Coal combustion studies in a fluidized-bed test facility

Coal combustion studies in a fluidized-bed test facility

Energy Vol. 17, No. 6, pp. 579-591, 1992 Printedin Great Britain. All rights reserved 0360s5442/92 $5.00 + 1).00 Copyright 0 1992 Pergamon Press plc ...

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Energy Vol. 17, No. 6, pp. 579-591, 1992 Printedin Great Britain. All rights reserved

0360s5442/92 $5.00 + 1).00 Copyright 0 1992 Pergamon Press plc



of Chemical



S. RAO, V. G.


and R. R.


The University of Illinois at Chicago. P.O. Box 4348, Chicago, IL 60680,



13 June 1991)

Abstract-A fluidized-bed combustor has been designed, constructed, installed, and tested for the combustion of coal. A cylindrical (0.152m i.d.) fluidized-bed, equipped with a specially designed propane burner to preheat the inert bed of sand and ignite an Illinois Basin coal sample (IBC-103), was employed in the test runs. The flue-gas composition is established by using a Cole-Parmer KM 9(!04 electronic combustion analyzer. A methodology is developed to compute the percentage utilization efficiency of carbon and the fuel-combustion efficiency, based on knowledge of fuel and flue-gas compositions and operating conditions of the combustor. Thirteen combustion-test results are reported, emphasizing the quality and efficiency of c~)mbustion for different coal-feed rates (0.66-1.58 g/s), bed temperatures (1003-1216 K), fractional excess air values (0.48-1.521, and gas-fluidization numbers (1.25-3.26). In addition. the hydrodynamic fluidization quality of the bed has been investigated on the basis of a series of measurements relating the calming-section pressure drops through the gas-distributor plates, pressure drops across the total bed and across its two sections, and the pressure-drop fluctuation history for various temperatures as unctions of the ~uidizing air velocity. These data provide bed-property parameters such as the minimum fluidization velocity, bed voidage, variance in bed pressure-drop fluctuation, and variations of these properties with operating conditions. The results are compared with limited related experimental data in the literature.


technology has been extensively investigated for the combustion of sulfur bearing and low rank coals. In this process, sulfur dioxide produced during combustion reacts with calcined dolomite or calcium carbonate present in the bed and thus the flue gas is almost free of this undesirable pollutant. In order to assure efficient occurrence of this heterogeneous gas-solid reaction in a fluidized bed environment it is essential that good gas-solids contacting and solids mixing is accomplished. We have been engaged in a program of research to establish diagnostic probes which can reliably establish the hydrodynamic fluidization quality of the bed. Saxena and Rao’ described a thermocouple temperature probe and showed how a temperaturehistory record of a probe immersed in a fluidized bed can be interpreted to establish the fluidization quality. Later these workers2 also developed a transducer pressure probe and demonstrated that a pressure-history record also is an effective way to delineate the different fluidization regimes, which may prevail in the bed with changing operating parameters. These techniques and procedures in turn can be employed to ensure that a fluidized-bed coal combustor is being operated under favorable conditions so that the desired hydrodynamic activity predominates in the reactor. Saxena et al3 extended the scope of these experimental techniques and data analyses to high temperatures (913-1173 K) by conducting experiments on two sand beds with average diameters of 641 and 1312 pm. The coal-combustion experiments were performed in a 0.152 m i.d. fluidized-bed combustor. The combustor is an extensively modified version of an earlier Fluidized-bed

t To whom ail correspondence $:Present



should be addressed. of Chemical



India. 579


of Technology,




S. C. SAXENA et al


design4 and combustion. entering the experiments parameters.

is now equipped with a propane burner to preheat the inert bed and initiate coal This gas sparger is installed in the calming section and it preheats the air before bed section. The experimental facility is described in the next section. A series of has been conducted on coal combustion under a variety of changing operating



The pilot plant includes air and propane gas-supply and metering systems, the fluidized-bed coal combustor, a coal feeder, a flue-gas cleanup system, and a combustion air preheater arrangement, as described by Saxena et al.‘T4 Extensive modifications incorporated in the

