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Circulating fluidized bed combustion in the turbulent regime: modelling of carbon combustion efficiency and sulphur retention
Fuel 80 (2001) 1405±1414
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Circulating ¯uidized bed combustion in the turbulent regime: modelling of carbon combustion ef®ciency and sulphur retention J. Adanez a,*, P. GayaÂn a, G. Grasa a, L.F. de Diego a, L. Armesto b, A. Cabanillas b a
Department of Energy and Environment, Instituto de CarboquõÂmica (CSIC), P.O. Box 589, 50080 Zaragoza, Spain b Ciemat. Avda. Complutense 22, 28040 Madrid, Spain Received 18 September 2000; revised 2 January 2001; accepted 9 January 2001
Abstract A model has been developed considering the hydrodynamic behaviour of a turbulent circulating ¯uidized bed, the kinetic of coal combustion and sulphur retention in the riser. The hydrodynamic characteristics of the turbulent ¯uidization regime were integrated together with the kinetic submodels of char combustion and sulphur retention by limestone. From the combustion of a lignite and an anthracite with limestone addition in a hot CBF pilot plant of 20 cm internal diameter and 6.5 m high, the effect of operating conditions such as temperature, excess air, air velocity, Ca/S molar ratio, coal and limestone particle size distributions on carbon combustion ef®ciency and sulphur retention were studied. The experimental results were compared with those predicted by the model and a good correlation was found for all the conditions used. q 2001 Elsevier Science Ltd. All rights reserved. Keywords: Turbulent ¯uidization; Sulphur retention; Combustion ef®ciency
1. Introduction Circulating ¯uidized bed coal combustion (CFBC) with sorbent addition is a well-known technology for achieving high carbon combustion ef®ciencies and sulphur retentions. To simulate and optimize the behaviour of a circulating ¯uidized bed reactor, the mathematical modelling of the hydrodynamic and kinetic characteristics are needed. Modelling CFBC has been attained from different approaches. The models can be grouped under three levels of details of sophistication. The simplest model is onedimensional with plug ¯ow reactor. These are models of Basu et al. [1], Malandrino et al. [2], Weiss et al. [3], Lee and Hyppanen [4], Saraiva et al. [5], Heinbockel and Fett [6] and Remberg et al. [7]. These models consider the char combustion but only three of them take into account the sulphur retention [3,5,6]. It is generally accepted that the CFB riser can be divided into two sections (the denser lower part and the lean upper zone). Some models consider a bubbling [8] or a turbulent regime in the bottom bed of the riser [9] and a dilute region in the upper part of the combustor [10] which can have a core±annulus structure [11±14]. Even there are models that consider an intermediate transition region or splash zone * Corresponding author. Tel.: 134-976733977; fax: 134-976733318. E-mail address: [email protected] (J. Adanez).
between the dense and the dilute region [15±17], as well as an exit zone at the top of the furnace [18,19]. The 1D models do not consider the solid ¯ow in the annular region of the riser, where temperature, gas concentration and velocity can differ from that in the core, in which an up-¯owing dilute region is considered. The models that consider a core±annulus solid ¯ow structure, such as Das and Bhattacharya [20], Sengupta and Basu [13], Basu et al. [12], Haider and Linzer [21], Hannes et al. [22], Talukdar [14], Talukdar and Basu [23], Montat et al. [17] and Park and Basu [24], not all consider the SO2 retention. Thus, only Sengupta and Basu [13], Basu et al. [12], Hannes [16], Talukdar [14] and Montat et al. [17] models included this submodel. Three-dimensional models, as HyppaÈnen et al. [25] and Tsuo et al. [26], use parameters evaluated from experimental results, analyzing the combustion and pollutant generation in the furnace. More rigorous 3D models are being developed [27±30] but provide only predictions of the hydrodynamics of the furnace, the combustion and sulphur retention submodels are yet to be included, as they need a tremendous calculation time even for small periods of time. In this work, a mathematical model for a CFBC combustor operating in the turbulent ¯uidization regime has been developed. The turbulent regime occurs between bubbling and fast ¯uidization [31]. The turbulent regime allows higher residence times of the particles in the riser due to
0016-2361/01/$ - see front matter q 2001 Elsevier Science Ltd. All rights reserved. PII: S 0016-236 1(01)00015-1
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Nomenclature a Ar COp 2 Cs Csup Dg dp Ec Fp F0 F0,coal F1 F2 F3 F6 F7 FA FC FD G Gp Go Gs hf jc k kc K Kc0 Ms P3(r)
decay constant, (m 21) Archimedes number oxygen concentration around the particle (kmol m 23) SO2 concentration at the outlet of the bed (kmol m 23) SO2 concentration in the surface of the particle (kmol m 23) diffusivity of oxygen in nitrogen (m 2 s 21) particle diameter (m) overall carbon combustion ef®ciency (%) carbon ¯owrate to the dense region from F0 1 F1 1 F2 streams (kg s 21) carbon ¯owrate in the feed (kg s 21) coal ¯owrate fed (kg s 21) carbon ¯owrate from the auxiliary bed to the principal bed (kg s 21) descending carbon ¯owrate in the annulus (kg s 21) ascending carbon ¯owrate in the core (kg s 21) carbon ¯owrate non-recovered in the cyclone (kg s 21) carbon drainage from the auxiliary bed (kg s 21) coal ¯owrate in the feed (kg s 21) solids ¯owrate recovered in the bag ®lter (kg s 21) solids ¯owrate drained (kg s 21) mass ¯ux of solids rising in the core (kg m 22 s 21) solid circulation ¯ux above TDH (kg m 22 s 21) mass ¯ux of solids rising from the dense region surface (kg m 22 s 21) external solid circulation ¯ux (kg m 22 s 21) height above dense-region/dilute-region interface (m) carbon fraction in char particles (kg kg 21) solid dispersion constant from core to annulus (m s 21) char combustion kinetic constant (m s 21) constant de®ned in Eq. (6) kinetic parameter at t 0 in Eq. (15) molecular weight of sulphur normalized size distribution function of char in the F3 stream (m 21)
lower gas velocities than in fast ¯uidization, and then better sulphur retentions. To predict the behaviour of a turbulent bed, knowledge of the axial and radial solid concentrations as a function of the main operating parameters (gas velocity, solid circulation ¯ux, etc.) are needed. The mixing pattern of solids and gas in turbulent ¯uidization is different from that of bubbling and fast beds. The main hydrodynamic features of the turbulent regime are that bubbles contain more solids than those in the bubbling regime, the bed surface is much more diffusive than that of a freely bubbling bed because of
Q R r(ri) rc Rep Res ri Dri RS Sh t t(ri) uo up ut Ve Ve(ri) Wcl xcA xcC xcD Xs xs Xs,max
volumetric gas ¯ow rate (m 3 s 21) riser radius (m) shrinking rate of char particles of size ri (m s 21) core radius (m) Reynolds number de®ned as [dpr g(ug 2 us)/m] Reynolds number de®ned as [dpr g(uo 2 ut)/m] mean radii of particles in the population i (m) size interval of the population i (m) sulphur retention Sherwood number time (s) mean residence time of particles of size i (s) super®cial gas velocity (m s 21) solids velocity in annulus (m s 21) single particle terminal velocity (m s 21) volume of solids in the riser (m 3) volume of solids of size i in the riser (m 3) total carbon holdup in the dense region (kg) carbon fraction in the coal feed carbon fraction in the bag ®lter carbon fraction in the drained solids sulfation conversion of limestone sulphur content in the coal maximum sulfation of limestone
Greek letters e cross-sectional averaged voidage ep voidage above TDH ed mean voidage in the dense region ea voidage in the annulus ec voidage in the core e mf minimum ¯uidization voidage e Ro e R at the surface of the dense region rg gas density (kg m 23) rs particle density (kg m 23) e co voidage in the core at the surface of the dense region n volumetric ¯ow (m 3 s 21) Subscripts i relative to the population of char with an average radius
considerable entrainment of particles and the bottom bed remains relatively dense with a porosity around 0.7. Bubbling and fast bed hydrodynamics have been extensively studied in the literature [32]. There have also been considerable studies on the transition to the turbulent ¯uidization regime, but there has been remarkably little investigation on the hydrodynamics of the regime itself [33]. To describe the structure of a turbulent bed, three different approaches have been used in literature. First, Avidan and Edwards [34] considered the bed as a homogeneous one.
