Control of in-bed agglomeration by fuel blending in a pilot scale straw and wood fueled AFBC

Biomass and Bioenergy Vol. 4, No. 2, pp. 117-133, 1993 0961-9534/93 $6.00 + 0.00 Printed in Great Britain. All rights reserved © 1993 Pergamon Press Ltd

CONTROL OF IN-BED AGGLOMERATION BY FUEL BLENDING IN A PILOT SCALE STRAW

AND WOOD FUELED AFBC

D. SALOUR, *'! B. M. JENKINS,t M. VAFAEI~ 2 and M. KAYHANIAN§ *Electricity Planning Resource Office, California Energy Commission, Sacramento, California, U.S.A.

tAgricultural Engineering Department, University of California, CA 95616, U.S.A. :~ABB Impell Corporation, San Ramon, California, U.S.A.

§Civil Engineering Department, University of California, Davis, U.S.A.

(Received 7 March 1992; accepted 12 August 1992)

Alrstract--Bed agglomeration in a pilot scale atmospheric fluidized bed combustor was controlled by blending wood into rice straw in concentrations of 50% wood or more and stoichiometrically holding reaction temperature at or below 800°C. At higher straw concentrations, agglomeration of the bed occurred over run times inversely proportional to the straw concentration in the blend. Bed agglomeration was preceded by a characteristic decline in bed pressure drop most likely related to combustion air channeling through the bed. Blend ash composition was not substantially affected until straw concentration decreased to 50% or below because of the four-fold higher ash concentration in the straw compared to wood. Blend ash base-to-acid ratio also was not substantially affected above 50% straw concentration, remaining essentially constant at 0.3, compared to the wood ash base-to-acid ratio of 1.1. Initial deformation temperature, as measured by standard cone test of the blend ash, increased from a minimum of 880°C for a 75% straw blend to 1120°C for a 10% straw concentration in the blend.

Keywords--Biomass combustion, rice straw, wood, fluidized bed, ash agglomeration, fouling, slagging.

1. INTRODUCTION

A c o m m o n disposa l technique for cereal s traws, and par t icu la r ly rice s traw, is open field burn ing fol lowing harvest . Such burn ing pro- duces subs tant ia l amoun t s o f po l lu tan t emissions. A recent survey o f Ca l i fo rn ia agr icul ture showed that 64% of the agr icul tura l res idual b iomass burned in the state was rice straw, the rest being p r imar i ly f rom a lmond , walnut , and whea t crops.~ The emissions o f par t icu la te ma t t e r and ca rbon m o n o x i d e f rom sta tewide agr icul tura l burn ing are higher than the emissions o f these po l lu tan ts f rom the s tate 's electri- cal uti l i ty industry. The emission o f biogenic silica fiber f rom the burn ing o f rice s t raw has also created a publ ic percept ion o f subs tant ia l heal th risk, which recently has led to the enact-

rFormerly post-graduate researcher, University of California, Davis.

2Formerly post-graduate researcher, University of California, Davis.

The authors gratefully acknowledge the support of the California Energy Commission, California Rice Research Board, and Thermoelectron Corporation in supporting this work, and that of North American Refractories Company for performing some of the analyses.

Mention of a proprietary trade name does not constitute an endorsement by the University of California.

ment o f legis lat ion (AB 1378) to phase down the a r e a o f rice bu rned in the state by 1998. 2 Cal i forn ia is no t a lone in its concern for burning, and is cer ta inly not the first to restrict agr icul tura l burning. Other areas have also ei ther e l iminated or regula ted agr icul tura l burning. In Asia , where burn ing is also a pr inc ipa l means o f d isposing o f rice straw, subs tant ia l interest exists in f inding al ternat ives to the burn ing p rac t i ce? In Cal i fornia , the recent legis- la t ion forc ing a reduct ion in a rea burned will require growers to seek a l ternat ive d isposa l methods . Ini t ial ly, the me thod employed will mos t l ikely be soil i ncorpora t ion , pe rhaps leading to a change in the pract ice o f con t inuous rice cul ture prevalent in the state since 1912.

M a n a g e m e n t o f rice s t raw by removal f rom the field is ano the r al ternat ive. There may be some advan tages to this technique c o m p a r e d to soil i nco rpo ra t ion in terms o f disease control , a l though it would represent an expor t o f nutr ients (pr incipal ly P and K) if ash were not re turned to the soil. The to ta l quant i ty of s t raw in the state is between 1 and 2 mil l ion tonnes, and no current m a r k e t exists that can handle this much straw. F r o m a power genera t ion perspective, the quant i ty o f s t raw in the state represents a potent ia l o f f rom 100 to 200 MWe,

117

118 D. SALOUR et al.

which could easily be absorbed into the power sector. Because of the pollution associated with field burning, rice straw is an emission offset fuel for new facilities. Several of the independently operated power plants recently constructed in the state were permitted on the basis of using rice straw to offset emissions. Rice straw has proved recalcitrant in its use as a fuel in combustion power facilities. The straw has a low ash deformation temperature which leads to slagging and in fluidized beds to bed agglomeration and defluidization. Other problems also exist, including tube fouling resulting from alkaline deposition. 4

The purpose of this study was to determine if bed agglomeration could be reduced or eliminated when burning rice straw in an atmospheric fluidized bed combustor (AFBC). Previous work had shown that rice straw, wheat straw, and a number of other agricultural biomass fuels could be gasified in a bubbling type fluidized bed when the temperature was held below 760°C with air-fuel ratios between 0.85 and 1.05. 5 Gas higher heating values between 4.85 and 5.87 MJ m -3 were obtained. Subsequent to this work, combustion of rice straw in a pilot scale fluidized bed combustor resulted in bed agglomeration at similar temperatures and above. Since most power plant boilers are designed to operate at temperatures above 760°C (1400°F) for reasons of efficiency, the use of low temperature ash deformation point fuels poses a problem. Because the plants permitted to use rice straw were intended to burn a mixture of fuels, the study described here sought to determine if fuel blending could control bed agglomeration at temperatures reasonably approximating a normal boiler temperature. To do this, mixtures of rice straw and an urban wood waste fuel obtained from a nearby 25 MW~ biomass combustion power plant were prepared and burned in a pilot scale AFBC. Investigations were also made into the effects of blending on fuel bed ash composition.