Fig. 1. Schematic of the fluidized-bed combustor. 1, tube-fired propane burner; 2 and 3, preheating sections; 4 and 5, elbows; 6, calming section; 7, perforated plate air distributor; 8, auxiliary propane sparger; 9, test-bed section; 10, multi-nozzle air distributor; 11, continuous feed pipe; 12, batch feed pipe; 13, overflow pipe; 14, freeboard section; 15, cooling coils; 16, auxiliary air inlet. All dimensions are in mm.


coal combustion



reactor and other units are described here. A flue-gas analyzer system has been added to determine the concentrations of 02, CO and COZ. The schematic and engineering design details of the fluidized-bed reactor are shown in Fig. 1. The main test bed and freeboard sections of the combustor consist of a 0.254 m dia, schedule 40, 304 stainless-steel pipe that comprised of four flanged sections, each 0.571 m long and bolted together with 0.038 m long spacer rings. Below these sections is a 0.305 m long calming section which communicates with the air preheating section through two elbows. The overall height of the two preheating sections is 1.52 m and it is provided with a tube-fired propane burner which is not used in the experiments reported here. A specially designed auxiliary propane burner in the calming section is used to preheat the fluidizing air. The entire preheater, elbows, calming, bed and freeboard sections are provided with a 50 mm thick Purolite-30 insulating material lining. Four external band heaters, type T, two-piece elements, 273 mm i.d., 51 mm wide, are wrapped around the test-bed section of the furnace above the distributor plate. These heaters are connected to three independent electrical circuits and an automatic temperature controller with thermocouple feedback is employed to energize these heaters and to maintain the column wall at a specified temperature. The combustor wall is externally insulated with Fiberflax Lo-Con felt insulating wrapping of 25 mm thickness. Design details of the propane burner are shown in Fig. 2; this burner is installed in the calming section. It consists of a central stainless steel housing made out of a 25.4 mm cylindrical rod with a 12.7 mm axial blind cylindrical cavity. Symmetrically attached to this cavity are four stainless steel pipes, 9.5 mm o.d., 3.18 mm i.d., and 63.5 mm long with their ends sealed with welded stainless steel plugs. Each of these four arms has two orifices, 0.5 mm dia, and located at distances of 25.4 and 50.8 mm from the center of the central support housing. One such orifice is also made at the center of the central support housing. To introduce propane in the sparger, a 3.18 mm dia pipe is attached to the central housing as shown in Fig. 2. The propane thus introduced distributes through these nine orifices into the calming section. Experiments revealed that better mixing of air and propane is achieved when the orifices face directly the calming section perforated gas distributor plate so that the incoming air stream meets

I /0”

Special Flttlng








Fig. 2. Design



of the 304 stainless steel propane gas-sparger unless specified otherwise.



All dimensions

are in mm



et al

Three layen of equal& spaoed 1.5 mm holes (27)


k 6.3 T Dlstrlbutor / Arc Snap Rtng

4 12.7w w I9 4

AlI dlmenslons

i Plate

are In mm

Fig. 3. A cylindrical conical-head multi-orifice nozzle (A) and the arrangement gas-distributor plate (B).

