1

THE CONDITIONS FOR BOILER BANK PLUGGING BY SUB-MICRON SODIUM SALT (FUME) PARTICLES IN KRAFT RECOVERY BOILERS W. J. Frederick, Jr. Department of Chemical Engineering and Environmental Science Chalmers University of Technology Kemivägen 10 SE-41296 Gothenburg, Sweden Esa Vakkilainen Andritz-Ahlstrom Corporation Sentnerinkuja 2, P.O. Box 5, FIN-00441 Helsinki, Finland Honghi Tran Pulp & Paper Centre, University of Toronto 200 College Street, Toronto, M5S 3E5, Canada S.J. Lien Institute of Paper Science and Technology 500 10th St., N.W. Atlanta, GA 30318-5794, USA

Presented at the TAPPI-CPPA International Chemical Recovery Conference, Whistler, BC, June, 2001

ABSTRACT

A method for predicting the conditions that lead to plugging of the boiler bank by sub-micron

condensation aerosols of alkali metal salts in kraft recovery boilers has been developed and evaluated

against field observations. The method is based on very recent data on the phase behavior and sintering

characteristics of compacts of dusts from kraft recovery boilers, both of which are very dependent on the

inorganic composition of the particles. The method also considers the thermal load on the boiler, and the

specific design characteristic of gas temperature entering the boiler bank versus thermal load on the boiler.

The prediction method presented in this paper has been tested against data from 20 different recovery

boilers. It predicted correctly whether the boiler encounters problematic fume deposits or not, for 80% of

those boilers.

INTRODUCTION

Kraft recovery boilers, which burn the waste liquor from production of papermaking fiber, have

historically encountered serious problems of fouling and plugging of gas passages in the superheater,

boiler, and economizer banks [1-3]. Except for some economizer deposits, the deposits that form in these

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regions are composed of sodium and potassium sulfate, carbonate, and chloride. The fouling problems are

a direct result of the very high alkali metal content (~20 wt-%) of the spent pulping liquor and the low

melting range of the resulting ash.

In the 1990’s considerable progress was made toward eliminating these fouling and plugging problems.

Recovery boilers built during that period have greater cross-sectional areas in the furnace region; those

built or rebuilt during the period have superior air delivery systems. Problems of gas passage blockage by

deposits formed from entrained residue of the spent liquor droplets have been greatly reduced or

eliminated in these boilers.

In spite of these advances, fouling and blockage of gas passages in the boiler bank of these recovery

boilers is still a persistent problem. A number of studies have been conducted to better understand the

causes of boiler bank fouling and plugging [1-10]. The objective of this paper is to present an analysis of

the physicochemical basis for densification of boiler bank fouling in kraft recovery boilers, and a method

to predict when serious deposition and plugging problems in the boiler bank are likely to be encountered.

The method presented can be used to evaluate the impact of changes in black liquor composition and

thermal load on boiler bank deposition and plugging.

BACKGROUND

Fine, sub-micron size alkali metal salt particles (fume) foul and plug the superheater, boiler, and

economizer banks of kraft recovery boilers, resulting in reduced heat transfer efficiency and lower overall

black liquor burning capacity. Sintering - the process of densification of porous solids - is responsible for

densification and hardening of these deposits. As deposits densify, their strength increases and they

become more difficult to remove by soot-blowing.

Fume particles begin to sinter at about 350oC, and the rate of sintering increases rapidly with temperature

[3-7]. There is wide variation in the sintering behavior and composition of dust samples from different

recovery boilers. The rate at which recovery boiler deposits sinter and harden is controlled by a

combination of dust composition, particle size and packing density, and temperature.

These fine particles apparently sinter and harden into difficult-to-remove boiler bank deposits by two

mechanisms: (a) evaporation and recondensation of sodium chloride from the particles, forming a neck at

the point of contact between two particles, and (b) transport of mass from the particles to the neck by solid-

state diffusion [7]. Figure 1 illustrates the changes that occurred during sintering. As sintering proceeds,

the neck between the fine particles grows until the identity of the initial particles is lost. Beyond that point

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the size of the grains that comprise the structure grows as mass from smaller, individual particles is

transported to the larger ones. These phenomena have been observed both in laboratory studies [3,5,6] and

in deposits formed in operating recovery boilers [9].