2. RELATED C O M B U S T I O N RESEARCH

Fluidized bed combustors have been used with a number of biomass fuels on laboratory, pilot and commercial scales. Fuels used include wood, manure, cotton gin trash, olive pits, municipal solid waste (MSW), refuse derived fuel (RDF), rice hulls, dewatered sewage sludge, and a number of others. 6-11

Experience with wood in bubbling bed units has shown a great deal of volatile combustion in the freeboard with temperatures almost 100 K greater than in the bed. 12,13 This reduces the need for in-bed heat exchange surfaces that can suffer from severe corrosion and erosion. It was also found that the combustion efficiency is dependent on the fuel particle size, excess air level, bed temperature, and gas velocity. 14 In bubbling beds with low freeboard heights it may be possible to reduce these dependencies by increasing the height of the freeboard. The alternative for highly volatile fuels of varying particle size is to operate in circulating fluidized bed (CFB) mode. Combustion efficiency of 99.5% has been reported for wood in a CFB at excess air levels above 10%, with limited dependence on bed temperature, particle size, or gas velocity. ~5

The fluidized bed combustion of rice straw has been reported by Rajaram and Malliah) 6 In this experimental study 300 × 300 × 300 mm straw bales were introduced through an over- bed chute into a 0.25 m 2 bubbling bed. The superficial velocities were kept between 1.5 and 3 m s -~ and stable operation was reportedly maintained at 600°C bed temperature. Combus- tion efficiencies on the order of 95-98% were reported without ash recirculation. The fuel energy input to the system was 465 kW. Free- board temperatures were 100K greater than those in the bed. No agglomeration was noticed after 50 h of operation. The success of this operation may be attributed to operating temperatures below the initial deformation temperature (a standard measure of the onset of ash softening as discussed later) of rice straw ash, and also to the high bulk density of the bales. Low temperature operation in the freeboard requires significant heat removal due to the volatility of the straw. The rate of heat removal from the bed or freeboard was not reported. The high bulk density of the bales allowed them to sink in the bed, increasing the residence time and apparently allowing more complete combustion.

Successful operation has also been reported with chopped wheat straw by Washbourne. ~7 The wheat straw used had an ash content of 8.8% with an initial deformation temperature of 980°C. Bed temperatures were maintained at 800°C and a thermal efficiency of 68.5% was achieved when the bed was intermittently fluidized to reduce fuel particle elutriation. No runs longer than 2h were reported and so

Control of in-bed agglomeration 119

the effects of longer term continuous operation are unknown.

Based on pilot plant experience, Energy Products of Idaho Inc. (EPI, now JWP Energy Products, Inc.) has incorporated certain design changes into its bubbling bed boilers to accommodate low density fuels such as straw materials. ~8,19 These modifications include greater bed depths and lower fluidization velocities to increase fuel particle residence times; greater freeboard heights to promote complete vOlatile combustion; and an over-bed pneumatic feed system that injects the fuel at high velocity into the bed. Modifications in the configur- ations and arrangement of the freeboard and convection pass heat exchangers have also been made.

Control of ash induced bed agglomeration has been attempted by adding limestone or kaolin clay to the bed material in order to increase ash fusion temperatures. Since the mechanism behind this phenomenon is not fully understood, additions to the bed have often been done quite liberally, 18 and with occasional adverse results on other plant equipment.

Sintering mechanisms by which particle-to- particle bonding can occur are viscous flow, vapor condensation, diffusion and surface ten- sion. 21 Typically, the ASTM cone test 2° has been used to determine ash fusion temperatures. A study done by Huffman et al. 22 using scanning electron microscopy and X-ray diffraction techniques indicates significant partial melting of some coal ashes at temperatures as low as 200-400 K below the initial deformation temperature (IDT). Their studies indicated that in oxidizing atmospheres, potassium seemed to be the most important fluxing agent below 1100-1200°C. They also found that to retard ash melting, aluminum oxide was the most desirable chemical constituent, whereas iron, calcium, and potassium oxides are the most undesirable. Sintering points as low as 448°C were found by Jung and Schobert 23 for a low rank coal.

There are currently a number of techniques used in the coal industry to predict ash fusion temperatures from composition. The most com- mon is the use of phase diagrams to determine liquid temperatures that parallel the ash fusion temperatures of coal . 24'25 In general the ash composition is normalized to three or four major constituents and the analysis is carried out on the appropriate phase diagram. The accuracy of this technique declines with an

increase in the number of primary constituent oxides in the ash. 26 Biomass fuels vary in ash composition, ranging from rice hull ash of almost pure silica composition, to wood ash with seven or more major species. Empirical models using multiple regression and stepwise regression techniques have also been used as predictors of ash fusion temperature. Where successful, the empirical models have been limited to coal from a certain field, z7 There are also many coal slagging and fouling parameters that are based on ash composition. Many of these have been compiled by Winegartner 28 and some are applicable to biomass ash. For example, the metal oxides in coal ash have been classified as either acids or bases. The total acid content is the sum of the mass fractions of SiO2, TiO2, and A1203 in the ash. The total base in coal ash is the sum of the mass fractions of Fe203 (FeO in reducing atmospheres), CaO, MgO, K20, and Na20. Using these definitions several fouling factors have been developed. An important one is the base-to-acid ratio:

Fe203 -k- CaO + MgO + K20 + Na20 Rb/a = Si02 + TiO2 q- AI203

At values less than 0.75, the melting temperature and viscosity of coal ash at a given temperature has been shown to decrease as Rb/a increases. Above 0.75, the melting temperature usually begins to increase. This trend is also seen with biomass ash as will be shown later, although the point of minimum deformation temperature may be other than that found for coal.

From the base-to-acid ratio various fouling and slagging indices have been developed by Attig and Duzy. 29 The fouling index is defned as Fy = Rb/a (Na20) and the slagging index as Fs = RD/, (S), where (Na20) is the percentage of soda in the ash as prepared by the ASTM method, and (S) is the percentage of sulfur in the dry fuel as found by ultimate analysis. These indices have limited application and can- not be applied to all coal ashes. Their applica- bility to biomass ash has been shown to be unreliable) ° A more fundamental understand- ing of biomass ash properties and behavior needs to be developed because of their import- ance to agglomeration, defluidization, and fouling. The role of sulfates, as well as the silicates, needs to be further investigated for biomass fuels. Sulfur balances on tests conducted in the pilot reactor described here typically show large


discrepancies between total quantities of fuel sulfur and sulfur emitted as SO2, which might be partially accounted for in sulfates.

3. EXPERIMENTAL PROCEDURES

3.1. Pilot scale reactor design

Fluidization characteristics of the reactor used in this study were developed from hydro- dynamic considerations and involved compu- tation of the minimum fluidization velocity, particle terminal velocity, bed height at minimum fluidization, expanded bed height, and the pressure drop across the bed.

The bed material selected for the experiments here was an alumina-silicate sand and prior to use was comprised of nonuniform angular particles. The mean particle diameter was determined through sieve analysis and a mass averaging process:

1 (1) (6/dO,

i = l

where 6i is the weight fraction of bed material captured in the interval between two sieve sizes, and dpi is the average particle diameter in this interval (see Table 1 for nomenclature).

Additionally it was necessary to know the particle sphericity, 42.