of the nozzles on the

counter-currently the propane stream. A spark plug is provided above the propane sparger in the calming section to ignite the air-propane mixture. A bank of four propane tanks suitably manifolded provided a regulated supply of the gas and is metered by two rotameters before entering into the sparger. The bed distributor plate has 19 multi-orifice nozzles. Design details of the single nozzle are shown in Fig. 3A, whiie the layout of all nozzles on the distributor plate is shown in Fig. 3B. The nozzle is made out of a 19 mm dia stainless steel cylindrical rod, having a 12.7 mm dia blind axial hole. It has a conical taper at its top end to avoid settling of solids. Air from each nozzle exits from twenty-seven 1.5 mm dia holes drilled at an inclination of 70” from the cylinder axis. These holes are located in three layers around the circumference of the cylinder with nine equidistant holes in each layer. The hole axis is directed to intersect the cylinder axis, The lowermost layer is located at a distance of 60.4 mm from the top of the distributor plate. The vertical separation between consecutive layers is 6.3 mm. The inclination of the holes prevents flow of solids into the nozzle and thence into the calming section. The centerlines of holes are offset by 13.3” between adjacent layers. This ensures a uniform distribution of air around the periphery of each nozzle. Nineteen nozzles are mounted on a 6.3-mm thick stainless-steel distributor plate with two arc-snap rings located on either side of the plate. Each of these nozzles have two grooves, 18 and 36 mm from the top surface to the distributor plate, and these provide additional grip when the nozzles are cast in a 50.8 mm thick refractory layer. The lay-out of the nozzles on the distributor plate is shown in Fig. 3B. These are arranged on the circular plate with one nozzle at the center and the remaining 18 nozzles in two concentric circles of diameters 69 and 126 mm. The inner circle accommodates six nozzles at equal separation while the remaining 12 are located on the outer circle. This geometric configuration of the nozzles on the distributor plate is arrived at by applying the criterion that the plate regions bounded by the base area shown as 1, 2, and 3 in Fig. 3B are supplied with the same air volume per unit area, which will, in turn, ensure uniform gas flow through the cylindrical bed. The calming section air distributor plate has 61 holes of 3.2 mm dia.’ The IBC-103 coal is fed into the combustor by an Acrison Model 105B solids feeder which is modified by incorporating a new 304 stainless steel metering auger (3/4-in. o.d. and 16 and l/8-in. long) housing cylinder and a downspout for smooth flow of solid particles up to 6 mm in diameter. The coal feed rates could be typically varied up to 10 kg/h by a single knob control


coal combustion studies


and a linear relationship is found between the feed rate and the auger speed (revolution per minute) as read on a dial. At the top end of the freeboard, a 96 mm gas sample probe is installed which communicates with the Cole-Parmer KM9004 electronic combustion analyzer. It monitors flue stack levels of oxygen in percentage, and carbon monoxide in ppm, and the temperature of the flue gas. The instrument is also capable of giving calculated carbon dioxide composition based on measured oxygen concentration for a given fuel of known carbon weight fraction. The unit has a display and a printer which will provide the print out on demand.



In order to ensure uniform air distribution and good fluidization, the pressure drops (APc and AP,) are measured across the calming section distributor and the bed distributor plate, respectively, as functions of increasing air superficial velocity U. The regression analysis of these data yielded the following relationships: AP,(Pa) = 205.2U’.62 AP,,(kPa) = 2.019U’,90 APo(kPa) = 3.707iJ’.96

for 0 s U(mls) G 2.24, after installation, after 40 h of operation.


From these equations, it is clear that both distributor plates have flow conditions approximating orifice flow.6 A series of experiments has been conducted with a sand of average dia 1312 pm, its size distribution is reported earlier,3 at ambient and higher temperatures. Higher temperatures are obtained by either propane and/or coal combustion. Pressure drop and pressure-drop history records are taken for two sections of the bed at each temperature as a function of fluidizing gas velocity. One lower section refers to the bed section contained between 4 and 14 cm above the gas distributor plate, and the other upper section to the bed region confined between 14 and 29 cm above the distributor plate. A typical slumped bed height in these experiments is about 35 cm. The pressure-drop data are employed to compute the voidages of the lower (sL) and upper (E”) bed sections as a function of gas velocity. The pressure-drop histories are recorded for a period of 92 s at each temperature and fluidizing gas velocity. These data are employed to compute the variance (at), probability density function (pdf), and power spectral density function (psdf) with a view to characterize the bed quality fluidization. For the calculation of coal or thermal combustion efficiency, it is essential to continuously monitor the steady state compositions of 02, CO, and C02. Thermal combustion efficiency of a fluidized-bed coal combustor is defined as the ratio of heat liberated to the heat input in a certain time period. For a carbonaceous material primarily containing carbon and hydrogen (such as coal) the difference in the rate of heat liberated and the rate of heat input is due to several factors. These are: incomplete combustion of C to CO. incomplete combustion of volatile components, and unburnt C that remains in the bed or in the entrained and overflow streams. The parameters influencing the combustion efficiency fall in three general categories6 viz. fuel characterization parameters, operating parameters, and equipment parameters. For a given fuel and combustor, the operating parameters which influence the combustion efficiency are: total air or excess air, bed temperature, feed rate (carbon loading), bed particle size, bed depth etc. In this work a bed depth of 35 cm is employed which changes nominally by the feed rates and combustion of coal over the period of our experimentation ranging up to 3 h for a given feed rate. Therefore the important parameters for our present discussion are: total and excess air, bed temperature and coal feed rates. The total and excess air is represented here in terms of the superficial fluidizing air velocity at ambient conditions. The total combustible heat loss at given operating conditions can be computed if the heat losses associated with elutriated