Figure 1. Progression of neck development and grain growth during sintering of an ash sample from the electrostatic precipitator of a kraft recovery boiler. Conditions were 450oC for (top left to bottom right) dust from an electrostatic precipitator prior to sintering, and sintered for 3 minutes, 15 minutes, and 60 minutes.

Our recent work has shown that the sintering rate of deposits from fume particles correlates well with both

the first melting temperature and the chloride content of the fine particles [7]. In general, materials begin to

sinter by solid-state diffusion when their homologous temperatures reach a value between 0.5 and 0.8 [11].

The homologous temperature is defined for pure substances as the ratio of the absolute temperature at

which particles are sintered to their melting temperature (T/Tm). For mixtures, the appropriate melting

temperature is the first melting temperature (T0). Recovery boiler dusts begin to sinter appreciably at about

350oC. The range of first melting temperatures is usually between 510 and 600oC. This corresponds to

homologous temperatures between 0.7 and 0.8. Deposits of particles with first melting temperatures below

535-540oC sinter rapidly, and the sintering rate increases as T0 decreases [7].

For dusts that contain at least some chloride, the first melting temperature is determined by the potassium

and carbonate content of the fume particles [10]. For the range of dust composition found in many

operating recovery boilers, increasing the potassium and/or carbonate content of the fume particles

increases their sintering rate.

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Deposits of larger particles, from 5 to 100 µm in diameter, can form simultaneously with fume deposits.

These larger particles form hard deposits by two mechanisms. The first depends upon the lower sticky

temperature of the particles and the gas temperature entering the boiler bank. The lower sticky temperature

is defined as the temperature at which 15 wt-% of the material in the particles is molten [1]. This

mechanism is the same as the deposition mechanism for larger, entrained particles produced by burnout of

individual black liquor droplets [1,2]. The second mechanism is based on the simultaneous deposition of

fume and intermediate-size particles, and the sintering of the finer fume particles to bind the larger

particles into the deposit [9]. Figure 2 shows both spherical intermediate size particles and large fragments

of deposits from an upstream boiler tube as well as fine fume particles. Since sintering rate is inversely

proportional to particle diameter, an assemblage of fume and intermediate-size particles can harden as the

fine fume particles sinter to each other and to the larger, intermediate-size particles. This can occur even at

temperatures below the first melting temperature of the intermediate-size particles.

Figure 2. SEM photomicrograph of a cross-section of an epoxy resin-imbedded deposit formed on a deposition probe inserted in the mid-boiler bank region of a kraft recovery boiler.

PREDICTING BOILER BANK FOULING DUE TO HARDENING OF FUME DEPOSITS

The model that we have developed to predict boiler bank fouling due to hardening of fume deposits is

based on the assumptions that

1. the rate of sintering is fast enough to harden at least part of the material deposited between local

soot-blowing events, to a strength sufficient to withstand the force of the steam jet at the next soot-

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blowing event, and

2. the rate of sintering at any location within the boiler bank is determined by the first melting

temperature and chloride content of the particles that make up the deposit and the gas temperature

at that location.

The composition of the fume particles in a recovery boiler can be measured, or it can be estimated from the

sodium, potassium, and chloride content of the black liquor solids, the expected potassium chloride

enrichment factors, and their expected carbonate content, with the constraints of mass and charged balance

closure. The enrichment factor for potassium or chloride in the fume particles reaching the boiler bank is

bls

i

i

f

i

i

fi

KNaMW

x

KNaMW

x

EF

+

+

=

1.3923

1.3923,

(1)

In Equation 1, EFK,f = 1.5 and EFCl,f = 3.0; xi is the mass fraction of either potassium or chloride in the

fume particles or black liquor solids, MWi is the molecular weight of element i, and Na and K are the mass

fraction of sodium and potassium in the fume particles or black liquor solids.