F surface area of sphere -]

~s = LSUF----~ac e ~ o-f ~ . J b o t h of . . . . . . lume

This value has been tabulated in the literature, 31 and an appropriate value for sharp sand was taken to be 0.67. The voidage of the bed material at the onset of fluidization can be measured experimentally or it can be estimated from random packing data. The requisite gas properties such as viscosity and density were determined to calculate the minimum fluidization velocity (Umr):

1.75 FdpUmfpg]2 150(1- £mf) [-dpUmrPg] ,/, ,3mr L ~ _ J + - - ~2E3mf L ~ J

= d~ p~(p~ - p~)g (2)

This quadratic equation was solved for the term within the square brackets and sub- sequently for Umf. In calculating Umr the mean particle diameter was used for the size distribution actually present in the bed. In computing the terminal velocity (Ut) the smallest size of solid present in appreciable quantity was used. The terminal velocity for spherical particles was

Table I. Nomenclature

Symbol Unit

A r

Dr g Hf

n~f

~T

D/air Rba Rep Tb L, Up Umf U, Ut. spherical APb~ Ef E m Ernf

P #g 6i

internal cross-sectional area of reactor m2 average diameter of smallest particle of appreciable quantity in the bed material m mean particle diameter of bed material m reactor diameter m gravitational constant m s -2 expanded bed height m static bed height m bed height at minimum fluidization m terminal velocity correction factor for non-spherical particles with Rep < 0.2 terminal velocity correction factor for non-spherical particles with Rep > 1000 terminal velocity correction factor for non-spherical particles with 0.2 < Rep < 1000 mass flow rate of air kg s-I acid-to-base ratio of fuel ash particle Reynolds number bulk gas temperature K standard reference temperature K superficial velocity m s - minimum fluidization velocity m s terminal velocity of a non-spherical particle m s-h terminal velocity of a spherical particle m s- bed pressure drop Pa expanded bed voidage static bed voidage bed voidage at minimum fluidization conditions particle sphericity density kg m 3 viscosity of product gas at temperature N s m--' weight fraction of bed material within a sieve interval


then calculated iteratively by the following equations:

Ut spherical - - g ( p s - - pg)d~ f o r Rep < 0 . 4 " 18pg '

Ot s p h e r i c a l [ 4 (ps-- P_g_g)Zg2-]'/3 d ' = _225 pglAg j p'

for 0.4 ~< Rep < 500

Ut spherical, = I3"lg(Ps-Pg)dp] I/2_pg "

for 500 ~< Rep < 200,000. (5)

A range was chosen for the particle Reynolds number and the corresponding terminal velocity calculated. The original assumption for Rep was then checked against the following equation:

Rep -- dp pg U t (6) /Ag

I f Rep lay outside the original range, a new range was chosen and the process iterated until closure was obtained on Rep. In order to adjust for the irregular shape of the particles, correction factors are applied 32 for particle sphericities between 0.67 and 0.906:

~ s Ksv = 0.843 log 0.065' for Rep < 0.2

KN = 5.31 -- 4.88~s, for Rep > 1000. (8)

In the transitional regime between the two equations above, an interpolation was made:

KTR=IKsT - /0.43](lO00--Rep'~ /0.43

for 0.2 < Rep < 1000. (9)

The terminal velocity for the nonspherical particles in each flow regime was found from:

Ut = Ksv Ut, sphe,cal, for Rep < 0.2 (10)

U t = g T R U t , spherical, for 0.2 < Rep < 1000 (11)

I-0.43-]I/2 U t = / - ~ - - / U,,sphe,ical, for Rep> 1000. (12)

L ~ N J

In a fluidized bed operation with little or no particle recirculation capability, the superficial gas velocity should be maintained between Umf

and Ut to avoid carry-over of solids. The linear or superficial velocity in the vessel at the bulk gas temperature is given by:

(3) Fhai r Tb u 0 - - - • - (13)

Pair, st * A r Tst

where the ratio Tb/Tst describes the expansion of the gas on heating. Ar is the cross-sectional area of the reactor. (4)

To determine the opt imum static bed height a number of cold flow tests were carried out on a 165 m m diameter plexiglass fluidized bed, using lone 420 grain (North American Refrac- tories Company (NARCO), Pleasanton, CA). The distributor used with this bed was similar to that used on the reactor. As part of the test, pressure differential readings were taken across the bed at a frequency of 2000 Hz, over a 1 s time span. These experiments were carried out at flow rates ranging from 2.4 to 14.2Ls -I for 25, 51, 102, 203, and 406mm bed depths. The pressure traces generated by these tests revealed the type of fluidization occurring in the bed. The results showed that the 203 mm and 406 mm beds tended to slug at high velocities. The 102 m m bed exhibited well behaved bubbling fluidization starting from 3 . 3 L s -1 (0.15m s -1) and continuing up to 14.2Ls -i (0.66 m s-1). Although fluidization became pro- gressively more turbulent as the flow rate was

(7) increased, no slugging occurred. The 102 mm static bed depth was taken to be adequate. Pressure data with this bed depth are shown in Fig. 1.

The relationship between bed height and voidage is:

nmr 1 - - £ m - - = ( 1 4 ) am 1 - - £mf

Hr 1 - - E m

nm 1 - - £f" (15)

3000'

2000'

'~ 1000"

2.36 L s "1

({.2 ().4 0'.6 ().8 1.0 Time (s)

Fig. 1. Bed pressure drop for three air flow rates (bed depth = 102 mm).


In eqn. (14) the static bed height (Hm) is controlled, and the voidage at minimum fluidization and static conditions can either be found from random packing data or experimentally. The height at minimum fluidization (Hmf) was calculated using eqn. (14), this being an intermediate step in determining the expanded bed height (Hf) under operating conditions.

At the high superficial velocities achieved when the bed is at operating temperature, the bubble dynamics equations (developed for low velocity fluidization) no longer provide an accurate estimate of bed expansion. More accurate estimates are given by correlations developed for turbulently fluidized beds? 2

U 0 - Umf n f _ 1 -~ (16) Hmr 0.35(gDr) 1/2"

The expanded bed voidage was computed from eqn. (15), and used to estimate the pressure drop across the expanded bed:

APb~d = Hr(1 -- Er)(Ps -- pg)g. (17)

The results of these calculations are shown in Table 2. In these calculations, a reactor diameter of 178 mm was assumed, the reactor temperature was taken to be an average 677°C along the length of the reactor, the bed material had the properties of Investocast 60 alumina-silicate grain (NARCO), the primary air flow rate was 22 L s-t , with a secondary air flow rate of 5 L s t admitted above the bed.

The results indicated that the smallest particles in the grain would become entrained on the gas flow in the lower part of the reactor. To avoid loss of bed particles, the upper portion of the reactor was expanded into a 432mm diameter disengaging zone where the super-

ficial velocity of the gas becomes less than the terminal velocity of the particles.

Estimates of reaction temperature as a function of stoichiometry were also made. For this purpose a simplified combustion model developed by Jenkins and Ebeling 33 was used as a preprocessor in computing the equilibrium adiabatic reaction temperature and product composition via the chemical equilibrium solver STANJAN. 34 While the reaction is not necess- arily expected to come to equilibrium within the residence time prescribed by the reactor length, the trend in temperature with air-fuel ratio was particularly useful in estimating required excess air levels for control by stoichiometry. Com- plete results of this analysis can be found in Salour. 35 The calculations outlined above were used to design the experimental fluidized bed combustor (Fig. 2).