S. C.



et al

unburnt C and that associated with the amount of CO present in the flue gas may be established. We have estimated this heat loss by making C balance and employing knowledge of flue-gas composition. This methodology is described below. Let C, H, 0, N, and S be the mass fractions of carbons, hydrogen, oxygen, nitrogen, and sulfur, respectively, in the feed. Further, let A and B be the mass fractions of unburnt and burnt carbon, respectively, in the fuel. Then, A+B=C.


We further define P=

C converted C converted to CO B C converted to CO + CO2 =

to CO


Taking unit mass of the fuel as basis, to CO = PB,


to CO* = (1 - P)B,


C converted C converted


mass of CO, in the flue gas = (44/12)(1-

mass of CO in the flue gas = (28/12)PB,



O2 consumed to produce CO, + CO = (2.666 - 1.333P)B.


Assuming that H, N, and S present in the fuel are completely converted to HzO, NO, and SO*, respectively, O2 consumed = (16/2.02)H


+ (16/14)N + (32/32)S =X1,


SO2 produced = (64/32)S,


NO produced = (60/28)N.


the total O2 required for stoichiometric


of fuel =

(32/12)C + (16/2.02)H + (16/14)N + (32/32&S - 0 =X2.


Let 2 be the fractional excess air supplied, which is defined as the excess air divided by the stoichiometric air. Therefore, O2 supplied = X2(1 + Z),


O2 consumed during combustion = (2.666 - 1.333P)B + X,,


mass of N2 in the flue gas = (79/21)(28/32)(1+


mass of O2 in the flue gas = X2(1 + Z) - [(2.666 - 1.333P)B +X1].

(17) (18)

Let F be the mass of dry flue gas per unit mass of fuel. Then F = (44/12)(1-

P)B + (28/12)PB + (64/32)S + (60/28)N

+ (79/21)(28/32)(1+ = B + 4.312(2.666C

Z)X, +X2(1 + Z) - [(2.666 - 1.333P)B +X1], + 7.92H + 1.143N - 0 + S) + 11.49C + 26.22H

+ 5.31s + 5.927N - 4.410.


The amount of dry flue gas can also be estimated from the flue gas composition. The flue gas flow rate and composition are not appreciably influenced if we neglect the presence of SOP and NO in the flue gas. Hence the flue gas may be taken as consisting of CO, C02, Nz and Oz. Let Y be the mass of dry flue gas per unit mass of C burnt in the fuel. Then, y = 44[COz] + 32[02] + 28[CO] + 28[N2]

WC01 + [CW



The square brackets represent gas. Recalling that

coal combustion studies


the volume fraction of the particular chemical species in the flue

[CO1+ [%I+ [C&l + [@I= 1,


Eq. (20) simplifies to Y = {4[C0,1+

[O,] + 7]/3{[CO] + [CO,]}.


Mass of dry flue gas per unit mass of the fuel is = F = Y(C -A).


From Eqs. (4), (19) and (23), the fraction of C burnt, B, can be written as follows: B = [4.312(2.666C

+ 7.92H + 1.143N + S - 0) + 11.49C

+ 26.22H + 5.9271\3- 4.410 + 5.31S]/(Y - 1).


Knowing B from Eq. (24), the fraction of unburnt C, A, may be computed from Eq. (4). Let Ma, be the average molecular weight of the flue gas which can be calculated from the knowledge of flue gas composition, and is given by the following relation: M,, = 44[COJ + 28[CO] + 28[NJ + 32[02]. Moles of the flue gas per unit mass of the fuel = F/M,,

(25) = Iv&.


Number of moles of CO in the flue gas per unit mass of the fuel = M,[CO]. Heat lost in CO per unit mass of the fuel = AH,,M,[CO]. Here AH,-, is the heat of combustion

(27) (28)

of CO and is 67,636 kcal/kg mol.