J. Adanez et al. / Fuel 80 (2001) 1405±1414
Abed [35], Lee and Kim [36] and Grace [37] suggested a two phases bed, with a dilute phase in the centre and a denser one near the walls. Finally Rowe and MacGillivray [38] and Brereton [39] described the bed with an intermediate structure of slugs and a homogeneous expansion. More recently, Chehbouni et al. [40] used a two phases structure, considering the bed as an intermediate between a bubbling and a circulating regime. They proposed two correlations to predict the average and local void fraction based on the work of Zhang et al. [41]. However, this equation cannot predict the axial voidage in the bed. Ege et al. [42] and Farag et al. [43] studied the hydrodynamic of a turbulent bed and proposed a three-region model, with an annular region constituted by a pseudo-homogeneous phase, and a core region composed by bubbles and an emulsion phase. In a previous work [44] we have investigated some characteristics of the turbulent regime like: the transition velocity (to establish the starting point of the regime), the axial voidage pro®le in the bed and the internal solid circulation ¯uxes using different particle sizes and type of solids in a circulating ¯uidized bed cold pilot plant. The experimental results were analyzed to ®nd a suitable submodel to predict the axial and radial voidage pro®les as a function of the operating conditions. The study revealed that the turbulent bed could be divided into a dense and a dilute region. The dilute region shows a core±annulus structure with a net dispersion of solids from the core to the annulus. The axial voidage pro®le in the dilute region was analyzed with an exponential decay model previously developed for fast bed regime [45]. Previous works [46,47] on modelling of fast bed CFBC were used in this stage to model the carbon combustion ef®ciency and sulphur retention submodels. Kinetics of char combustion was modelled with a shrinking core model with mixed control by chemical reaction and gas ®lm diffusion, assuming that the ash separates once formed. The coal devolatilization was considered as uniform in the dense region with instantaneous combustion of volatiles. The limestone calcination is considered instantaneous. The primary fragmentation of coal, the secondary fragmentation of char and the attrition of limestone have been considered as negligible in the modelling approach carried out in this work. The SO2 generation rate from char depends on the char combustion rate which depends on the riser height due to the existence of axial oxygen concentration pro®le. The SO2 generation rate from volatile is determined through the assumption of uniform devolatilization in the dense region and complete volatile combustion. The SO2 disappearance rate was modelled with a semi-empirical model [48], with a ®rst order kinetic reaction rate for SO2. The sulphation reactivity depends on the SO2 concentration in each height and the residence time for each particle size. The model was validated with experimental results of carbon combustion ef®ciencies and sulphur retentions obtained in a CFBC pilot plant working with a lignite and
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an anthracite. The effect of operating conditions such as coal and limestone particle size distributions, temperature, excess air, air velocity and Ca/S molar ratio on carbon combustion ef®ciency and sulphur retention were analyzed. 2. Experimental Fig. 1 shows a schematic diagram of the experimental set up. The pilot plant is 300 kWth. The apparatus comprised a combustor of 20 cm i.d. and 6.5 m high, a solids (coal and limestone) feeding system, an air line, a cyclone, a line of return of solids to the bed, a heat exchanger, a bag ®lter and a continuous analysis of exhaust gases (CO, CO2, SO2, NOx and O2). The coal and limestone are fed into the bed by a screw feeder. The combustion and ¯uidization air is divided into primary air and secondary air. The primary air is distributed in the bottom of the bed by a distributor plate with bubble caps. The secondary air is introduced through the wall at 1.5 m above the distributor plate. The reactor is refractory lined in the bottom zone and has water walls in the upper 2 m. In the lower zone a heat exchanger is introduced to extract the reaction heat generated. Moreover, the distributor plate has a line for solids draining. At the top of the bed a cyclone is installed for the recovery of the entrained solids which are fed into the riser by an auxiliary line of 9 cm i.d. and a solids valve. The exhaust line has a bag ®lter for the recovery of solid losses in the cyclone. The pilot plant was equipped with analyzers for continuous measurement of exhaust gas mass ¯owrate, gas composition, temperature at different heights in the bed and the total pressure drop in the bed. Two coals with different rank, sulphur and ash contents (Mequinenza lignite: sulphur 7.4%, ash 33.7%; and Bierzo anthracite: sulphur 1.5%, ash 35.7%) were used in the
Fig. 1. Experimental set-up.
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Table 1 Analysis of the coals used Coal
Mequinenza
Bierzo
Proximate analysis (wt%) Moisture Ash Volatile matter Fixed carbon
6.9 33.7 43.2 16.2
0.9 35.7 8.9 54.5
Ultimate analysis (wt%) C H N S
42.9 3.7 0.7 7.9
57.1 1.9 0.9 1.5
experimental work. Table 1 shows the analysis of these coals. Fig. 2 shows the particle size distributions of the coals and the limestone used during experimental work. The experimental work with Mequinenza lignite was made using as sorbent the ArinÄo limestone with two different particle size distributions. Besides, two particle size distributions were used for the Bierzo anthracite to study the effect of this parameter on carbon combustion ef®ciency. The operating variables studied were the air velocity in an interval from 3.2 to 5 m s 21 (assuming that the bed is in the turbulent regime), bed temperature (800, 850 and 9008C), excess air from 10 to 30 and the Ca/S molar ratio (2, 2.5 and 3). During the experimental work, a steady state was maintained for 4 h. At the end of the steady state the different solid streams (bed drained and bag ®lter) were weighed and analyzed for unburned carbon and CaSO4, CaO and CaCO3 content. To avoid analysis errors due to the low C concentration these solids samples were concentrated. Solid samples were leached with HCl which dissolves CaSO4, CaO and CaCO3 increasing the organic C and decreasing the C analysis errors. Carbon analysis was made in a Carlo Erba CHN±O
analyzer. CaSO4 content was analyzed by the Eschka method. The carbon combustion ef®ciency was calculated considering the C feed in and the C losses in the different solid streams (drainage and cyclone) by Eq. (1). The losses due to the CO were not considered because they were negligible: Ec %
FA xcA 2 FD xcD 1 FC xcC £ 100 F A x cA
1
For calculating the sulphur retention, only combustible sulphur was considered, that is organic and pyritic. The sulphur retention was determined taking into account the inlet sulphur, the gas outlet ¯owrate and the SO2 concentration by the following equation: Rs
F0;coal xs =Ms 2 QCs £ 100 F0;coal xs =Ms
2
3. Hydrodynamics of the turbulent bed As it has been pointed out previously, at present, very little data on the hydrodynamic modelling of the turbulent ¯uidization regime are available. Thus, in a previous work [44] the characteristics of the gas±solid ¯ow in this regime of ¯uidization were studied. Measurements of axial voidage pro®les and internal and external solid circulation ¯uxes were carried out in a cold circulating ¯uidized bed (0.1 m i.d. and 4 m high) described elsewhere [45]. The solids used were sand and coal with different particle sizes. The experimental results were analyzed to ®nd a mathematical model suitable for predicting the axial and radial voidage pro®les as a function of the operating conditions. The axial voidage pro®les were measured at different air velocities (inside the range of the turbulent ¯uidization regime for each particle size) obtaining S-shaped pro®les. Therefore, the axial voidage pro®les show a dense zone at the bottom and a dilute zone at the top of the riser, as was found for fast ¯uidization. 3.1. Voidage of the dense region From the literature, two papers were found that calculated the voidage in the dense region of a CFB. Johnsson et al. [49] adapted the equations for a dense bed consisting of an emulsion phase at minimum ¯uidization state penetrated by solid free bubbles. Werther and Wein [9] proposed a correlation for the voidage of the dense region determined in two pilot-scale cold ¯uidized beds working in the turbulent regime: 20:13 ed 2 emf = 1 2 emf 0:14Re0:4 p Ar
Fig. 2. Accumulated weight (%) particle size distributions of coals and limestone used. Mequinenza, Bierzo I and II, ArinÄo limestone I and II.