Sulfur dioxide and oxygen present in the flue gas reacts with NaCl and KCl, producing sulfates and HCl

[12,13]. As a result, the enrichment factor for chloride in the fume particles at the superheater exit depends

upon the sulfur dioxide concentration pointing the upper furnace and superheater regions. When the sulfur

dioxide partial pressure is nearly zero, the enrichment factor for chloride in the fume particles reaching the

boiler bank is as described by Equation 1. However, when significant SO2 is present in the gases exiting

the superheater, the molar ratio of chloride to alkali metals in the fume particles entering the boiler bank is

independent of the chloride to alkali metal mole ratio in black liquor, and

5.11.3923

5.35 =

+f

KNa

Cl (2)

The mass balance equation for the ionic components of fume is

Na + K + Cl + SO4 + CO3 = 1 (3)

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In Equations 2 and 3, Cl, CO3, and SO4 are the mass fractions of chloride, carbonate and sulfate in the

fume particles.

The electroneutrality constraint is

030485.351.392334 =−−−+ COSOClKNa (4)

The mass fraction of carbonate in the fume is estimated from a correlation with the total solids content of

the black liquor burned, and the sulfur to alkali metal mole ratio in the black liquor solids, Equations 5 and

6. This correlation is based on data from 22 industrial recovery boilers.

When Xbls > 65 + 0.5(SRbls-35), then

CO3 = 110.55 - 3.633(Xbls - 0.548SRbls + 18.48)

+ 0.03021(Xbls - 0.548SRbls + 18.48)2 (5)

Otherwise,

CO3 = 0.4207 - 0.00645(65 + 0.5(SRbls-35) - Xbls) (6)

In Equations 5 and 6, SR, the molar ratio of sulfur to alkali metal, is defined as

+=

2.7846

32KNa

SSR

(7)

To calculate the composition of fume particles, the mass fraction of carbonate in the fume particles is first

calculated from Equations 5 or 6 and 7. Equations 1-4, are then solved simultaneously to obtain the mass

fractions of the other four species. Figures 3, 4 and 5 show how well this model predicts the composition

of fume particles, as compared with the measured fume composition from the 22 kraft recovery boilers

from which data had been collected. The mean square differences between the measured and calculated

values were 0.7 wt-% for potassium, 0.9 wt-% for chloride, and 3.9 wt-% for carbonate.

The first melting temperature is estimated from Equations 8, 9, or 10, a correlation developed from the first

melting temperature measurements of Tran et al. [10] and Backman et al. [14]. The correlation is

applicable for fume particles and deposits that contain at least a small amount of chloride and less than 30

wt-% carbonate.

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If KRf < 5%, then

T0 = 633.02 - 17.43 KRf +1.222 KRf2 - 1.357 CO3,f + 0.00975 CO3,f

2 (8)

if 5% < KRf < 15%, then

T0 = 623.04 – 11.00 KRf + 0.3406 KRf2 - 1.401 CO3,f + 0.01188 CO3,f

2 (9)

or, if KRf > 15%, then

T0 = 534.63 - 1.401 CO3,f + 0.01188 CO3,f2 (10)

In Equations 8 and 9, KRf is defined as

+=

1.3923

1.39KNa

KKR

(11)

Figure 6 shows the agreement between first melting temperature values predicted using equations 8-10

versus experimentally measured results for ESP ash samples from 20 different kraft recovery boilers. The

mean square difference between them was 6oC.

Figure 3. Predicted potassium content of fume particles versus the corresponding measured values for 22 kraft recovery boilers.

y = 0.930xR2 = 0.974

0

5

10

15

20

0 5 10 15 20

Measured potassium in fume, wt-%

Calc

ulat

ed p

otas

sium

in fu

me,

wt-

%

8

y = 1.114xR2 = 0.876

0

2

4

6

8

10

0 2 4 6 8 10

Measured chloride in fume, wt-%

Cal

cula

ted

chlo

ride

in fu

me,

wt-

%

Figure 4. Predicted chloride content of fume particles versus the corresponding measured values for 22 kraft recovery boilers.

y = 1.23xR2 = -0.07

0

2

4

6

8

10

12

14

16

18

0 2 4 6 8 10 12 14

Measured carbonate in fume, wt-%

Cal

cula

ted

carb

onat

e in

fum

e, w

t-%

Figure 5. Predicted carbonate content of fume particles versus the corresponding measured values for 20 kraft recovery boilers.