Heat lost due to unburnt C per unit mass of the fuel = AH,(A/12.0).


Here AH, is the heat of combustion of C and is 94,052 kcal/kg mol. Percentage Total combustible

C utilization efficiency = qcu = (B/C)lOO.


heat loss per unit mass of the fuel = Q,. = Eq. (28) + Eq. (29).


fuel combustion

efficiency = nCE = [(Qr - QL)/QF]lOO.


Here, QF is the heating value of the fuel and is known. For IBC-103 coal, the reported heating value is 13,432 Btu/lb (7455 kcal/kg). Percentage of heat lost through incomplete combustion of CO to CO2 in relation to the total heat lost (QL) may be computed from



In the computation of nCE from Eq. (31), for the case of mixed fuel a special explanation is needed. QF for the mixed fuel is computed in proportion to their weight fractions present in the feed. Thus,

QF=WCQFC+(1-KG&w Here QFcand QFPare the heating values of the coal and propane respectively, mass fraction of coal in the mixed fuel feed.


(33) and W, is the


In Fig. 4, cU and cL are plotted at six temperatures as a function of fluidization number U/U,,,,. The agreement among the individual values of .sU and cL at a given temperature is good, irrespective of whether heating is obtained by either pure coal combustion or by combined coal and propane combustion. Hence, it is inferred that the quality of bed fluidization is not influenced by the method adopted to obtain a particular bed temperature. EGY 17:6-E

S. C.





A Upper





: u12pm








E 0.6




.-_-::. .



Sand 1312pm T-233K

CP 0.4,





0.4 5-




n 0






Sand 1312pm T - 1203K







3 tJ/Urnf


0.9 Sand





--Lower -Upper












0.6 :



0.6 ‘=._






1 - 633K 0.4,










sand 1312pm T - 1103K



in bed voidage

3 UIunf


Fig. 4. The variations







of the lower and upper functions of U/V,,.

bed regions

at different



This is understandable because coal forms only a small weight percentage (2-3) of the inert bed. The two sets of bed voidage values are in reasonable agreement with each other at higher temperatures implying that the bed may be regarded as much more uniform in its quality under combustion conditions. In Fig. 5 are presented several typical plots of pressure drop (AP) fluctuation history records over a period of 92 s at a sampling or recording speed of about 11 Hz. The readings refer to the lower bed section and several gas fluidizing velocities at 833 K. Saxena and Rae’ have pointed out that the variation of the variance in the pressure-fluctuation can be regarded as a good indicator of the quality of bed fluidization. This approach is adopted here. The variance cr’, of 1.25






2 x


a Q

0.75 0.25

0.25 0.75



~rcuurlrpuulrrw~ 0.99

0.25 0





60 Time

Fig. 5. Pressure-fluctuation








at 833 K and at 8 different


of U/U,,,,.



coal combustion



the fluctuation signals is defined as

a;= (l/N’)

2 [E(t) - PI’. i=l


Here, c(t) is the ith instantaneous value of the pressure signal at time t and p is the average of such N’ instantaneous values taken over a time period t at a recording speed off Hz. Further, N’ = ft.


up is referred to as the standard deviation of the pressure drop fluctuation data. Computed values of up for the two bed regions as a function of V/V,, at various temperatures are presented in Fig. 6. At ambient temperature the standard deviation is the same for the lower and upper bed regions. As the temperature is increased the variance for the upper bed region has a more pronounced peak, in general, than for the lower bed region. Further, the nature of variation suggests a transition in the fluidization regime. This clearly suggests that in a tall bed the nature of fluidization may change along the bed height. The implications of this on the overall quality of bed fluidization are planned to be investigated in beds of different size particles heated by propane combustion alone. A detailed nature of the bed hydrodynamics and its structure is understood by computing the probability density function and investigating the departure of this distribution from normal distribution in terms of skewness and kurtosis. For an ergodic process, this function gives the probability that the data will have a value within the defined range at a given instant. If N, is the number of data points in the range dr around X, then the probability density function p(x), is given byX p(x) = N:/N’ dx.