3
A comparison of the voidages measured at the dense region in the experimental work and the values predicted by these two equations were made. The ®t was not good in both cases. For this reason, it was decided to modify the
J. Adanez et al. / Fuel 80 (2001) 1405±1414
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equation of Werther and Wein [9] since it was developed for turbulent beds. Thus, two new coef®cients were found using the ®tting method of Nelder and Mead [50] to ®nd the best ones. The modi®ed equation is now:
to be valid for both fast and turbulent beds [44]: u Res 0:31 Go 131:1 o rg eco Ar
20:31 ed 2 emf = 1 2 emf 0:19Re0:77 p Ar
The rising solid ¯ux at each height, G, was calculated starting from Go and considering the rate of solids transfer. This rate was calculated with the following equation:
4
3.2. Axial voidage pro®les in the dilute region The Kunii and Levenspiel [51] model modi®ed by Adanez et al. [45] was used to predict the experimental axial voidage pro®les. This model considers an exponential decrease of the bulk density with height. The solid fraction in the freeboard region of the riser becomes: 1 2 e 1 2 ep 1 ep 2 eRo exp 2ahf 2 K
5
The K parameter was determined considering a core± annulus structure in the freeboard. By means of a mass balance the following equation was determined by Adanez et al. [45]: p
K GS 2 G =rs up
6
In the literature there are some equations to calculate the decay constant as a function of the operating conditions for a CFB operating in the fast ¯uidization regime. None of the equations appeared to predict the decay constant of our data, which could be due to the fact that all were determined in the fast regime. Thus, a new equation was proposed by modifying the constants of the equation of Adanez et al. [45] through the ®tting of the experimental `a' values found in our work: a uo 2 ut
1:44
3:45 2 2196 dp
7
3.3. Radial voidage pro®les in the dilute region The internal solid circulation ¯uxes measured in the cold pilot plant showed a parabolic shape with the solids rising mainly through the core and descending near the wall of the riser. These results of a core±annulus structure were similar to those observed in the fast regime. Also, a decrease in both the rising solid ¯ux through the core and the descending solid ¯ux through the annulus could be observed when the height was increased, indicating the existence of a net solid dispersion from the core to the annulus. The model of Rhodes [52] modi®ed by de Diego et al. [53] was used to predict the solid ¯uxes as a function of the operating conditions. This model proposed a core±annulus structure with solid dispersion from the core to the annulus. The model requires the knowledge of the solid ¯ux at the beginning of the dilute region, that is, rising from the dense region surface. The ¯owrate of entrained solids from the dense region was calculated by the following equation, which was found
dG 2r k 1 2 ec 2 1 2 ep 2 s dhf rc
8
9
where the constant `k' is calculated with the equation of de Diego et al. [53] for fast beds: k 0:14= uo 2 ut
10
Taking into account that the solids were distributed between the core and the annulus, the following equation was attained, where the mean voidage at each dilute region height was calculated through Eq. (5):
rc R
2
e 2 ea ec 2 ea
11
The hydrodynamic model was solved in an iterative way and the external solid circulation ¯uxes predicted were compared with the experimental ones. A good correlation was found showing the validity of Eqs. (8)±(11). 4. Coal combustion and sulphur retention As the hydrodynamics characteristics of the turbulent regime have been found to be similar to those of the fast regime, the mathematical models for carbon combustion and sulphur retention developed for fast beds [46,47] have been modi®ed considering the new equations for the turbulent regime. 4.1. Combustion of a coal particle In the modeling of coal combustion, the coal devolatilization is considered to be uniform throughout the dense region with an instantaneous combustion of the volatiles. The shrinking core model with mixed control by chemical reaction and ®lm mass transfer for particles which lose their ashes was assumed to model the char combustion kinetics. The rate at which particles of size ri shrink is given by the following equation: r ri 2
12COp 2 dri dt rc jc 1=kc 1 dp =ShDg
12
The kinetic constants for the different coals were determined in a previous paper [54].
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bed were known. As the contribution of carbon losses in the gas phase is neglected, the theoretical carbon combustion ef®ciency was calculated with the following expression as a function of the carbon ¯owrates in the feed (F0), in the cylone outlet (F6) and in the bed drainage (F7): Ec
F0 2 F6 2 F7 £ 100 F0
14
4.3. Sulphur retention
Fig. 3. The circulating ¯uidized bed system considered.
4.2. Char population balances in the bed The carbon mass balances in a bed with shrinking particles must be solved by taking a population balance of each family of char particles with sizes between ri and ri 1 Dri in the dense region and in the dilute region at different heights. In Fig. 3 a schematic representation of the circulating ¯uidized bed system has been shown, with the different streams considered in the modelling. For a discrete particle size distribution, the population balance of char particles in the dense region has the following equation: Dri Fip Dri 1 Wcl i11 r ri11 Wcl i Dri11 P3 ri Dri Wcl F3 Dri 1 Wcl r ri 1 3Wcl r ri Dri =ri 13 being Fip F0 i 1 F1 i 1 F2 i : These equations are solved iteratively acting over the total carbon weight in the dense region, Wcl, as the iteration parameter and calculating values for Wcl i at each size interval until the convergence condition is reached. Knowing P3(r) in the dense region the ¯owrates of carbon F1 and F2 can be calculated for each char population, solving the population balances in the dilute region and taking into account the ef®ciency of solids recovery by the cyclone. The population balances in the dilute region are solved by dividing the region into compartments in series, where solids are in perfect mixing inside the compartment although in plug ¯ow between them. The char particles passing from the core to the annulus have the same particle size distribution as those present in the respective core compartment. Details about the solving procedure were given elsewhere [46]. When the calculation for this system reached convergence, the size distributions and mass ¯ow of all char streams in the
To know the SO2 concentration in each position of the riser and to make the mass balance it is necessary to know the generation and disappearance of SO2 in each element of volume considered, thus the solution is made iteratively. As the sulphur in coal is distributed between char and volatiles, the generation rate depends on the char combustion rate and the place of devolatilization considered (assuming instantaneous volatile combustion). In the model, the coal devolatilization is assumed to be uniform in the whole dense region, as the devolatilization times (for the typical particle sizes used in the coal combustors) are bigger than the mixing times of the particles. The SO2 generation rate from char depends on the rate of char combustion, which varies with the height in the riser because there is an oxygen pro®le along the riser. As the gas is assumed to have plug ¯ow, the char particle population balances and the SO2 concentration have to be solved simultaneously. The disappearance rate depends on the SO2 concentration in each point. The kinetics of the sulphation reaction of limestone are determined by the sorbent reactivity. A semi-empirical model proposed by Rubiera et al. [48] was used to de®ne limestone reactivity: Kc0 Csup t dXs Kc0 Csup exp 2 Xs;max dt
! 15
The kinetic parameters of this equation Kc0 and Xs,max), which are speci®c of each limestone and particle size, were determined in a batch ¯uidized bed for each limestone and particle size used in the simulation [55]. As can be seen, the sulphation reaction rate depends on the SO2 concentration around the particle and thus the axial SO2 concentration pro®les are needed. It was assumed that the particle size of limestone particles does not change during reaction and the attrition and fragmentation of limestone particles is not considered. Moreover, the calcination of limestone particles was assumed instantaneous. The limestone reactivity also depends on the time of reaction for each particle size and thus on the mean residence time of the particles in the bed. So, it is necessary to know the age distribution of the different particle sizes of limestone present in the riser. The mean residence time is
J. Adanez et al. / Fuel 80 (2001) 1405±1414
de®ned by the following expression: n4 n r 21 t ri 2 1 i Ve ri Ve
100
16
98
In this equation the mean residence time for each particle size depends on its concentration (Ve(ri)) in the bed through the cyclone recovery ef®ciency for each particle size. Thus the sulphation conversion of each limestone particle size is calculated from the age distribution of the sorbent in the bed, assuming perfect mixing of solids in the reactor. To calculate the mean residence times of particles it is necessary to know all the solid ¯uxes, which are dependent on the solids recovery ef®ciency by the cyclone. In this work the ef®ciency of solids recovery is the resultant of the standard curve of a high gas throughput cyclone [56], but modi®ed considering the operating conditions of the combustor.
96
5. Results and discussion An algorithm was developed with the equations shown above to determine the carbon combustion ef®ciencies and sulphur retentions during CFBC combustion in turbulent regime. 5.1. Carbon combustion ef®ciencies Figs. 4±6 show a comparison between experimental carbon combustion ef®ciencies (Ec) and those predicted by the model, showing the effect of the operating conditions on Ec. In all experimental work a Ca/S molar ratio of 2.5 and 10% of secondary air was used, with a pressure drop in the riser of 10,000 N m 22. We have to emphasize the high values of carbon combustion ef®ciencies obtained with lignite, being lower for the anthracite. Effect of bed temperature. Fig. 4 shows the effect of the temperature on the carbon combustion ef®ciency when using Bierzo anthracite at 4 m s 21. As expected, increasing the temperature increases the carbon combustion ef®ciency due to the increase in the reaction rates.
Ec (%)
Bierzo I
94 92 90 88 2.5
3.5
4.5
5.5
u (m/s) Fig. 5. Effect of linear gas velocity on carbon combustion ef®ciency: T 8508C; excess air 20%; ArinÄo I limestone with lignite and anthracite. Mequinenza (B), Bierzo I (X), Bierzo II (W).
Effect of linear gas velocity. Fig. 5 shows the Ec values obtained with the lignite and the anthracite when working at different linear gas velocities maintaining the temperature and the excess air. Also, it shows Ec values obtained with two different particle size distributions of the anthracite. In all cases, as expected, an increase in the linear gas velocity gives a light decrease in the carbon combustion ef®ciencies Ec. A modi®cation of the gas velocity when the excess air and working temperature are constant has two effects. On the one hand, the coal feed increases linearly with gas velocity, which tends to reduce the combustion ef®ciency by increasing the throughput per unit bed area on the system and by the other hand, the solid circulation ¯owrate increases when gas velocity increases and so the ¯owrate of solid losses by the cyclone increases. Moreover, the gas velocity acts in other ways generating ®ne particles different to those formed by shrinkage in the combustion, such as attrition or fragmentation which have not been considered by the model. These effects mainly act on the mean residence time of char particles in the bed. Their effect on the ef®ciency Ec is 100
98
98
96
96 Ec (%)
Ec (%)
Mequinenza Bierzo II
100
94 92
94 92
90
90
88 86 750
1411
88 800
850
900
950
T ( C) Fig. 4. Effect of temperature on carbon combustion ef®ciency with coal Bierzo I and ArinÄo I limestone: uo 4 m s21 ; excess air 20%.
