa Energy Associates LLC, 1745 Terrace Drive, Belmont, California, 94002 USA 650 654 3182; 650 654 3179 (fax); [email protected]


49


80

Char Burnout, %

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

Predicting the Conversion Efficiencies of Any Coal Type in CFBCs

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