The reactor vessel was of modular design, with the modules bolted together at the flanges. Starting from the bottom, modules included the windbox, distributor, bed section, first stage freeboard, second stage freeboard, transition to the disengagement section, and the disengagement section itself. The designation of the freeboard stages was largely arbitrary, and is used only for future reference of locations on the reactor vessel. This design was employed to facilitate easy repair of each module and future installation of other modules carrying specialized equipment.

The outer shell of the reactor was made of carbon steel. The inside of the reactor was lined with 38 mm of high temperature castable refractory (Pyrominerals Inc., Oakland, CA). The internal diameter of the reactor up to the transition was 178 mm. The vessel expanded to 432 mm in the transition and maintained this inner dimension through the disengagement

Table 2. Fluidization parameters at 22 L s-~ total primary air flow

Parameter Variable Value

Mean particle diameter (Investocast 60) dp 260 #m Diameter of smallest particle dp 178/~m Particle sphericity ~s 0.67 Minimum fluidization velocity Umr 0.04 m s-t Superficial velocity (bed) Uo.b 2.54m s -t Superficial velocity (disengagement zone) U0.d 0.43 m s- Terminal velocity of smallest particle U t 1.43 m s-1 Static bed voidage E m 0.50 Voidage at minimum fluidization velocity Emf 0.52 Expanded bed voidage Ef 0.93 Static bed height Hm 0.10 m Bed height at minimum fluidization nmf 0.11 m Expanded bed height H r 0.68 m Expanded bed pressure drop APbe d 1 377 Pa


section. The overall length of the reactor was 4.6 m. As discussed previously, the disengagement section was intended to reduce the velocity of the combustion gas in the disengagement zone and promote the internal recirculation of fuel and bed particles entrained on the flow. No heat exchange capacity was directly incorporated into the reactor vessel in these tests. Multiple ports were located along the length of the reactor to accommodate instruments and heat exchange tubes if desired. The reactor was otherwise of the internally air cooled type.

The distributor consisted of a 60 mesh stainless screen sandwiched between two perforated stainless steel plates. The holes on the plates were directly aligned and provided 53% open area. The distributor was bolted in place

between the windbox and the bed section. The reactor was fitted with 3 sight glasses, 2 on the lid above the disengagement section and 1 on the bed. The feed system consisted of a metering auger providing fuel to a series of two rotary air locks from which fuel was gravity fed to a high speed auger pushing the fuel into the bed 114 mm above the distributor plate. The rotary air locks were incorporated to prevent backflow of combustion gases from the reactor. Because the air locks were not completely gas tight, overpressure air was introduced into a chamber between the two to induce a positive pressure gradient from bed to environment. Pneumatic feed was also attempted, but discarded in prefer- ence to the mechanical system because of undue bed disturbance.

4570

432

178

o'

O

Disengagement Section

i ~- Sweep Air m

|

I

Secondary Air Port / 1

- Refractory Lining |

|

|

mum

~[ ~ [~ InstmmentPort(typical)

Bed Drains .11 ~1 • [ ~ Fuel injection Port

m ~ ' ~ " Distributor Plate I I Propane Preheat [-----

~ Windbox

Fig. 2. Schematic of pilot scale AFBC (dimensions in mm).


Table 3. Fuel particle size distributions (% mass retained)

Screen 5/16" 1/8" 75% 65% 50% 10% Tyler opening (8 mm) (3 mm) S t r aw S t r a w S t r a w Straw mesh (mm) Straw S t r a w b l e n d b l e n d b l e n d blend

4 4.76 0.10 0.00 0.00 0.00 0.03 0.80 6 3.36 0.25 0.00 0.07 1.50 1.60 3.20 8 2.38 1.37 0.00 3.52 4.28 7.49 I 1.70

20 0.84 31.36 8.71 33.90 37.37 37.41 54.00 40 0.42 33.69 44.62 30.90 27.52 26.11 17.70

100 0.15 28.23 40.99 26.59 23.46 21.68 9.20 pan - - 5.00 5.68 5.02 5.87 5.68 3.40

Primary and secondary combustion air were supplied by a ring compressor capable of 28 kPa at the flow rates desired. Any auxiliary air was supplied from the laboratory compressed air supply. The two air supply systems were completely separate, each with its own manifold and series of air flow meters. The metering system consisted entirely of rotameters (Dwyer Instruments Co., Anaheim, California).

The bed preheating system comprised a propane burner external to the fluidized bed combustor. Combustion gas was used to heat the bed directly. The propane burner included all automatic safety interlocks and overrides. During operation, the preheat system was manually shut off after the bed had achieved a temperature greater than 650°C and solid fuel had been introduced into the reactor. Approximately 1.9 L s ~ of air were continu- ously passed through the preheat system to prevent diffusion of combustible gases from the FBC to the propane burner. After preheating, the solid fuel and combustion air rates were adjusted as necessary to maintain the desired temperature. These rates could be adjusted only within a limited range because of the competing demands of heat transfer and stoichiometry. Although bed drains were installed in the reactor, routine ash removal was by elutriation, and superficial velocities high enough to accomplish this were also required.

Grab samples of reactor exhaust gas were collected and analyzed by gas chromatography for permanent gases and light hydrocarbons. Gas composition was used in completing material and energy balances for the system. Instability of operation when firing rice straw alone sometimes precluded the acquisition of gas samples for these experiments.

3.2. Instrumentation

The data acquisition system used to monitor temperature and pressure fluctuations in the FBC included an electronic CR-7 data logger

(Campbell Scientific, Logan, Utah). The temperature inside the FBC was monitored by 18 inconel-sheathed chromel-alumel thermocouples, each 1.6 mm in diameter and 305 mm long. The thermocouples were adjusted to measure temperatures 51 mm from the inner wall. Two other thermocouples of the same type were used to measure temperature of air coming from the blower and the building compressor. Bed and distributor pressure drop was monitored using a strain gauge type pressure transducer (Microswitch 162 PC).

3.3. Test procedure

Four series of experiments were conducted. The first three series were all conducted using rice straw alone, whereas the fourth series utilized straw and wood fuel blends. All the fuels used in these tests were processed through a hammer mill with a 5/16" (8 mm) round hole screen, with the exception of the fuel used in the third series. The straw used in this series was reduced through a 1/8" (3 mm) screen. Particle size distributions are given in Table 3. Rice straw was collected from a single field of variety M202 in California. The urban wood waste was a mixed demolition waste obtained from a wood-fueled power plant in Woodland, California, and consisted principally of pine and Douglas fir wood. The fuel size and bed grain used in each series are given in Table 4. The properties of the fuels are listed in Table 5.