5
10
15
20
25
30
35
Exc (%) Fig. 6. Effect of excess air on carbon combustion ef®ciency with Bierzo I anthracite: T 8508C; uo 4 m s21 :
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J. Adanez et al. / Fuel 80 (2001) 1405±1414
100 Sulphur retention (%)
Sulphur retention (%)
100
90
80
70
60 1.5
2.0
2.5
3.0
3.5
Ca/S Fig. 7. Effect of Ca/S on sulphur retention with Mequinenza lignite and two particle size distributions for the limestone: uo 4 m s21 : ArinÄo I (B), ArinÄo II (A).
small with lignite, since the combustion times of lignites are always much lower than the mean residence times of the char particles in the bed and it was more important when working with the anthracite due to its lower combustion reactivity. The effect of gas velocity was higher using the particle size distribution with a minor proportion of ®ne particles due to the increases in the solid circulation ¯uxes and then in the residence times of the particles. Effect of excess air. As seen in Fig. 6 an increase of excess air gives a small increase in the mean oxygen concentration in the bed, thus increasing the carbon combustion ef®ciency. It can be said that the model applied to the results obtained in the combustion of lignite and anthracites in a pilot plant gives a good prediction of carbon combustion ef®ciencies in the operating conditions of the turbulent regime. 5.2. Sulphur retentions To validate the model for predicting the sulphur retention, a series of experiments was made in the pilot plant described above. The effect of the Ca/S molar ratio, air velocity, particle size distribution of sorbent and the coal type, on the SO2 retention reached in the bed was studied using ArinÄo limestone. Figs. 7 and 8 show a comparison between experimental SO2 retention (Rs) and those predicted by the model, showing the effect of the operating conditions on Rs. In all the experimental work an excess air of 20 and 10% of secondary air injected at 1.5 m was used, with a pressure drop in the riser of 10,000 N m 22. The temperature was set at 8508C. Effect of Ca/S molar ratio. Fig. 7 shows the experimental and predicted sulphur retentions obtained at different Ca/S molar ratios. An air velocity of 4 m s 21 and two different particle size distributions of the limestone were used. As can be seen in the ®gure an increase in the Ca/S molar ratio gives a signi®cant increase in the sulphur retention reached in the bed.
90 80 Mequinenza
70 60
Bierzo I
50
Bierzo II
40 30 3.0 3,0
4.0
5.0
u (m/s) Fig. 8. Effect of air velocity on sulphur retention using ArinÄo I and two types of coals: Ca=S 2:5: Mequinenza (B), Bierzo I (X), Bierzo II (W).
Moreover, the Rs reached with the particle size distribution II (which had a higher proportion of ®ne particles) was higher than those obtained with distribution I. This was due to the effect of the particle size on the limestone reactivity, which increases when the particle size decreases. It has to be pointed out that Ca/S molar ratios higher than 2.5 were necessary with this limestone and particle size distributions to reach Rs higher than 90%. Effect of linear gas velocity. Fig. 8 shows a plot of the experimental and predicted sulphur retentions as a function of the air velocity at Ca=S 2:5 using Mequinenza lignite and Bierzo anthracite. As can be seen in the ®gure, an increase in the air velocity decreases sulphur retention mainly due to two effects. Firstly, it increases the coal throughput increasing the SO2 generation and secondly, it increases the circulation ¯owrates of solids and thus decreases the mean residence time of limestone particles and their conversion in the bed. As shown in the modelling section, the kind of coal acts on the sulphur retention through its sulphur distribution and content. The sulphur content in the char determines the SO2 concentration in the bed. The experimental work was made with two coals with very different characteristics: Mequinenza lignite with a high sulphur content (7.3%) and Bierzo anthracite with lower sulphur content (1.5%). Thus with the anthracite, the low SO2 concentration in the bed acted on the sulphation rates decreasing the limestone conversion as ®rst order kinetics for SO2 is assumed. In addition to the kind of coal, the other characteristic of the coal studied was the particle size distribution fed in. Fig. 8 also shows the effect of the particle size distribution of the coal fed. As can be seen, with bigger mean particle size in the bed, the external solid circulation ¯ux is minor and therefore the mean residence time for limestone particles will be increased (for the same air velocity). If the residence time increases the sulfation conversion increases, and therefore the sulphur retention.
J. Adanez et al. / Fuel 80 (2001) 1405±1414
Fig. 9. Comparison between experimental and predicted by the model carbon combustion ef®ciencies.
Finally, to check the validity of the model, a comparison between the experimental carbon combustion ef®ciencies and those predicted by the model is shown in Fig. 9 including all the experimental points. In general, model predictions for carbon combustion are slighly higher than those obtained in the pilot plant. This could be due to the fact that the model does not consider attrition and fragmentation of coal, which increases the amount of ®nes leaving the riser without being burnt. The same comparison was made between the experimental sulphur retentions and those predicted by the model in Fig. 10. In this case, the predictions for the sulphur retentions are lower for the same reason. The attrition of the limestone generates very ®ne particle sizes where maximum attainable sulfation conversion is greater than the one of the size fed into the riser. However, it can be said that the model gives a good ®t for carbon combustion ef®ciencies and sulphur retentions in the range of operating conditions used. 6. Conclusions A mathematical model for the prediction of carbon combustion ef®ciencies and sulphur retentions in circulating ¯uidized bed combustors previously developed for fast beds has been modi®ed considering the hydrodynamics of the turbulent regime. The model allows prediction of the effect of the main operating conditions, such as air velocity, excess air, temperature, particle size distribution, Ca/S molar ratio, limestone reactivity and coal type. The model was used to predict experimental results obtained in a CFBC pilot plant working in the turbulent ¯uidization regime, burning a lignite and an anthracite. A good ®t was found between the predictions of the model on carbon combustion ef®ciency and sulphur retentions obtained in the pilot plant.
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Fig. 10. Comparison between experimental and predicted by the model sulphur retentions.
References [1] Basu P, Sett A, Gbordzoe EAM. In: Mustonen JP, editor. Ninth International Conference on FBC. New York: ASME, 1997. p. 738. [2] Malandrino A, Arena U, Massimilla L. In: Anthony EJ, editor. 11th International Conference on FBC. New York: ASME, 1991. p. 841. [3] Weiss V, SchoÈler J, Fett FN. In: Basu P, Large JF, editors. Circulating ¯uidized bed technology II. Oxford: Pergamon Press, 1988. p. 289. [4] Lee YY, HyppaÈnen T. International Conference on FBC, 1989;753. [5] Saraiva PC, Azevedo JLT, Carvalho MG. In: Rubow LN, editor. 12th International Conference on FBC, San Diego. New York: ASME, 1993. p. 375. [6] Heinbockel I, Fett FN. In: Basu P, editor. Heat recovery system and CHP, vols. 15, 2. Oxford: Pergamon Press, 1995. p. 171. [7] Remberg CG, Nemet A, Fett FN. In: Preto FDS, editor. 14th International Conference on FBC. New York: ASME, 1997. p. 1139. [8] Leckner B, Golriz MR, Zhang W, Andersson BA, Johnsson F. In: Anthony EJ, editor. 11th International Conference on FBC. New York: ASME, 1991. p. 771. [9] Werther J, Wein J. AIChE Symp Ser 1994;90(301):31. [10] Sotudeh-Garebaagh R, Legros R, Chaouki J, Paris J. Fuel 1998;77(4):327. [11] Huilin L, Guangbo Z, Rushan B, Yongjin C, Gidaspow D. Fuel 2000;79:165. [12] Basu P, Talukdar J, Wu S. AIChE Symp Ser 1994;90(301):114. [13] Sengupta SP, Basu P. In: Anthony EJ, editor. 11th International Conference on FBC, Montreal. New York: ASME, 1991. p. 1295. [14] Talukdar J. PhD thesis, Technical University of Nova Scotia, Canada, 1996. [15] Johnsson F, Leckner B. In: Heinschel KJ, editor. 13th International Conference on FBC. New York: ASME, 1995. p. 671. [16] Hannes J. PhD thesis, Delf University of Technology, The Netherlands, 1996. [17] Montat D, Fauquet P, Lafanechere L, Bursi J. In: Preto FDS, editor. 14th International Conference on FBC. New York: ASME, 1997. p. 1023. [18] Phillippek C, Knobig T, Schonfelder H, Werther J. In: Preto FDS, editor. 14th International Conference on FBC. New York: ASME, 1997. p. 983. [19] Johnsson F, Vrager A, Tiikma T, Leckner B. 15th International Conference on FBC. New York: ASME, 1999. [20] Das A, Bhattacharya SC. Appl Energy 1990;69:227. [21] Haider M, Linzer W. Circulating ¯uidized bed technology VI. In: Avidan A, editor. AIChE. New York, 1994;685. [22] Hannes JP, Renz U, van den Bleek CM. In: Henschel KJ, editor. 13th International Conference on CFB, 1995. p. 287.