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y = 0.998xR2 = 0.937

500

520

540

560

580

600

620

640

500 520 540 560 580 600 620 640

Measured T0, oC

Cal

cula

ted

T0,

o C

Figure 6. First melting temperature values calculated using equations 8-10, versus experimentally measured results for ESP ash samples from 20 different kraft recovery boilers.

Lien et al. [7] measured the rate of linear shrinkage (∆L/L) of pellets pressed from particles collected in the

electrostatic precipitators of eight different recovery boilers. They correlated the linear shrinkage data with

the chloride content of the dusts, the temperature at which the particles were sintered, and time. Equation

12, the correlating equation, was adapted from Kingery and Berg [15], who used it to correlate sintering

rates for particles sintered by the evaporation-condensation mechanism. Raoult’s law was used to express

Po, the partial pressure of NaCl in this case, in terms of the pure component vapor pressure of NaCl and the

mole fraction NaCl (yNaCl) in the dusts.

=∆

5.0

3/13/1

3/2sin

Tpt

rk

LL ot

(12)

The slopes of linear shrinkage versus t1/3 (Po)1/3 yNaCl/T1/2 are the ratios of ksint/r2/3 for each dust, where ksint

is the rate constant for sintering and r is the surface area-weighted mean radius of the fume particles. These

slopes correlated well with the first melting temperature of the pellet materials [7]. An equation that

describes the relationship between these slopes and the first melting temperatures of the materials is

03/2

sin 115800,61 Trk

f

t −= for T0 < 537oC (13)

and

10

3413/2

sin =f

tr

k

for T0 > 537oC (14)

Equations 13 and 14 describe a strong correlation for ksint with first melting temperature for first melting

temperatures below 537°C, but no relationship between first melting temperature and sintering above that.

The tendency of a deposit of fume particles to plug a boiler bank can therefore be estimated from ksint of

the fume particles or deposit and the gas temperature entering the boiler bank. For a given chloride content

and temperature, deposits whose first melting temperature is less than 537oC are likely to sinter rapidly

because the gas temperatures will almost always be high enough to drive rapid sintering. Deposits sinter

faster, the lower the first melting temperature. Deposits with first melting temperatures higher than 537oC

will sinter far less rapidly.

Equations 15-17 describe the impact of the first melting temperature (T0), the mole fraction of NaCl (yNaCl)

in the deposit, time (t, min), and temperature (T, K) on the linear shrinkage (∆L/L), porosity (P), and

bending strength (σ, MPa) of a fume deposit. Equation 18 was obtained by fitting the bending strength

versus porosity data of Piroozmand et al. [16] for sintered pellets made from recovery boiler ESP catch.

3/12/13/2

1/3sin

0

12.231))+(-21,229/Te( tTr

xpykLL

fume

NaClter=∆

(15)

3

3/12/13/2

1/3sin

03

0

0

12.231))+(-21,229/Te(1

1

1

1)(1

−

−=

∆−

−=−

tTr

xpyk

P

LL

PtP

fume

NaClter

(16)

s= 42.8 e-6.64 P (17)

Figure 7 illustrates how the first melting temperature of fume particles deposited in the boiler bank

influences boiler bank plugging. Typically, gases enter the boiler bank between 550 and 700°C. They exit

at about 400°C. Deposits of fume with a first melting temperature near 510°C will sinter and harden

rapidly, developing a bending strength in excess of 20 MPa throughout the boiler bank. It will be very

difficult to remove them by soot-blowing at 2-hour intervals. At the other extreme, deposits of fume with

a first melting temperature greater than 537°C sinter so slowly that they would easily be removed by soot-

blowing at 2-hour intervals.

2

0.0

0.2

0.4

0.6

0.8

1.0

300 400 500 600Gas temperature at boiler bank entrance, oC

1 -P

oros

ity, -

T0 = 510oC

520oC

530oC

>537oC

30

0.5

10

5.0

1.0

Ben

ding

str

engt

h,M

Pa

0.0

0.2

0.4

0.6

0.8

1.0

300 400 500 600Gas temperature at boiler bank entrance, oC

1 -P

oros

ity, -

T0 = 510oC

520oC

530oC

>537oC

30

0.5

10

5.0

1.0

Ben

ding

str

engt

h,M

Pa

Figure 7. Impact of the first melting temperature of fume particles deposited in the boiler bank on hardening of the resulting deposits. Calculated results, for fume containing 3.0 wt-% chloride, sintered for two hours.