For the first series of tests using rice straw alone, fuel rates were controlled to between 2 and 3 g s -I ( 8 0 - 1 2 0 g m 2s-~ or 288-432kg

Table 4. Fuel sizes and bed grain used for four test series

Straw Nominal concentration size

Series (%) (mm) Bed grain

1 100 8 lone 420 2 100 8 lone 420 3 100 3 Investocast 60 4 Variable 8 Investocast 60

Control of in-bed agglomeration

Table 5. Physical properties of the fuel blends (all values dry basis except bulk density and moisture)

125

Straw:Wood concentration in blend

Property I00:0 100:0 75:25 6 5 : 3 5 5 0 : 5 0 10:90 0:I00 (8 ram) (3 mm) (8 mm) (8 mm) (8 mm) (8 mm) (8 mm)

Bulk density (kg m -3 w.b.) 117 131 122 140 140 ND ND Moisture (% w.b.) 8.9 8.6 9.4 9.7 10.0 22.7 11.I Heating value (MJ kg -t) 15.68 15.68 16.59 16.95 17.50 18.92 19.32 Proximate analysis:

Volatiles 68.36 68.36 70.41 71.23 72.46 75.73 76.55 Ash 15.60 15.60 12.78 11.66 9.97 5.40 4.33 Fixed carbon 16.04 16.04 16.81 17.11 17.57 18.87 19.12

Ultimate analysis: C 39,31 39.31 41.23 41.99 43.15 46.22 46.98 H 4.91 4.91 5.08 5.15 5.26 5.58 5.60 O 37.05 37.05 39.25 40.14 41.45 45.02 45.85 N 0.45 0.45 0.49 0.50 0.53 0.59 0.60 S 0.08 0.08 0.06 0.06 0.05 0.02 0.01 C1 0.20 0.20 0.17 0.15 0.13 0.08 0.06 Residue 18.00 18.00 13.73 12.02 9.45 2.50 0.90

ND = not determined.

m -2 h -l) and during stable periods excess air levels were about 27% and superficial velocity around 2 m s -l (the stoichiometric air-fuel ratio for the straw used was 4.6). Without the inclusion of internal heat exchange surfaces other than the reactor wall, fuel rates in excess of 3g s -~ were found to result in high and largely uncontrollable reaction temperatures in the freeboard. The second series of tests was conducted using the same fuel material but at fuel rates less than 2 g s -~ and with approximately 100% excess air (1.4 ~ Uo, b ~< 1.9 m s -~ at bed temperature). The third test series was conducted using rice straw at the nominal size of 3 mm. Fuel rates were maintained between 2 and 2.7g s -I, with approximately 75% excess air during periods of limited stability (2.1 ~<Uo, b~<2.5m S--I). In the last two experiments of the third series, the bed was preheated by burning urban wood waste prior to the introduction of the rice straw.

The blends used in the fourth series of experiments had the following ratios of rice straw to urban wood waste: 75:25, 65:35, 50:50, and 10:90. All the blended fuels were of 8 mm nominal size and were mixed on a wet weight basis. The moisture contents were nearly equal at about 10% wet basis. The 10% blend had a somewhat higher moisture content and the actual dry weight straw concentration was 11.6%. Fuel rates were generally maintained between 2 and 3 g s-~ with excess air between 50 and 100% (1.8 ~< Uo, b ~< 2.5 m s-l).

The bed medium used was the same sharp alumina-silica sand described for the cold flow tests. The first and second test series were carried out using the Ione 420 sand grain. The

third and fourth series of tests were conducted using the Investocast 60 grain which has a smaller size distribution than lone 420 (Table 6). This change was made to obtain a more uniform temperature distribution along the length of the reactor. Both grains have the same chemical composition. The sand bed was replaced after each run and was maintained at a static bed depth of 102 mm.

Chemical analyses of the bed materials and fuel ash were conducted by the NARCO Western Research Laboratory using X-Ray Fluorescence (XRF) techniques. Fusion temperatures of the fuel ash and reacted bed materials were determined by standard ASTM cone test 2° in a high temperature Kanthal furnace.

4. COMBUSTION TEST RESULTS

4.1. Unblended straw

All experiments in the first series of tests using rice straw alone ended in bed agglomeration. High temperatures generally extended from the top of the bed section to the middle of the second stage of the freeboard. A drop in the maximum pressure differential across the bed (APbed) below 1 kPa indicated loss of fluidization and the onset of agglomeration within the bed. The decline in bed pressure differential was most likely associated with channeling through the bed as agglomeration proceeded. This occurrence generally preceded the develop- ment of high temperatures (greater than 900°C) in the first stage. As primary air flow rate was increased in an attempt to cool the freeboard, lower-bed temperatures could not be

JBB 4/2--E


Table 6. Chemical and physical properties of bed media

Grain size distribution: Screen opening Tyler mesh (mm)

Percent total weight

Investocast 60 lone 420

8 2.380 10 1.680 14 1.190 20 0.840 28 0.595 35 0.420 48 0.297 65 0.210

100 0.149 200 0.074

Chemical analysis (% mass): SiO2 AI~O 3 Fe203 TiO 2 53 54 43-44 0.4 0.7 1.8-2.4

Physical data: Grain Bulk density (g cm -3) lone 420 1.35 Investocast 60 2.5-2.6

- - trace - - 0-3 - - 25-35 - - 50-60

0-trace 10 15 0-1 1-4

22-42 44-64

9-16 0-1

CaO MgO Na20 K20 P205 0-0.2 0-0.2 0.2 0.4 0.2 0.4 - -

Apparent porosity (%) Specific gravity (g cm 3) 1-4 2.71 2 -4 2.71

Source: NARCO.

maintained above 650°C. Post-test inspections revealed large masses of agglomerates in the bed and a layer of agglomerates on the reactor wall extending from the top of the splash zone up to the transition section.

All three experiments in the second series resulted in extinction in the lower bed after a period of 20 min or so and when the fuel feed was not stopped immediately, the bed filled with fuel because of the lack of reaction in this zone. The fuel build-up resulted in a subsequent ignition from above and pile burning leading to agglomeration. The highest temperatures were achieved in the first stage of the freeboard. High temperatures extended to the second stage of the freeboard as air flow rates were increased. The bed pressure drop was above 1 kPa (indicating good fluidization) until temperatures in the bed dropped below 600°C. The only test in this series that exhibited stability for a reason- able period resulted from a low primary air flow rate at the beginning of the test. At these fuel rates, the heat generated in the bed was not sufficient to maintain temperatures above the ignition temperature, probably as a result of low fuel retention in the bed.

The third series of experiments was conducted with the smaller straw particle size to determine if faster ignition and reduced residence time requirements could help stabilize the bed. The first experiment was conducted in the same manner as before, was difficult to control, and ultimately ended in agglomeration. However, operating times were increased by using the smaller fuel size. In the next two experiments,

the reactor was preheated by burning urban wood waste prior to the introduction of straw. This was done to thermally stabilize the reactor. In a test at a fuel rate of 2.7 g s -l, and with the excess air level kept close to 75%, stable temperatures below 800°C were maintained in the freeboard and temperatures close to 670°C were maintained in the bed. However, 28 min after switching to rice straw there was a sudden temperature rise to 900°C within the lower bed. The pressure drop across the bed had fallen below 1 kPa 6 rain before the temperature surge. Post-test inspection revealed large masses of agglomerates in the bed and a layer of agglomerates along one side of the reactor wall.