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[23] Talukdar J, Basu P. In: Kwauk M, Li J, editors. Circulating ¯uidized bed technology V. Beijing: Science Press, 1997. p. 307. [24] Park C, Basu P. Chem Engng Sci 1998;52(20):3499. [25] HyppaÈnen T, Lee YY, Ranio A. In: Basu P, Horio M, Hasatani M, editors. Circulating ¯uidized bed technology III. Oxford: Pergamon Press, 1991. p. 563. [26] Tsuo YP, Lee YY, Rainio A, HyppaÈnen T. In: Kwauk M, Li J, editors. Circulating ¯uidized bed technology V. Beijing: Science Press, 1997. p. 466. [27] Balzer G, Simonin O. In: Preto FDS, editor. 14th International Conference on FBC. New York: ASME, 1997. p. 1017. [28] Mathiesen V, Solberg T, Hjertager BH. In: Werther J, editor. Circulating ¯uidized bed technology VI. Wurzburg: DECHEMA, 1999. p. 249. [29] Kuipers JAM, van Swaaij WPM. In: Werther J, editor. Circulating ¯uidized bed technology VI. Wurzburg: DECHEMA, 1999. p. 267. [30] Peirano E, Begis J, Johnsson F, Leckner B, Balzer G. In: Werther J, editor. Circulating ¯uidized bed technology VI. Wurzburg: DECHEMA, 1999. p. 281. [31] Yerushalmi J. Gas ¯uidization technology. New York: Wiley, 1986. [32] Lim K, Zhu J, Grace J. Int J Multiphase Flow 1995;21:141. [33] Berruti F, Chaouki J, Godfroy L, Pugsley T, Patience G. Can J Chem Engng 1995;73:579. [34] Avidan A, Edwards M. Fluidization V. New York: Engineering Foundation, 1986. p. 457. [35] Abed R. Fluidization IV. New York: Engineering Foundation, 1984. p. 137. [36] Lee GS, Kim SD. Chem Engng Commun 1989;86:91. [37] Grace JR. Chem Engng Sci 1990;45:1953. [38] Rowe PN, MacGillivray HJ. Fluidization. New York: Plenum Press, 1980. p. 545.
[39] Brereton CMH. PhD thesis, University of British Columbia, Canada, 1987. [40] Chehbouni A, Chaouki J, Guy C, Klvana D. Chem Engng J 1996;61:73. [41] Zhang W, Tung Y, Johnsson F. Chem Engng Sci 1991;46:3045. [42] Ege P, Grislingas A, de Lasa HI. Chem Engng J 1996;61:179. [43] Farag HI, Ege P, Grislingas A, de Lasa HI. Can J Chem Engng 1997;75(10):851. [44] Adanez J, GayaÂn P, de Diego LF, Grasa G. Progress Report No. 4, ECSC Project 7220-ED-081, 1999. [45] Adanez J, Gayan P, Garcia-Labiano F, de Diego LF. Powder Technol 1994;81(3):259. [46] Adanez J, de Diego LF, Gayan P, Armesto L, Cabanillas A. Fuel 1995;74(7):1049. [47] Adanez J, de Diego LF, Gayan P, Armesto L, Cabanillas A. Fuel 1996;75(3):262. [48] Rubiera F, Garcia-Labiano F, Fuertes AB, Pis JJ, Adanez J. 11th International Conference on FBC. New York: ASME, 1991. p. 1489. [49] Johnsson F, Andersson S, Leckner B. Powder Technol 1991;68:117. [50] Nelder JA, Mead RA. Comput J 1964;7:308. [51] Kunii D, Levenspiel O. Powder Technol 1990;61:193. [52] Rhodes MJ. Powder Technol 1990;60:27. [53] de Diego LF, GayaÂn P, Adanez J. Powder Technol 1995;85:19. [54] Adanez J, de Diego LF, AbaÂnades JC. Seventh International Conference on Coal Science, vol. II, 1993. p. 125. [55] Adanez J, Garcia-Labiano F, AbaÂnades JC, de Diego LF. Fuel 1994;73(3):355. [56] Stairmand CJ. Trans Inst Chem Engng 1951;29:356.
www.fuel®rst.com
Circulating ¯uidized bed combustion in the turbulent regime: modelling of carbon combustion ef®ciency and sulphur retention J. Adanez a,*, P. GayaÂn a, G. Grasa a, L.F. de Diego a, L. Armesto b, A. Cabanillas b a
Department of Energy and Environment, Instituto de CarboquõÂmica (CSIC), P.O. Box 589, 50080 Zaragoza, Spain b Ciemat. Avda. Complutense 22, 28040 Madrid, Spain Received 18 September 2000; revised 2 January 2001; accepted 9 January 2001
Abstract A model has been developed considering the hydrodynamic behaviour of a turbulent circulating ¯uidized bed, the kinetic of coal combustion and sulphur retention in the riser. The hydrodynamic characteristics of the turbulent ¯uidization regime were integrated together with the kinetic submodels of char combustion and sulphur retention by limestone. From the combustion of a lignite and an anthracite with limestone addition in a hot CBF pilot plant of 20 cm internal diameter and 6.5 m high, the effect of operating conditions such as temperature, excess air, air velocity, Ca/S molar ratio, coal and limestone particle size distributions on carbon combustion ef®ciency and sulphur retention were studied. The experimental results were compared with those predicted by the model and a good correlation was found for all the conditions used. q 2001 Elsevier Science Ltd. All rights reserved. Keywords: Turbulent ¯uidization; Sulphur retention; Combustion ef®ciency
1. Introduction Circulating ¯uidized bed coal combustion (CFBC) with sorbent addition is a well-known technology for achieving high carbon combustion ef®ciencies and sulphur retentions. To simulate and optimize the behaviour of a circulating ¯uidized bed reactor, the mathematical modelling of the hydrodynamic and kinetic characteristics are needed. Modelling CFBC has been attained from different approaches. The models can be grouped under three levels of details of sophistication. The simplest model is onedimensional with plug ¯ow reactor. These are models of Basu et al. [1], Malandrino et al. [2], Weiss et al. [3], Lee and Hyppanen [4], Saraiva et al. [5], Heinbockel and Fett [6] and Remberg et al. [7]. These models consider the char combustion but only three of them take into account the sulphur retention [3,5,6]. It is generally accepted that the CFB riser can be divided into two sections (the denser lower part and the lean upper zone). Some models consider a bubbling [8] or a turbulent regime in the bottom bed of the riser [9] and a dilute region in the upper part of the combustor [10] which can have a core±annulus structure [11±14]. Even there are models that consider an intermediate transition region or splash zone * Corresponding author. Tel.: 134-976733977; fax: 134-976733318. E-mail address: [email protected] (J. Adanez).
between the dense and the dilute region [15±17], as well as an exit zone at the top of the furnace [18,19]. The 1D models do not consider the solid ¯ow in the annular region of the riser, where temperature, gas concentration and velocity can differ from that in the core, in which an up-¯owing dilute region is considered. The models that consider a core±annulus solid ¯ow structure, such as Das and Bhattacharya [20], Sengupta and Basu [13], Basu et al. [12], Haider and Linzer [21], Hannes et al. [22], Talukdar [14], Talukdar and Basu [23], Montat et al. [17] and Park and Basu [24], not all consider the SO2 retention. Thus, only Sengupta and Basu [13], Basu et al. [12], Hannes [16], Talukdar [14] and Montat et al. [17] models included this submodel. Three-dimensional models, as HyppaÈnen et al. [25] and Tsuo et al. [26], use parameters evaluated from experimental results, analyzing the combustion and pollutant generation in the furnace. More rigorous 3D models are being developed [27±30] but provide only predictions of the hydrodynamics of the furnace, the combustion and sulphur retention submodels are yet to be included, as they need a tremendous calculation time even for small periods of time. In this work, a mathematical model for a CFBC combustor operating in the turbulent ¯uidization regime has been developed. The turbulent regime occurs between bubbling and fast ¯uidization [31]. The turbulent regime allows higher residence times of the particles in the riser due to
0016-2361/01/$ - see front matter q 2001 Elsevier Science Ltd. All rights reserved. PII: S 0016-236 1(01)00015-1
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J. Adanez et al. / Fuel 80 (2001) 1405±1414
Nomenclature a Ar COp 2 Cs Csup Dg dp Ec Fp F0 F0,coal F1 F2 F3 F6 F7 FA FC FD G Gp Go Gs hf jc k kc K Kc0 Ms P3(r)
decay constant, (m 21) Archimedes number oxygen concentration around the particle (kmol m 23) SO2 concentration at the outlet of the bed (kmol m 23) SO2 concentration in the surface of the particle (kmol m 23) diffusivity of oxygen in nitrogen (m 2 s 21) particle diameter (m) overall carbon combustion ef®ciency (%) carbon ¯owrate to the dense region from F0 1 F1 1 F2 streams (kg s 21) carbon ¯owrate in the feed (kg s 21) coal ¯owrate fed (kg s 21) carbon ¯owrate from the auxiliary bed to the principal bed (kg s 21) descending carbon ¯owrate in the annulus (kg s 21) ascending carbon ¯owrate in the core (kg s 21) carbon ¯owrate non-recovered in the cyclone (kg s 21) carbon drainage from the auxiliary bed (kg s 21) coal ¯owrate in the feed (kg s 21) solids ¯owrate recovered in the bag ®lter (kg s 21) solids ¯owrate drained (kg s 21) mass ¯ux of solids rising in the core (kg m 22 s 21) solid circulation ¯ux above TDH (kg m 22 s 21) mass ¯ux of solids rising from the dense region surface (kg m 22 s 21) external solid circulation ¯ux (kg m 22 s 21) height above dense-region/dilute-region interface (m) carbon fraction in char particles (kg kg 21) solid dispersion constant from core to annulus (m s 21) char combustion kinetic constant (m s 21) constant de®ned in Eq. (6) kinetic parameter at t 0 in Eq. (15) molecular weight of sulphur normalized size distribution function of char in the F3 stream (m 21)
lower gas velocities than in fast ¯uidization, and then better sulphur retentions. To predict the behaviour of a turbulent bed, knowledge of the axial and radial solid concentrations as a function of the main operating parameters (gas velocity, solid circulation ¯ux, etc.) are needed. The mixing pattern of solids and gas in turbulent ¯uidization is different from that of bubbling and fast beds. The main hydrodynamic features of the turbulent regime are that bubbles contain more solids than those in the bubbling regime, the bed surface is much more diffusive than that of a freely bubbling bed because of
Q R r(ri) rc Rep Res ri Dri RS Sh t t(ri) uo up ut Ve Ve(ri) Wcl xcA xcC xcD Xs xs Xs,max
volumetric gas ¯ow rate (m 3 s 21) riser radius (m) shrinking rate of char particles of size ri (m s 21) core radius (m) Reynolds number de®ned as [dpr g(ug 2 us)/m] Reynolds number de®ned as [dpr g(uo 2 ut)/m] mean radii of particles in the population i (m) size interval of the population i (m) sulphur retention Sherwood number time (s) mean residence time of particles of size i (s) super®cial gas velocity (m s 21) solids velocity in annulus (m s 21) single particle terminal velocity (m s 21) volume of solids in the riser (m 3) volume of solids of size i in the riser (m 3) total carbon holdup in the dense region (kg) carbon fraction in the coal feed carbon fraction in the bag ®lter carbon fraction in the drained solids sulfation conversion of limestone sulphur content in the coal maximum sulfation of limestone