The application of pdf to physical data is to establish a probabilistic description of the instantaneous values of the data. The existence of the periodic phenomena can be confirmed by the saddle shape about the mean of the probability density function. The shapes of the pdfs also reveal the extent of interference of the smaller fluctuations superimposed on the major fluctuation. In Fig. 7, the variations of pdf at several gas fluidizing velocities and at 923 K are presented. Similar plots have been generated at 833 and 1033 K The nature of these plots are 0.6 1



U/U,f 0.6



2 a, 5


C? 0.2

Fig. 6. The variations

of standard deviation of pressure-drop fluctuations in the lower and upper regions as a function of U/V,,,,, at different temperatures.


et al





AP &Pa)

M Wa)

Fig. 7. The variations of probability density function at 923 K and at several U/U,,,, values for lower and upper bed sections.

similar and their variations with U/U,,,, are also almost identical. The change in pdf with temperature is insignificant but with increase in air velocity there is a clear increase in the range over which the pressure drop changes. In fact our data not presented here have indicated that such a trend is more pronounced for the upper region than for the lower region. This implies that in the upper region there are larger coalesced bubbles and hence larger scatter in the AP values than in the lower bed region for a given gas fluidizing velocity. The coal combustion runs in an inert bed of sand and in the temperature range 1003-1216 K Table

1. Data relating to combustion


mixed (coal and propane) bed = 1312 pm. _



J’U, d


Fc (g/S)



Tb (K)




mean diameter


I 5

co Wn

Corn ETrsitio LC CO2 (70)

02 (70)




the inert sand



9 CE


1.033 1.189 1.430 1.648

1.88 2.16 2.60 2.99

0.661 0.661 0.788 0.778

0.275 0.344 0.473 0.509

1.16 1.20 1.15 1.36

1003 1023 1023 1018

1476 1151 1191 824

6.66 6.38 6.04 5.61

12.6 12.5 13.2 13.7

0.766 0.770 0.772 0.773

0.702 0.697 0.655 0.671

91.6 90.6 84.8 86.8

93.1 92.7 88.8 90.4

1.196 1.547 1.856

2.17 2.81 3.37

1.016 1.016 1.016

0.086 0.344 0.430

1.44 1.30 1.52

1103 1103 1088

513 405 367

6.99 5.96 5.62

12.8 13.6 13.9

0.751 0.763 0.766

0.744 0.647 0.684

99.1 84.8 89.8

98.9 88.9 92.1

0.712 0.834 1.030 1.196 1.547 1.856

1.25 1.46 1.81 2.10 2.71 3.26

0.788 0.778 1.016 1.283 1.582 1.582

0.215 0.301 0.000 0.232 0.245 0.430

0.48 0.52 1.39 0.69 0.89 0.90

1193 1203 1213 1201 1216 1211

385 357 349 412 430 372

8.12 7.81 6.86 7.37 7.85 7.71

11.0 11.2 13.2 12.1 11.6 11.5

0.761 0.765 0.745 0.756 0.755 0.760

0.550 0.559 0.686 0.557 0.660 0.675

72.4 73.1 92.0 73.7 87.4 88.8

79.8 80.8 93.5 80.3 90.4 91.7



coal combustion



%u 80

Fig. 8. The variations

01015 T .1095 + +1205_1

of carbon fluidization

utilization efficiency (A), number at three different

and fuel combustion bed temperatures.