4.2. Wood and straw blends

Experiments with 75% rice straw and 25% urban wood waste in the fourth series of tests also ended in agglomeration, but exhibited more stable reactor temperatures for longer periods of time than with rice straw alone. Freeboard temperatures in the first stage reached equilibrium below 800°C, and maintained stability until agglomeration occurred in the bed. The second stage, transition, and disengagement sections never attained thermal equilibrium. In one experiment at 56% excess air, the bed pressure differential fell below 1 kPa 60 min before the end of the test, which lasted 90 min. This behavior was a departure from the other tests where the pressure drop occurred much closer to the time of agglomeration.

Experiments with 65% rice straw and 35% urban wood waste ran for about 2 h prior to


agglomeration. The reactor achieved thermal ~ 5000 equilibrium after the first hour and maintained ~ 4ooo stable temperatures in the bed and first stage up to the point of agglomeration. Temperatures ~ 3000 did show a periodic behavior consistent with ~ 2000 the filling and emptying of the fuel metering auger which varied the air-fuel ratio somewhat. .~ 1000 Bed pressure differential was again observed to ~ 0 ,--

5 decline prior to agglomeration. The end of the = -1000 runs was highlighted by a sudden temperature 0 increase in the bed and first stage, due most likely to the loss of fluidization resulting in a pile burn. Constant adjustment of the air flow rate was necessary. During the post-test inspection large masses of agglomerates were found in the bed. A layer of friable agglomerated material covered the reactor walls and accumulated in stagnation zones behind the thermocouples.

The experiment with 50% rice straw concentration lasted 6 h and was terminated without any indication of agglomeration or loss of fluidization. For the last 3 h of the run the primary air flow rate was constant at 22 L s while the fuel rate was maintained at 2.7 g s -~ (109 g m 2 s-~ or 390 kg m Lh-~). This resulted in an excess air level of 93%. Since high air fuel ratios were maintained from the start of the test it took approximately 1.5 h to achieve thermal equilibrium. The maximum bed pressure drop was above 2 kPa during the first 2.5 h, after which it declined to 1.5 kPa at the end of the experiment. This behavior is similar to that exhibited in the experiments at higher straw concentrations and raises the question as to whether the run would have terminated successfully if continued for a longer time. The only

5000 agglomerates found were in stagnation zones E ! on the trailing edges of the thermocouples, and g~. 4 0 0 0 ~--~,-" these agglomerates were highly friable. The .~ 3o00 -I reactor temperature profile for this test is ~ ' shown in Fig. 3. The temperature control was ~5 2000 ........ i

E much less erratic, but a distinct convexity exists ~ 1000 in the freeboard, indicating greater burning of - ~ volatiles and fine fuel particles in this zone. The ~.~ 0 "v__t__ lower bed temperature is more stable as a result ~:-~00o

o 200 of the longer residence time of the denser wood fuel particles. The operating features of this test are given in Table 7.

200 400 600 800 1000 Temperature (°C)

1200

Fig. 3. Vertical temperature profile along reactor when burning a 50% straw blend.

The experiment with a 10% rice straw concentration maintained excellent stability at an excess air level of 83%. Fuel rates were kept between 2 and 3g s -~ throughout. This test was terminated normally after 3h and no decline in bed pressure drop was observed. No agglomerates were found in the reactor after the run. The temperature profile is shown in Fig. 4. The temperature distribution was uniform in the freeboard up to the transition, after which expansion and heat loss from the disengagement section caused a decline in temperature (the temperature depression at 3 m height is due to secondary sweep air injection occurring at this time). The temperature uni- formity displayed by this test is very typical of the reactor when burning wood fuels, or blends containing large concentrations of wood fuels. The operating features of this experiment are also given in Table 7.

E E E E

X . . . . . . . . .

L E

E

400 600 800 1 000 1200 Temperature (°C)

Fig. 4. Vertical temperature profile along reactor when burning a 10% straw blend.

Table 7. Operating conditions for 50% and 10% rice straw blends

Primary Bed pressure Straw Fuel rate airflow Excess air drop concentration (%) (g s ~) (L s -~) Air fuel ratio (%) (Pa)

50 2.7 19.4 8.7 93 950 10 2.6 19.4 7.5 83 1120


Smaller particles in the Investocast 60 grain were predicted to become entrained on the gas flow in the lower sections of the reactor and drop out in the disengagement section. This was confirmed early on by post test observations which showed that bed material had accumulated on the transition. The 45 ° slope of the transition wall was evidently insufficient. To overcome what resembled a loss of bed material, secondary air injection was employed at the transition to sweep these particles back into the bed.

During many of the experiments using rice straw alone, an attempt was made to reduce temperatures in the freeboard through secondary air injection, but without much success. The ring compressor could not generate sufficient air flow through the flow metering and secondary air nozzles to provide the necessary cooling effect. Instead, the secondary air apparently provided only sufficient oxygen to promote further burning of the volatiles and fine particles in the freeboard.

The expanded bed height was calculated to be 68 cm, which agrees very well with the observed expanded bed height of 59-66 cm. The actual expanded bed height was determined in a post- test inspection by measuring the point where agglomerates first appeared on the reactor wall. This method was only viable in those tests where agglomeration occurred, but is likely representative of bed heights in other tests because of the similar velocities used. The average bed pressure drop was computed from the bed weight to be 1.4 kPa. The actual average pressure drop of about 1.5kPa, observed during periods of stable operation without agglomeration, approximated the predicted value reasonably well. At the superficial velocities utilized, there exist substantial pressure fluctuations (Fig. 1), such that peak pressure differentials exceed 2 kPa. There appears to be a close correlation between the measured fluidization parameters and those calculated from theory.

Power balances on the reactor for the 50% and 10% straw concentrations showed that of the fuel power input (40 and 36 kW respectively), approximately 25% was dissipated through the reactor shell by heat transfer (the other 75% being principally in the sensible power of the exhaust stream). Carbon losses in the exhaust were not determined. Computations of external heat loss from measurements of outer shell temperature confirm losses of 7-10 kW for these experiments.

5. DISCUSSION

The agglomeration phenomena observed here appear to be dependent on both temperature and time. Temperatures in excess of 800°C anywhere in the reactor led to eventual agglomeration in all tests of rice straw alone. Because of the volatile and fine particle burning above the bed, the temperature was most likely to exceed this value in the freeboard. An attempt to control by air cooling with higher primary air flow rates generally led to extinction in the lower part of the bed. Without direct heat extraction from the bed, adequate temperature control by air cooling alone was infeasible. Temperature control in the reactor was there- fore limited by the shell heat transfer and the interaction of increased air flow, higher burning rates, and higher entrainment rates of particles. Direct heat transfer from the freeboard might have enabled better control of agglomeration at higher straw concentrations.