Greek letters e cross-sectional averaged voidage ep voidage above TDH ed mean voidage in the dense region ea voidage in the annulus ec voidage in the core e mf minimum ¯uidization voidage e Ro e R at the surface of the dense region rg gas density (kg m 23) rs particle density (kg m 23) e co voidage in the core at the surface of the dense region n volumetric ¯ow (m 3 s 21) Subscripts i relative to the population of char with an average radius
considerable entrainment of particles and the bottom bed remains relatively dense with a porosity around 0.7. Bubbling and fast bed hydrodynamics have been extensively studied in the literature [32]. There have also been considerable studies on the transition to the turbulent ¯uidization regime, but there has been remarkably little investigation on the hydrodynamics of the regime itself [33]. To describe the structure of a turbulent bed, three different approaches have been used in literature. First, Avidan and Edwards [34] considered the bed as a homogeneous one.
J. Adanez et al. / Fuel 80 (2001) 1405±1414
Abed [35], Lee and Kim [36] and Grace [37] suggested a two phases bed, with a dilute phase in the centre and a denser one near the walls. Finally Rowe and MacGillivray [38] and Brereton [39] described the bed with an intermediate structure of slugs and a homogeneous expansion. More recently, Chehbouni et al. [40] used a two phases structure, considering the bed as an intermediate between a bubbling and a circulating regime. They proposed two correlations to predict the average and local void fraction based on the work of Zhang et al. [41]. However, this equation cannot predict the axial voidage in the bed. Ege et al. [42] and Farag et al. [43] studied the hydrodynamic of a turbulent bed and proposed a three-region model, with an annular region constituted by a pseudo-homogeneous phase, and a core region composed by bubbles and an emulsion phase. In a previous work [44] we have investigated some characteristics of the turbulent regime like: the transition velocity (to establish the starting point of the regime), the axial voidage pro®le in the bed and the internal solid circulation ¯uxes using different particle sizes and type of solids in a circulating ¯uidized bed cold pilot plant. The experimental results were analyzed to ®nd a suitable submodel to predict the axial and radial voidage pro®les as a function of the operating conditions. The study revealed that the turbulent bed could be divided into a dense and a dilute region. The dilute region shows a core±annulus structure with a net dispersion of solids from the core to the annulus. The axial voidage pro®le in the dilute region was analyzed with an exponential decay model previously developed for fast bed regime [45]. Previous works [46,47] on modelling of fast bed CFBC were used in this stage to model the carbon combustion ef®ciency and sulphur retention submodels. Kinetics of char combustion was modelled with a shrinking core model with mixed control by chemical reaction and gas ®lm diffusion, assuming that the ash separates once formed. The coal devolatilization was considered as uniform in the dense region with instantaneous combustion of volatiles. The limestone calcination is considered instantaneous. The primary fragmentation of coal, the secondary fragmentation of char and the attrition of limestone have been considered as negligible in the modelling approach carried out in this work. The SO2 generation rate from char depends on the char combustion rate which depends on the riser height due to the existence of axial oxygen concentration pro®le. The SO2 generation rate from volatile is determined through the assumption of uniform devolatilization in the dense region and complete volatile combustion. The SO2 disappearance rate was modelled with a semi-empirical model [48], with a ®rst order kinetic reaction rate for SO2. The sulphation reactivity depends on the SO2 concentration in each height and the residence time for each particle size. The model was validated with experimental results of carbon combustion ef®ciencies and sulphur retentions obtained in a CFBC pilot plant working with a lignite and
1407
an anthracite. The effect of operating conditions such as coal and limestone particle size distributions, temperature, excess air, air velocity and Ca/S molar ratio on carbon combustion ef®ciency and sulphur retention were analyzed. 2. Experimental Fig. 1 shows a schematic diagram of the experimental set up. The pilot plant is 300 kWth. The apparatus comprised a combustor of 20 cm i.d. and 6.5 m high, a solids (coal and limestone) feeding system, an air line, a cyclone, a line of return of solids to the bed, a heat exchanger, a bag ®lter and a continuous analysis of exhaust gases (CO, CO2, SO2, NOx and O2). The coal and limestone are fed into the bed by a screw feeder. The combustion and ¯uidization air is divided into primary air and secondary air. The primary air is distributed in the bottom of the bed by a distributor plate with bubble caps. The secondary air is introduced through the wall at 1.5 m above the distributor plate. The reactor is refractory lined in the bottom zone and has water walls in the upper 2 m. In the lower zone a heat exchanger is introduced to extract the reaction heat generated. Moreover, the distributor plate has a line for solids draining. At the top of the bed a cyclone is installed for the recovery of the entrained solids which are fed into the riser by an auxiliary line of 9 cm i.d. and a solids valve. The exhaust line has a bag ®lter for the recovery of solid losses in the cyclone. The pilot plant was equipped with analyzers for continuous measurement of exhaust gas mass ¯owrate, gas composition, temperature at different heights in the bed and the total pressure drop in the bed. Two coals with different rank, sulphur and ash contents (Mequinenza lignite: sulphur 7.4%, ash 33.7%; and Bierzo anthracite: sulphur 1.5%, ash 35.7%) were used in the
Fig. 1. Experimental set-up.
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Table 1 Analysis of the coals used Coal
Mequinenza
Bierzo
Proximate analysis (wt%) Moisture Ash Volatile matter Fixed carbon
6.9 33.7 43.2 16.2
0.9 35.7 8.9 54.5
Ultimate analysis (wt%) C H N S
42.9 3.7 0.7 7.9
57.1 1.9 0.9 1.5
experimental work. Table 1 shows the analysis of these coals. Fig. 2 shows the particle size distributions of the coals and the limestone used during experimental work. The experimental work with Mequinenza lignite was made using as sorbent the ArinÄo limestone with two different particle size distributions. Besides, two particle size distributions were used for the Bierzo anthracite to study the effect of this parameter on carbon combustion ef®ciency. The operating variables studied were the air velocity in an interval from 3.2 to 5 m s 21 (assuming that the bed is in the turbulent regime), bed temperature (800, 850 and 9008C), excess air from 10 to 30 and the Ca/S molar ratio (2, 2.5 and 3). During the experimental work, a steady state was maintained for 4 h. At the end of the steady state the different solid streams (bed drained and bag ®lter) were weighed and analyzed for unburned carbon and CaSO4, CaO and CaCO3 content. To avoid analysis errors due to the low C concentration these solids samples were concentrated. Solid samples were leached with HCl which dissolves CaSO4, CaO and CaCO3 increasing the organic C and decreasing the C analysis errors. Carbon analysis was made in a Carlo Erba CHN±O
analyzer. CaSO4 content was analyzed by the Eschka method. The carbon combustion ef®ciency was calculated considering the C feed in and the C losses in the different solid streams (drainage and cyclone) by Eq. (1). The losses due to the CO were not considered because they were negligible: Ec %
FA xcA 2 FD xcD 1 FC xcC £ 100 F A x cA
1
For calculating the sulphur retention, only combustible sulphur was considered, that is organic and pyritic. The sulphur retention was determined taking into account the inlet sulphur, the gas outlet ¯owrate and the SO2 concentration by the following equation: Rs
F0;coal xs =Ms 2 QCs £ 100 F0;coal xs =Ms
2
3. Hydrodynamics of the turbulent bed As it has been pointed out previously, at present, very little data on the hydrodynamic modelling of the turbulent ¯uidization regime are available. Thus, in a previous work [44] the characteristics of the gas±solid ¯ow in this regime of ¯uidization were studied. Measurements of axial voidage pro®les and internal and external solid circulation ¯uxes were carried out in a cold circulating ¯uidized bed (0.1 m i.d. and 4 m high) described elsewhere [45]. The solids used were sand and coal with different particle sizes. The experimental results were analyzed to ®nd a mathematical model suitable for predicting the axial and radial voidage pro®les as a function of the operating conditions. The axial voidage pro®les were measured at different air velocities (inside the range of the turbulent ¯uidization regime for each particle size) obtaining S-shaped pro®les. Therefore, the axial voidage pro®les show a dense zone at the bottom and a dilute zone at the top of the riser, as was found for fast ¯uidization. 3.1. Voidage of the dense region From the literature, two papers were found that calculated the voidage in the dense region of a CFB. Johnsson et al. [49] adapted the equations for a dense bed consisting of an emulsion phase at minimum ¯uidization state penetrated by solid free bubbles. Werther and Wein [9] proposed a correlation for the voidage of the dense region determined in two pilot-scale cold ¯uidized beds working in the turbulent regime: 20:13 ed 2 emf = 1 2 emf 0:14Re0:4 p Ar
Fig. 2. Accumulated weight (%) particle size distributions of coals and limestone used. Mequinenza, Bierzo I and II, ArinÄo limestone I and II.