are summarized in Table 1 and presented in Fig. 8. In each case, the flue-gas composition is measured for CO, O2 and CO2 and these are reported in this table. The general trend evident from these data is that the concentrations of CO and O2 decrease while that of CO2 increases as the combustor temperature increases. The quality of combustion, as represented by the decreasing concentration of CO, improves as the temperature increases, and temperatures above 1000 K lead to good combustion. More specifically, when the temperature is changed from 1018 to 1213 K at excess air level of 1.35 +0.05, the concentrations of CO and O2 decreased and that of CO* increased. These findings are consistent with the reported results of Hampartsoumian and Gibbs,6 and Gibbs and Headley.’ The combustion efficiencies are also found to increase with increase in temperature in conformity with other works.“,9.1” In such data, the level of carbon loading is also important, an increase in carbon loading will decrease the combustion efficiency for otherwise identical conditions. As pointed out by Gibbs and Beer,” increase of carbon loading enhances the rate of particle attrition resulting in greater elutriation loss. This causes the combustion efficiency to decrease. The quality of combustion at a given temperature also depends on the value of excess air, U/U,,,,, and carbon loading or feed rate. The excess air level can be varied either by increasing the fluidizing velocity at a constant coal feed rate, or by changing the fuel feed rate at a constant fluidizing velocity. In Table 1, data at a constant feed rate (1.016 g/s) and temperature of about 1100 K, suggest that increasing excess air decreases the concentrations of CO and CO;?, increases the concentration of 02, and increases the combustion efficiency. These results are again in accord with reported trends. 6.9,“’The data of Table 1 also indicate that under such conditions the amount of unburnt carbon decreases with increase in excess air. This also implies that the carbon loss associated with elutriation decreases with increase in excess air. At a constant temperature and excess air, the effect of fluidizing gas velocity can be observed only by making the corresponding change in fuel feed rate which implies a change in carbon loading. In Table 1, such data are available at a temperature of about 1020 K for excess air level of 1.15. At 1020 K, the concentrations of CO and CO, decrease and the concentration of O2 increases as the fluidizing velocity and carbon loading increase. These observations are similar to those reported by Gibbs and Beer.” The combustion efficiency decreases and the amount of unburnt carbon increases with increase in carbon loading and fluidizing velocity. This as noted earlier is due to increase in attrition with increase in carbon loading. A few interesting comments can now be made about the concentration of CO in the flue gas of CO and its dependence on operating parameters. At about 1023 K, the concentration decreases rapidly as the U/U,,,, value is increased. This decrease is primarily due to increase in excess air which dilutes the CO in flue gas and thereby decreasing its concentration. At about 1103 K, the carbon loading is constant and increase in U/U,,,, values decrease the CO concentration in the flue gas. On the other hand at about 1205 K, the carbon loading increases

S. C. SAXENA et al


as the U/U,,,, values increase and this also changes the level of excess air. The combined effect of these two parameters results in the observed trend of initial increase and then decrease in CO concentration with increase in U/U,,,, values. The variations of CO, and O2 concentrations with U/U,,,, at three different temperatures also should be noted. These trends are of course consistent with our earlier discussion of these gases relating the variations of their concentrations with excess air and carbon loading. There is a clear correspondence between the standard deviation plots of Fig. 6 and the carbon utilization and combustion efficiency plots of Fig. 8. This is expected as the hydrodynamic activity in the bed is related to solids mixing and gas-solids contacting and these in turn are directly related to carbon utilization and combustion efficiencies. In Fig. 6, crP changes its trend of variation around the fluidization number (U/U,,,,) of about 2, signaling a transition from bubbling to turbulent regime.77’23’3 In the turbulent regime, the carbon combustion efficiency is a maximum and a further increase in the fluidization number will only insignificantly influence the bed hydrodynamics and hence the combustion efficiency, as is evident from Fig. 8. It also appears from Fig. 8 that the bed temperature in the range 1000-1200 K has only marginal influence on ncU and nCE in the turbulent fluidization regime. Both the figures together clearly demonstrate that there is a definitive relationship between the bed hydrodynamics and combustion efficiency. The parameter U/U,,,, has a pronounced influence through related changes in bed quality fluidization as compared to some other parameters varied in the present work. Acknowledgemen&-Financial support received from the Illinois Department of Energy and Natural Resources and its Coal Development Board through the Center for Research on Sulfur in Coal is gratefully acknowledged. We are also grateful to the Illinois State Geological Survey for supplying us with the coal sample used in combustion experiments. Cooperation of K. K. Ho is gratefully acknowledged. This work is also supported in part from the office of Solid Waste Research of the University of Illinois at Urbana-Champaign under grant number OSWR-04-002.