That agglomeration is also a function of time can be observed from the experiments in which straw was blended with wood waste. As the fraction of wood increased in the fuel blend, the time to agglomeration was extended, until at the 50% concentration, no agglomeration was observed to occur within the duration of the experiment. For this experiment, however, the bed pressure differential was seen to decline over time, indicating a potential for later bed agglomeration had the experiment been continued. This decline in bed pressure differential was observed in all experiments terminating in agglomeration, and appears to be a useful predictor of physical changes occurring in the bed.

The adequate cooling and more uniform temperature distribution in the experiments conducted with wood waste appear to be due to the higher particle densities of the wood relative to the straw. At high superficial velocities, the residence time of the wood particles is longer than the straw particles, and implies a more uniform burning profile throughout the reactor. The stability of the reactor on wood is excellent, as it is on the fuel blends with up to 50% weight fraction of straw, when temperature is maintained at or below 800°C.

The delay or prevention of agglomeration by fuel blending appears to be due largely to the changes in ash composition. Agglomeration and bed sintering probably occur when a liquid phase appears in the ash, and acts to bond the

Control of in-bed agglomeration

Table 8. Fuel ash composition (% mass)

129

Fuel SiO2 A1203 Fe203 TiO2 CaO MgO Na20 K20 P205 Rice straw 76.7 1.43 0.46 0.05 2.74 4.28 4.08 8.91 0.42 Wood 37.6 6.61 1.94 0.28 36.8 1.57 1.82 8.69 4.12

solid particles. The time dependent nature of agglomeration may be related to two effects: the first a build-up of ash in the bed with time which could lead to sufficient liquid phase eventually being present to sinter the bed; the second effect being a possible deposition and enrichment of elements from the fuel ash on the bed particles, which also eventually leads to bed sintering. Vapor phase alkali compounds can condense to form alkali-metal aluminosilicates which form a low-melting point "glue" to sinter the particles. 23 The rate at which these two effects proceed would be related both to the concentration of ash in the bed, and the composition of the fuel ash.

In the system representative of straw ash, partial melting can occur at very low temperatures, possibly at temperatures below 600°C, as an inspection of appropriate phase diagrams reveals. 36 Compounds which are responsible for lowering the fusion temperature of silicate systems include Na20, K20, Li20, CaO (at concentrations < 5%) and MgO. Alumina (A1203) and titania (TiO2), however, tend to increase fusion points.

The composition of rice straw ash (Table 8) includes high levels of silicon (Si), potassium (K), moderate amounts of magnesium (Mg) and sodium (Na), and very low concentrations of aluminum (A1). The wood waste fuel ash com-

position is also given in Table 8, and includes lower concentrations of Si, Mg, and Na, about equal concentrations of K, and higher concentrations of A1, Ti, Ca, Fe, and P. The ash concentration of the rice straw was four times higher than the ash concentration of the wood fuel.

The initial deformation temperature of the fuel blend ash increases as the fraction of urban wood waste increases as shown in Table 9. However, the changes in ash composition (Table 10, Figs 5 and 6) and fusion temperatures (Fig. 7) are relatively minor as the concentration of rice straw in the fuel decreases from 100 to 50% (in fact, there is a decline in initial deformation temperature for the 75% and 65% straw concentrations). This is the direct result of the predominance of rice straw ash in the composite ash of the blend. The ash content of the rice straw was 16% as compared to 4% for urban wood waste (residue values from ultimate analysis in Table 5 indicate higher straw ash and lower wood ash concentrations than does the proximate analysis. The proximate analysis is considered more representative because of the much larger sample size involved). To obtain a 50:50 blend of ashes in the fuel, the fuel blend itself would have to be roughly 20% rice straw and 80% urban wood waste. The dominating effect of rice straw shows again in the plot of

Table 9. Ash fusion temperatures and base-to-acid ratios of blended fuels

Ash fusion temperature ( °C) Base-to-acid ratio Straw: wood (%) IDT ST HT FT Rb, ~

100:0 909 997 1077 1274 0.26 75:25 880 990 1095 1225 0.31 65:35 890 1000 1125 1240 0.33 50:50 935 1010 1100 1250 0.38 10:90 1120 1170 1210 1270 0.74 0:100 1175 1200 1210 1230 1.14

IDT = Initial deformation temperature. ST = Softening temperature.

HT = Hemispherical temperature. FT = Fluid temperature.

Table 10. Ash composition of blended fuels (% mass)

Straw : wood (%) SiO 2 A120 ~ Fe203 TiO 2 CaO MgO Na20 K20 P205

75 : 25 73.39 1.87 0.59 0.07 5.62 4.05 3.89 8.89 0.73 65:35 71.62 2.10 0.65 0.08 7.17 3.93 3.79 8.88 0.90 50:50 68.21 2.56 0.78 0.10 10.14 3.69 3.59 8.86 1.22 10:90 50.15 4.95 1.46 0.21 25.87 2.44 2.55 8.76 2.93


-----w--- SiO 2 J- CaO lOO 1

80

g 60

" 40

~ 20 . . . . . . . . .

20 40 60 80 100

Straw Concentration (%)

Fig. 5. Concentrations of major oxides in the blend ash.

A1203 # Fe203 •

• N a 2 0 ~ K20 10

o !

6

4

rO 2'

TiO 2 ~ MgO .t P205

0 20 40 60 80 100

Straw Concentration (%)

Fig. 6. Concentrations of minor oxides in the blend ash.

the fusion temperature for the different fuels against the base-to-acid ratios (Rb/a) in Fig. 8. Those fuels with Rb/a between 0.25 and 0.35 tend to have very similar fusion temperatures. The fuels within this range carry a 50-100% concent ra t ion of rice straw. The initial defor- mat ion temperature (IDT), softening temperature (ST), and hemispherical temperature (HT) increase with an increase in Rb/a above 0.3. The fluid temperature (FT) remains reasonably con-

stant over the full range of Rbi a. The narrow range is remarkable between the initial defor- mat ion temperature and fluid temperature for the wood waste fuel (55 K).