3
A comparison of the voidages measured at the dense region in the experimental work and the values predicted by these two equations were made. The ®t was not good in both cases. For this reason, it was decided to modify the
J. Adanez et al. / Fuel 80 (2001) 1405±1414
1409
equation of Werther and Wein [9] since it was developed for turbulent beds. Thus, two new coef®cients were found using the ®tting method of Nelder and Mead [50] to ®nd the best ones. The modi®ed equation is now:
to be valid for both fast and turbulent beds [44]: u Res 0:31 Go 131:1 o rg eco Ar
20:31 ed 2 emf = 1 2 emf 0:19Re0:77 p Ar
The rising solid ¯ux at each height, G, was calculated starting from Go and considering the rate of solids transfer. This rate was calculated with the following equation:
4
3.2. Axial voidage pro®les in the dilute region The Kunii and Levenspiel [51] model modi®ed by Adanez et al. [45] was used to predict the experimental axial voidage pro®les. This model considers an exponential decrease of the bulk density with height. The solid fraction in the freeboard region of the riser becomes: 1 2 e 1 2 ep 1 ep 2 eRo exp 2ahf 2 K
5
The K parameter was determined considering a core± annulus structure in the freeboard. By means of a mass balance the following equation was determined by Adanez et al. [45]: p
K GS 2 G =rs up
6
In the literature there are some equations to calculate the decay constant as a function of the operating conditions for a CFB operating in the fast ¯uidization regime. None of the equations appeared to predict the decay constant of our data, which could be due to the fact that all were determined in the fast regime. Thus, a new equation was proposed by modifying the constants of the equation of Adanez et al. [45] through the ®tting of the experimental `a' values found in our work: a uo 2 ut
1:44
3:45 2 2196 dp
7
3.3. Radial voidage pro®les in the dilute region The internal solid circulation ¯uxes measured in the cold pilot plant showed a parabolic shape with the solids rising mainly through the core and descending near the wall of the riser. These results of a core±annulus structure were similar to those observed in the fast regime. Also, a decrease in both the rising solid ¯ux through the core and the descending solid ¯ux through the annulus could be observed when the height was increased, indicating the existence of a net solid dispersion from the core to the annulus. The model of Rhodes [52] modi®ed by de Diego et al. [53] was used to predict the solid ¯uxes as a function of the operating conditions. This model proposed a core±annulus structure with solid dispersion from the core to the annulus. The model requires the knowledge of the solid ¯ux at the beginning of the dilute region, that is, rising from the dense region surface. The ¯owrate of entrained solids from the dense region was calculated by the following equation, which was found
dG 2r k 1 2 ec 2 1 2 ep 2 s dhf rc
8
9
where the constant `k' is calculated with the equation of de Diego et al. [53] for fast beds: k 0:14= uo 2 ut
10
Taking into account that the solids were distributed between the core and the annulus, the following equation was attained, where the mean voidage at each dilute region height was calculated through Eq. (5):
rc R
2
e 2 ea ec 2 ea
11
The hydrodynamic model was solved in an iterative way and the external solid circulation ¯uxes predicted were compared with the experimental ones. A good correlation was found showing the validity of Eqs. (8)±(11). 4. Coal combustion and sulphur retention As the hydrodynamics characteristics of the turbulent regime have been found to be similar to those of the fast regime, the mathematical models for carbon combustion and sulphur retention developed for fast beds [46,47] have been modi®ed considering the new equations for the turbulent regime. 4.1. Combustion of a coal particle In the modeling of coal combustion, the coal devolatilization is considered to be uniform throughout the dense region with an instantaneous combustion of the volatiles. The shrinking core model with mixed control by chemical reaction and ®lm mass transfer for particles which lose their ashes was assumed to model the char combustion kinetics. The rate at which particles of size ri shrink is given by the following equation: r ri 2
12COp 2 dri dt rc jc 1=kc 1 dp =ShDg
12
The kinetic constants for the different coals were determined in a previous paper [54].
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J. Adanez et al. / Fuel 80 (2001) 1405±1414
bed were known. As the contribution of carbon losses in the gas phase is neglected, the theoretical carbon combustion ef®ciency was calculated with the following expression as a function of the carbon ¯owrates in the feed (F0), in the cylone outlet (F6) and in the bed drainage (F7): Ec
F0 2 F6 2 F7 £ 100 F0
14
4.3. Sulphur retention
Fig. 3. The circulating ¯uidized bed system considered.
4.2. Char population balances in the bed The carbon mass balances in a bed with shrinking particles must be solved by taking a population balance of each family of char particles with sizes between ri and ri 1 Dri in the dense region and in the dilute region at different heights. In Fig. 3 a schematic representation of the circulating ¯uidized bed system has been shown, with the different streams considered in the modelling. For a discrete particle size distribution, the population balance of char particles in the dense region has the following equation: Dri Fip Dri 1 Wcl i11 r ri11 Wcl i Dri11 P3 ri Dri Wcl F3 Dri 1 Wcl r ri 1 3Wcl r ri Dri =ri 13 being Fip F0 i 1 F1 i 1 F2 i : These equations are solved iteratively acting over the total carbon weight in the dense region, Wcl, as the iteration parameter and calculating values for Wcl i at each size interval until the convergence condition is reached. Knowing P3(r) in the dense region the ¯owrates of carbon F1 and F2 can be calculated for each char population, solving the population balances in the dilute region and taking into account the ef®ciency of solids recovery by the cyclone. The population balances in the dilute region are solved by dividing the region into compartments in series, where solids are in perfect mixing inside the compartment although in plug ¯ow between them. The char particles passing from the core to the annulus have the same particle size distribution as those present in the respective core compartment. Details about the solving procedure were given elsewhere [46]. When the calculation for this system reached convergence, the size distributions and mass ¯ow of all char streams in the
To know the SO2 concentration in each position of the riser and to make the mass balance it is necessary to know the generation and disappearance of SO2 in each element of volume considered, thus the solution is made iteratively. As the sulphur in coal is distributed between char and volatiles, the generation rate depends on the char combustion rate and the place of devolatilization considered (assuming instantaneous volatile combustion). In the model, the coal devolatilization is assumed to be uniform in the whole dense region, as the devolatilization times (for the typical particle sizes used in the coal combustors) are bigger than the mixing times of the particles. The SO2 generation rate from char depends on the rate of char combustion, which varies with the height in the riser because there is an oxygen pro®le along the riser. As the gas is assumed to have plug ¯ow, the char particle population balances and the SO2 concentration have to be solved simultaneously. The disappearance rate depends on the SO2 concentration in each point. The kinetics of the sulphation reaction of limestone are determined by the sorbent reactivity. A semi-empirical model proposed by Rubiera et al. [48] was used to de®ne limestone reactivity: Kc0 Csup t dXs Kc0 Csup exp 2 Xs;max dt
! 15
The kinetic parameters of this equation Kc0 and Xs,max), which are speci®c of each limestone and particle size, were determined in a batch ¯uidized bed for each limestone and particle size used in the simulation [55]. As can be seen, the sulphation reaction rate depends on the SO2 concentration around the particle and thus the axial SO2 concentration pro®les are needed. It was assumed that the particle size of limestone particles does not change during reaction and the attrition and fragmentation of limestone particles is not considered. Moreover, the calcination of limestone particles was assumed instantaneous. The limestone reactivity also depends on the time of reaction for each particle size and thus on the mean residence time of the particles in the bed. So, it is necessary to know the age distribution of the different particle sizes of limestone present in the riser. The mean residence time is
J. Adanez et al. / Fuel 80 (2001) 1405±1414
de®ned by the following expression: n4 n r 21 t ri 2 1 i Ve ri Ve
100
16
98
In this equation the mean residence time for each particle size depends on its concentration (Ve(ri)) in the bed through the cyclone recovery ef®ciency for each particle size. Thus the sulphation conversion of each limestone particle size is calculated from the age distribution of the sorbent in the bed, assuming perfect mixing of solids in the reactor. To calculate the mean residence times of particles it is necessary to know all the solid ¯uxes, which are dependent on the solids recovery ef®ciency by the cyclone. In this work the ef®ciency of solids recovery is the resultant of the standard curve of a high gas throughput cyclone [56], but modi®ed considering the operating conditions of the combustor.