As illustrated by Tran et al. [10], the first melting temperature for particles containing sodium, potassium,

chloride, sulfate, and carbonate decreases with increasing potassium and carbonate content. The values in

Table I are representative of the range of composition of fume particles collected in kraft recovery boilers.

The range for potassium is typically from 4 to 7 wt-%, with extremes of 2.3 wt-% and 11 wt-%. The

range for carbonate is typically 5-15 wt-%, with extremes of 0 wt-% to nearly 50 wt-%. Within this range

of composition, the range of first melting points is from 513°C to nearly 600°C as long as the chloride

content is nonzero. When both with the carbonate and chloride content are zero, the first melting

temperature can be as high as 810°C. The boundary for potassium and chloride composition to stay above

a first melting temperature of 537°C is

CO3 + 17.14 ln (K) < 32.3 (18)

Equation 18 is applicable as long as the chloride content is nonzero. When the chloride content is zero,

the first melting temperature is significantly higher.

Figure 7 indicates that the hardening rate of fume deposits increases very rapidly as the first melting

temperature of the fume particles drops below 537°C. This means that boiler bank plugging may be rather

sensitive to black liquor composition and recovery boiler firing conditions, both of which can influence

first melting temperature.

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Table I. Range of composition for ESP catch samples from 47 kraft recovery boilers.

Low Typical High Potassium 0.3% 3-9% 17% Chloride 0.0% 2-7% 14% Carbonate 0.0% 2-8% 24% Range of T0’s, oC 810 530-560 513

PREDICTING BOILER BANK PLUGGING DUE TO STICKY INTERMEDIATE PARTICLES

Deposits of intermediate particles on boiler bank tubes develop strength primarily by impinging on boiler

bank tubes while partially molten, forming strong deposits as they solidify. The mechanism is the same as

that for the formation of deposits from partially molten carryover particles [1,2]. These intermediate size

particles are partially depleted of potassium and chloride, mainly from evaporation of NaCl and KCl.

Analysis of data reported for black liquor char burned in a laboratory char bed combustor [17,18] shows

that the enrichment factors for intermediate particles formed during char bed burning are 0.49±0.30 for

chloride and 0.68±0.06 for potassium. The composition of these intermediate particles can be estimated

from the sodium, potassium, and chloride content of the black liquor solids, using these enrichment factor

values in Equations 1 and 2.

A third constraint is that the molar ratio sulfate to sodium in these particles is approximately the same as

the molar ratio of sulfur to sodium in black liquor, i.e.

(S/Na)ip = (S/Na)bls (19)

The last two constraints are the mass and charge balance constraints, Equations 3 and 4.

The composition of intermediate particles is calculated from the sodium, sulfur, potassium, and chloride

content of the black liquor solids, solving simultaneously Equations 1 and 2 with the chloride and

potassium enrichment factors set equal to 0.49 and 0.68, respectively, and Equations 3, 4, and 19.

4

0

2

4

6

8

10

12

14

0 2 4 6 8 10K in BL solids, wt-%

K o

rCli

n in

term

edia

te p

artic

les,

wt-%

Cl

K

T15

Low

er s

ticky

tem

pera

ture

(T15

), o C

780

760

740

720

700

800

820

840

780

760

740

720

700

800

820

840

Figure 8. Impact of potassium content of black liquor on the potassium and chloride content of the intermediate particles produced, and on their lower sticky temperature (T15). Calculations are for black liquor solids containing 1% chloride, 20% Na + K as Na equivalent, and a sulfur to alkali metal mole ratio of 30%.

The temperature of the flue gas at the boiler bank entrance must been known to determine whether the

intermediate particles are sticky at that point. Väätäinen [19] developed a correlation to estimate the gas

temperature at the boiler bank entrance as a function of thermal load (LT, % of design load), oxygen

concentration in the flue gas (O2, mole-%, wet basis), and the dry solids content of the black liquor (ds,

wt-%). His correlation, Equation 20, includes a term Tbase which is a function of the heat transfer area in

the superheater per unit of black liquor solids fired. It can be obtained from a few measurements of the

gas temperature at the boiler bank entrance, or estimated from boiler design information. Values of Tbase

typically fall between 575 and 625oC.