Chemical analyses were also done on the bed materials recovered after each run, except for the 10% straw concentrat ion. Table 11 indicates that as the rice straw concent ra t ion

was decreased, CaO, MgO, Na20, Fe203, and

K20 concentra t ions increased while AI20 3 and TiO2 contents decreased. This can be interpreted as an increase in the fuel ash cont r ibu t ion to the total mass of the bed materials. The increase in ash content is a direct result of the longer operat ing times possible with lower concentrat ions of rice straw in the fuel. When concen- t rat ions are plotted against run time, as in Figs 9 and 10, these trends can be seen more clearly. The concentra t ions of CaO, MgO, K20, and Na20 all increased as the test dura t ion increased. This can also be seen from an approximat ion of the reacted bed composi t ion by comput ing best fit combina t ions of blend ash with original bed grain from Table 6. Such a calculat ion is a rather gross approx imat ion because of the potential for enr ichment to occur in the bed. The results of this calculat ion are given in Table 12. Consider ing the typical

Table 11. Spent bed composition (% mass)

Straw: wood (%) SiO 2 AI203 Fe203 TiO 2 CaO MgO Na20 K20 P205

Investocast 60: 100:0 54.90 40.40 0.58 2.05 0.41 0.46 0.50 1.54 0.15 75:25 54.40 39.90 0.55 2.12 0.40 0.34 0.39 1.40 0.11 65:35 55.40 36.80 0.57 1.96 0.92 0.90 0.76 2.11 0.78 50:50 55.40 34.70 0.89 1.71 2.60 1.50 0.99 2.63 0.49

lone 420: 100:0 54.90 39.50 0.56 2.05 0.32 0.39 0.33 1.20 0.I0

Table 12. Estimated spent bed composition from best fit mixtures of fuel ash and unused grain (% mass)

Estimated bed ash

Straw: w o o d concentration (%) (%) SiO 2 AI203 Fe203 TiO 2 CaO MgO Na20 K20 P205

Investocast 60: 100:0 10.0 54.88 39.74 0.54 1.90 0.45 0.52 0.59 1.25 0.04 75:25 9.5 54.44 40.00 0.55 1.91 0.72 0.48 0.55 1.21 0.07 65:35 16.0 55.52 37.30 0.57 1.78 1.31 0.71 0.77 1.76 0.14 50:50 22.0 56.10 34.91 0.61 1.66 2.56 0.90 0.95 2.30 0.29

lone 420: 100:0 10.0 54.88 39.74 0.54 1.90 0.45 0.52 0.59 1.25 0.04

Cont ro l o f in-bed agg lomera t ion 131

----o-- IDT ~- ST ----o--- HT ~ FT

1 2 0 0 ,~

p 11oo-

10oo ..............................

900- ~

8 ~ • , . 0 20 40 60 80 I00

Rice Straw Concentration (%)

Fig. 7. Ash fusion t empera tu res wi th vary ing concen t ra t ions of s t raw in the fuel b lend ( IDT = ini t ial de fo rma t ion temperature , ST = sof tening tempera ture , H T = hemispher ica l

tempera ture , FT = fluid temperature) .

variations in XRF analysis and the dynamics of the fluidized bed combustion process, the com- parisons are rather good. For the cases shown in Table 12, the best fit ash concentrations are 10% for 100% rice straw and lone 420 grain, 10% for 100% rice straw and Investocast 60 grain, 9.5% for 75% rice straw, 16% for 65% rice straw, and 22% ash for 50% rice straw (P205 consistently shows up surplus by this calculation). The amount of ash remaining in the bed increases as the percentage of rice straw in the fuel blend decreases, which appears to be a direct result of the increased operating time available at lower straw concentrations. The 22% ash concentration for the 50% straw fuel mix corresponds to a retention of 32% of fuel ash in the bed over the course of the experiment.

IDT A ST ~ H T ---, ,---FT 1300-

1200 -

v ~ 1100-

~m. lOOO- te

900-

800 - 0.2 0.4 0.6 0.8 1.0 1.2

Base-to-Acid Ratio

Fig. 8. Ash fusion tempera tures and base- to-acid ra t io of the fuel b lend ( IDT = ini t ial de fo rma t ion tempera ture , ST = sof tening tempera ture , H T = h e m i s p h e r i c a l tempera ture ,

FT = fluid temperature) .

60

50

4 0 " ~ o ,

3O 0

SiO2 --'--'am A1203

---

1 2 3 4 Total run t ime (h)

5 6

Fig. 9. Concen t ra t ions (% mass) of ma jo r oxides in the spent bed mater ia ls .

6. S U M M A R Y A N D C O N C L U S I O N S

Unblended rice straw was not successfully combusted under sustained operation in this reactor, despite attempts to control reaction temperature through stoichiometric means. Agglomeration occurred in all experiments with straw alone where bed temperature was held above extinction. A drop in the differential pressure across the bed was observed prior to the onset of agglomeration, accompanied by a decrease in the lower bed temperature to below 650°C, most likely a result of channeling through the bed by primary fluidizing air. The blending of urban wood waste with rice straw at straw concentrations of 65% and 75% resulted in longer operating times (2 and 1.5 h respectively), but did not avoid agglomeration. Successful operation was achieved only with blends containing straw concentrations of 50% or less. Bed pressure drop did undergo a very gradual decline during the 50% test run, which indicates the possibility of agglomeration with longer operating periods when bed material is not replaced.

Control over agglomeration was obtained by fuel blending for both physical and chemical

Fe203 ~ TiO 2 ~ C a O • MgO ---o--- Na20 A K20 • p205

3 I

~l - ~ I . . . . '

0 • ~ • i . 0 1 2 3 4 5 6

Total run time (h)

Fig. 10. Concen t ra t ions (% mass) o f mino r oxides in the spent bed mater ia ls .

132 D. SALOUR et

reasons . As the pe r cen t age o f u r b a n w o o d

was te inc reased in the fuel b lend , fuel ash was

i n t r o d u c e d in to the r eac to r at a l ower rate.

T h e h igher pa r t i c le dens i ty o f the w o o d also

resu l ted in l o n g e r par t ic le res idence t imes in

the lower bed , wi th m o r e u n i f o r m t e m p e r a t u r e

d i s t r i bu t ions and less fine par t i c le b u r n i n g in

the f r eeboa rd . By inc reas ing the w o o d concen -

t r a t ion , the ini t ial d e f o r m a t i o n t e m p e r a t u r e o f

the fuel b l end ash was a lso increased. T h e b lend

ash c o m p o s i t i o n was no t s ignif icant ly af fec ted

by the a d d i t i o n o f w o o d unt i l c o n c e n t r a t i o n s

a b o v e 5 0 % were ach ieved because o f the

m u c h g rea te r c o n c e n t r a t i o n o f ash in the s t raw.

T h e b lend ash base - to -ac id ra t io r e m a i n e d

near ly c o n s t a n t at a b o u t 0.3 unt i l s t raw concen -

t r a t i on was r educed to 5 0 % or be low. A f t e r tha t

po in t , ba se - to -ac id ra t io inc reased (up to 1.1 for

w o o d a lone) a l o n g wi th an increase in the ini t ial

d e f o r m a t i o n t e m p e r a t u r e o f the fuel. A l t h o u g h

this suggests t ha t fuel b l end ing cou ld be a v iab le

m e c h a n i s m for c o n t r o l l i n g bed a g g l o m e r a t i o n

in the f lu idized bed c o m b u s t o r , the ex ten t to

which tube fou l ing m i g h t be in f luenced was no t

assessed, and requ i res fu r the r analysis .

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Control of in-bed agglomeration by fuel blending in a pilot scale straw and wood fueled AFBC

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