96
5. Results and discussion An algorithm was developed with the equations shown above to determine the carbon combustion ef®ciencies and sulphur retentions during CFBC combustion in turbulent regime. 5.1. Carbon combustion ef®ciencies Figs. 4±6 show a comparison between experimental carbon combustion ef®ciencies (Ec) and those predicted by the model, showing the effect of the operating conditions on Ec. In all experimental work a Ca/S molar ratio of 2.5 and 10% of secondary air was used, with a pressure drop in the riser of 10,000 N m 22. We have to emphasize the high values of carbon combustion ef®ciencies obtained with lignite, being lower for the anthracite. Effect of bed temperature. Fig. 4 shows the effect of the temperature on the carbon combustion ef®ciency when using Bierzo anthracite at 4 m s 21. As expected, increasing the temperature increases the carbon combustion ef®ciency due to the increase in the reaction rates.
Ec (%)
Bierzo I
94 92 90 88 2.5
3.5
4.5
5.5
u (m/s) Fig. 5. Effect of linear gas velocity on carbon combustion ef®ciency: T 8508C; excess air 20%; ArinÄo I limestone with lignite and anthracite. Mequinenza (B), Bierzo I (X), Bierzo II (W).
Effect of linear gas velocity. Fig. 5 shows the Ec values obtained with the lignite and the anthracite when working at different linear gas velocities maintaining the temperature and the excess air. Also, it shows Ec values obtained with two different particle size distributions of the anthracite. In all cases, as expected, an increase in the linear gas velocity gives a light decrease in the carbon combustion ef®ciencies Ec. A modi®cation of the gas velocity when the excess air and working temperature are constant has two effects. On the one hand, the coal feed increases linearly with gas velocity, which tends to reduce the combustion ef®ciency by increasing the throughput per unit bed area on the system and by the other hand, the solid circulation ¯owrate increases when gas velocity increases and so the ¯owrate of solid losses by the cyclone increases. Moreover, the gas velocity acts in other ways generating ®ne particles different to those formed by shrinkage in the combustion, such as attrition or fragmentation which have not been considered by the model. These effects mainly act on the mean residence time of char particles in the bed. Their effect on the ef®ciency Ec is 100
98
98
96
96 Ec (%)
Ec (%)
Mequinenza Bierzo II
100
94 92
94 92
90
90
88 86 750
1411
88 800
850
900
950
T ( C) Fig. 4. Effect of temperature on carbon combustion ef®ciency with coal Bierzo I and ArinÄo I limestone: uo 4 m s21 ; excess air 20%.
5
10
15
20
25
30
35
Exc (%) Fig. 6. Effect of excess air on carbon combustion ef®ciency with Bierzo I anthracite: T 8508C; uo 4 m s21 :
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J. Adanez et al. / Fuel 80 (2001) 1405±1414
100 Sulphur retention (%)
Sulphur retention (%)
100
90
80
70
60 1.5
2.0
2.5
3.0
3.5
Ca/S Fig. 7. Effect of Ca/S on sulphur retention with Mequinenza lignite and two particle size distributions for the limestone: uo 4 m s21 : ArinÄo I (B), ArinÄo II (A).
small with lignite, since the combustion times of lignites are always much lower than the mean residence times of the char particles in the bed and it was more important when working with the anthracite due to its lower combustion reactivity. The effect of gas velocity was higher using the particle size distribution with a minor proportion of ®ne particles due to the increases in the solid circulation ¯uxes and then in the residence times of the particles. Effect of excess air. As seen in Fig. 6 an increase of excess air gives a small increase in the mean oxygen concentration in the bed, thus increasing the carbon combustion ef®ciency. It can be said that the model applied to the results obtained in the combustion of lignite and anthracites in a pilot plant gives a good prediction of carbon combustion ef®ciencies in the operating conditions of the turbulent regime. 5.2. Sulphur retentions To validate the model for predicting the sulphur retention, a series of experiments was made in the pilot plant described above. The effect of the Ca/S molar ratio, air velocity, particle size distribution of sorbent and the coal type, on the SO2 retention reached in the bed was studied using ArinÄo limestone. Figs. 7 and 8 show a comparison between experimental SO2 retention (Rs) and those predicted by the model, showing the effect of the operating conditions on Rs. In all the experimental work an excess air of 20 and 10% of secondary air injected at 1.5 m was used, with a pressure drop in the riser of 10,000 N m 22. The temperature was set at 8508C. Effect of Ca/S molar ratio. Fig. 7 shows the experimental and predicted sulphur retentions obtained at different Ca/S molar ratios. An air velocity of 4 m s 21 and two different particle size distributions of the limestone were used. As can be seen in the ®gure an increase in the Ca/S molar ratio gives a signi®cant increase in the sulphur retention reached in the bed.
90 80 Mequinenza
70 60
Bierzo I
50
Bierzo II
40 30 3.0 3,0
4.0
5.0
u (m/s) Fig. 8. Effect of air velocity on sulphur retention using ArinÄo I and two types of coals: Ca=S 2:5: Mequinenza (B), Bierzo I (X), Bierzo II (W).
Moreover, the Rs reached with the particle size distribution II (which had a higher proportion of ®ne particles) was higher than those obtained with distribution I. This was due to the effect of the particle size on the limestone reactivity, which increases when the particle size decreases. It has to be pointed out that Ca/S molar ratios higher than 2.5 were necessary with this limestone and particle size distributions to reach Rs higher than 90%. Effect of linear gas velocity. Fig. 8 shows a plot of the experimental and predicted sulphur retentions as a function of the air velocity at Ca=S 2:5 using Mequinenza lignite and Bierzo anthracite. As can be seen in the ®gure, an increase in the air velocity decreases sulphur retention mainly due to two effects. Firstly, it increases the coal throughput increasing the SO2 generation and secondly, it increases the circulation ¯owrates of solids and thus decreases the mean residence time of limestone particles and their conversion in the bed. As shown in the modelling section, the kind of coal acts on the sulphur retention through its sulphur distribution and content. The sulphur content in the char determines the SO2 concentration in the bed. The experimental work was made with two coals with very different characteristics: Mequinenza lignite with a high sulphur content (7.3%) and Bierzo anthracite with lower sulphur content (1.5%). Thus with the anthracite, the low SO2 concentration in the bed acted on the sulphation rates decreasing the limestone conversion as ®rst order kinetics for SO2 is assumed. In addition to the kind of coal, the other characteristic of the coal studied was the particle size distribution fed in. Fig. 8 also shows the effect of the particle size distribution of the coal fed. As can be seen, with bigger mean particle size in the bed, the external solid circulation ¯ux is minor and therefore the mean residence time for limestone particles will be increased (for the same air velocity). If the residence time increases the sulfation conversion increases, and therefore the sulphur retention.
J. Adanez et al. / Fuel 80 (2001) 1405±1414
Fig. 9. Comparison between experimental and predicted by the model carbon combustion ef®ciencies.
Finally, to check the validity of the model, a comparison between the experimental carbon combustion ef®ciencies and those predicted by the model is shown in Fig. 9 including all the experimental points. In general, model predictions for carbon combustion are slighly higher than those obtained in the pilot plant. This could be due to the fact that the model does not consider attrition and fragmentation of coal, which increases the amount of ®nes leaving the riser without being burnt. The same comparison was made between the experimental sulphur retentions and those predicted by the model in Fig. 10. In this case, the predictions for the sulphur retentions are lower for the same reason. The attrition of the limestone generates very ®ne particle sizes where maximum attainable sulfation conversion is greater than the one of the size fed into the riser. However, it can be said that the model gives a good ®t for carbon combustion ef®ciencies and sulphur retentions in the range of operating conditions used. 6. Conclusions A mathematical model for the prediction of carbon combustion ef®ciencies and sulphur retentions in circulating ¯uidized bed combustors previously developed for fast beds has been modi®ed considering the hydrodynamics of the turbulent regime. The model allows prediction of the effect of the main operating conditions, such as air velocity, excess air, temperature, particle size distribution, Ca/S molar ratio, limestone reactivity and coal type. The model was used to predict experimental results obtained in a CFBC pilot plant working in the turbulent ¯uidization regime, burning a lignite and an anthracite. A good ®t was found between the predictions of the model on carbon combustion ef®ciency and sulphur retentions obtained in the pilot plant.
1413
Fig. 10. Comparison between experimental and predicted by the model sulphur retentions.
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