TBB = Tbase -1.45(100- LT) -5.62(3- O2) + 0.89(70- ds) (20)

The lower sticky temperature can be estimated from the work of Tran et al. [10] and Backman et al. [14].

Equations 21 and 22 are a correlation developed between these lower sticky temperature values and the

potassium and chloride content of the particles. This correlation was used to estimate the lower sticky

temperature in the present work.

5

0

2

4

6

8

10

12

14

0 2 4 6 8 10Cl in BL solids, wt-%

K o

rCli

n in

term

edia

te p

artic

les,

wt-%

Cl

K

T15

Low

er s

ticky

tem

pera

ture

(T15

), o C

760

720

680

640

600

800

880

860

0

2

4

6

8

10

12

14

0 2 4 6 8 10Cl in BL solids, wt-%

K o

rCli

n in

term

edia

te p

artic

les,

wt-%

Cl

K

T15

Low

er s

ticky

tem

pera

ture

(T15

), o C

760

720

680

640

600

800

880

860

760

720

680

640

600

800

880

860

Figure 9. Impact of chloride content of black liquor on the potassium and chloride content of the intermediate particles produced, and on their lower sticky temperature (T15,ip). Calculations are for black liquor solids containing 3% potassium, 20% Na + K as Na equivalent, and a sulfur to alkali metal mole ratio of 30%.

When

PP KNaK

KNaCl

++<

+002.06.5

then

++

+−

+−−+

++

+−=

22

15 105.0168.6174.10431.011(9043.21898.60171.871PPPPP KNa

ClKNa

KKNa

ClKNa

ClKNa

ClT

(21)

Otherwise,

2

15 105.0168.683.624PP KNa

KKNa

KT

++

+−=

(22)

1

Figures 8 and 9 show how the intermediate particle composition and lower sticky temperature is predicted

to change with the chloride and potassium content of black liquor, based on the model described here. It

decreases more rapidly with increasing chloride content than with increasing potassium content.

COMPARISON OF BOILER BANK PLUGGING PREDICTIONS WITH FIELD

OBSERVATIONS

In our analysis, the term “problematic fume deposits” refers to deposits in boilers that require excessive

soot-blowing steam to operate without frequent boiler bank washing, or that require frequent water

washing as a result of fume deposits. Four variables were selected as possible predictors of problematic

fume deposits in the boiler bank of recovery boilers. These variables were (a) the first melting

temperature (T0), (b) the chloride content of the fume particles (xCl,f), (c) the difference between the gas

temperature at the boiler bank entrance and the lower sticky temperature (∆T15,f = Tg,BB – T15,f), and (d) a

predicted bending strength (σ2hr) greater than 5 MPa for deposits sintered for two hours at the gas

temperature at the boiler bank entrance. T0 was selected because fume compacts with lower first melting

temperatures were observed to sinter more rapidly [7]. xCl,f was selected because these same fume

compacts were observed to sinter more rapidly when they contained more chloride; a value of xCl,f > 2%

was chosen as the criterion for problematic boiler bank deposits. ∆T15,f was included because fume

deposits can become sufficiently molten to densify by viscous transport [20] when Tg,BB exceeds T15,f. σ2hr

was chosen because it combines the impact of T0 and xCl,f on sintering rate with the conditions of ambient

temperature five time for sintering at the lower bank entrance; a minimum value of σ2hr > 5 was used to

indicate problematic boiler bank fume deposits.

Table II shows the furnace cross-sectional area, and the design capacity and operating load for 20 of the

22 kraft recovery boilers for which data on the operation and performance has been collected. These

boilers cover a 20-fold capacity range from smallest to largest.

Table III contains the concentrated black liquor solids content, and the average potassium, chloride, and

carbonate content of the black liquor and the fine particles collected in the electrostatic precipitator.