REFERENCES 1. S. C. Saxena and N. S. Rao, Energy-The International Journal 14, 811 (1989). 2. S. C. Saxena and N. S. Rao, Energy-The International Journal 15,489 (1990). 3. S. C. Saxena, N. S. Rao, and S. J. Zhou, Energy-The International Journal 15,1001 (1990). 4. S. C. Saxena and A. Mathur, Energy-The International Journal 10,57 (1985). 5. S. C. Saxena and A, Chatterjee, Energy-The International Journal, 4, 349 (1979). 6. E. Hampartsoumian and B. M. Gibbs, “The Influence of Fuel Burning Characteristics on the Performance of a Fluidized-Bed Combustor,” in International Conference on Combustion Engineering, Vol. II, p. 1.55 (1983). 7. S. C. Saxena and N. S. Rao, Energy-The International Journal 16, 1199 (1991). 8. J. S. Bendat and A. G. Piersal, Random Data: Analysis and Measurement Procedures, Wiley, New

York, NY (1986). 9. B. M. Gibbs and A. B. Headly, Proceedings

of the Second



Keairns eds., Cambridge, U.K. (1978). 10. D. P. Naude and R. K. Dutkiewicz, Proceedings

of the Second



of Large Coal Particles in a Fluidized-Bed”, in Conference, p. 235, J. F. Davidson and D. L.

“Fluidization-Bed Foundation

Keairns eds., Cambridge, U.K. (1978). 11. B. M. Gibbs and J. M. Beer, “A Pilot Plant

Combustion of Poor Quality Coal,” in p. 280, J. F. Davidson and D. L.


Study of Fluidized-Bed

Coal Combustion,”


Proceedings of the High Temperature Chemical Reaction Engineering Symposium, pp. 23-l to 23-9, London, U.K. (1975). 12. Y. R. Yang, S. L. HOU, B. C. Chen, and G. T. Chen, Proceedings of the Fourth Asian Congress of Fluid Mechanics, pp. l-4, Hong Kong (19-23 August 1989). 13. S. Anderson, F. Johnson, and B. Leckner, Proc. 1989 Znt. Conf. Fluidized-Bed Combustion, San

Francisco, CA, Vol. 1, p. 239 (30 April-3

May 1989).


coal combustion




A B C dx

= = = =

Mass fraction Mass fraction Mass fraction Increment in

of unburnt C in the fuel of burnt C in the fuel of C in the fuel the pressure-drop value

Th= t= U= U,t =

(Pa) F = Mass of dry flue gas per unit mass of

4. = Fp =

f = H = Mav= Mf =

N= N’ = N: = 0 = P = p = c(t) = p(x) = QF = QF,. = QFp = QL =


fuel Mass flow rate of coal (g/s) Mass flow rate of C,H, (g/s) Frequency (Hz) Mass fraction of H2 in fuel Average molecular weight of flue gas Moles of flue gas per unit mass of fuel Mass fraction of N, in fuel Number of data points Number of data points in the range dr around x Mass fraction of 0, in fuel Mass fraction of burnt C converted to co Average of N’ instantaneous pressuredrop signal values (Pa) ith instantaneous value of pressuredrop signal (Pa) Probability density function (Pa-‘) Fuel heating value (kcal/kg) Coal heating value (kcal/kg) C,H, heating value (kcal/kg) Total heat lost through incomplete combustion of CO to CO2 and unburnt C (kcal/kg) Mass fraction of S in fuel

Wc = x= X, = X2 = Y= Z =

Bed temperature (K) Time (s) Superficial gas velocity (m/s) Superficial gas velocity at minimum fluidization (m/s) Coal mass fraction in mixed fuel Pressure-drop value (Pa) 0, consumed for complete oxidation of H, N and S in the fuel (kg/kg fuel) 0, required for stoichiometric combustion of fuel (kg/kg fuel) Mass of dry flue gas per unit mass of C burnt (kg/kg of C) ratio of excess air to stoichiometric air

Greek letters

AHc = Heat of combustion of C (kcal/kg mol) AH,, = Heat of combustion of CO (kcal/kg mol) AP = Pressure drop (Pa) AP, = Pressure drop across perforated distributor plate (Pa) APn = Pressure drop across multi-orifice distributor plate (Pa) E = Bed voidage Ed = Voidage of lower bed section sU = Voidage of upper bed section rlcE = Fuel combustion efficiency in percent rlcU = Carbon utilization efficiency in percent t = Time period (s) a$ = Variance (Pa’)