Table IV contains the measured values for chloride composition and first melting temperatures of

precipitator ash samples from 20 of the recovery boilers. The first melting temperatures were calculated

from the measured ash composition, using Equations 9-11. Table IV also includes calculated values of

the gas temperature entering the boiler bank, and the estimated bending strength of deposits formed in the

boiler bank and exposed for two hours to the temperature of the gas entering the boiler bank entrance

(Tg,BB). Four of these boilers (boilers C, F, N, and W) have been observed to have boiler bank fouling

2

problems. Also included in Table IV is each mill's experience with boiler bank fouling.

Two of these boilers (H and J) are normally fired 10-20% below their design load. These boilers would

be expected to encounter severe problems at the firing rate that they were evaluated (see Table II).

However, they would not be expected to encounter fume deposition problems during their normal, low

load operation. Boilers H and J are not considered in the analysis that follows.

Table II. The furnace cross-sectional area, design capacity and operating load for 20 kraft recovery boilers.

Design capacity Boiler

designation

Furnace cross-section

area, m2 Tons dry solids/day kW/m2

Operating thermal load, % of design

A 118 1700 2428 130 B 48 831 2715 134 C 104 1500 2373 150 D 159 3100 3047 113 E 105 1600 2608 114 F 105 2000 3376 100 G 117 1837 2644 100 H 80 1211 2486 111 I 60 904 2416 127 J 88 1362 2514 123 K 72 1220 2849 107 L 60 950 2552 100 M 95 1600 2927 100 N 39 670 2733 100 O 105 1900 2959 100 P 102 1470 2344 110 Q 10 156 2634 100 W 167 3150 3125 122 X Y 131 2041 2520 125 Z 194 3150 2845 100 Å

3

Table III. Black liquor and ESP dust composition and first melting temperatures of the ESP dusts for the recovery boilers in Table II.

Black liquor composition

Average wt-% S/(Na2+K2)ESP dust composition

Average wt-% ESP dust

first melting

Boiler

Black liquor dry solids

content, wt-% Na Cl K mole-% Cl K CO3 temp., oC A 73 20.2 0.3 2.9 33.9 0.8 7.2 3.2 559 B 73 20.2 0.5 1.6 29.7 1.1 3.5 3.3 586 C 74 20.6 0.6 2.0 34.7 2.4 4.4 4.0 571 D 78 22.7 0.3 1.6 30.6 1.3 3.8 4.0 580 E 72 19.4 0.2 1.8 32.7 0.5 4.3 2.4 582 F 78 20.3 0.4 1.4 33.8 2.3 3.2 6.8 577 G 72 18.6 0.1 1.5 35.9 0.5 3.9 5.5 574 H 74 17.7 1.2 3.8 20.7 6.3 9.3 12.4 513 I 65 20.4 0.5 1.9 26.7 2.9 4.3 5.6 565 J 67 18.5 0.7 3.5 34.4 3.4 8.5 2.3 557 K 71 19.6 0.4 3.5 32.7 1.0 9.0 0.7 567 L 67 18.6 0.2 3.3 26.5 0.1 9.0 2.2 558 M 70 20.0 0.6 2.3 25.0 3.8 6.1 9.6 542 N 64 16.7 1.8 7.0 25.1 5.1 17.6 0.5 610 O 65 17.7 0.3 2.0 31.4 1.7 6.0 6.1 551 P 65 18.6 0.6 2.1 41.5 2.0 6.5 0.3 577 Q 57 19.1 0.4 0.1 26.1 0.8 0.3 0.2 634 W 81 21.1 0.3 2.7 34.2 0.9 6.4 10.4 536 X 20.5 1.7 1.4 32.8 7.9 4.6 0.6 584 Y 73 20.3 1.6 2.2 28.1 7.2 5.4 11.1 536 Z 70 24.2 0.1 3.6 23.5 0.2 6.9 0.0 578 Å 21.1 0.2 0.9 30.2 0.7 1.9 8.7 591

Five of the 20 recovery boilers (excluding boilers H and J) in Table IV consistently have problematic

deposits in their boiler banks. In comparing the four possible predictor variables (T0,f, xCl,f, ∆T15,f, and σ2hr)

for these five boilers, we found that

• xCl,f > 2% in all five boilers,

• σ2hr > 5 MPa in four of the five boilers,

4

• in two of the boilers, Tg,BB>>T15,f, and

• in one of the boilers, T0,f < 537oC.

It is clear from these comparisons that boilers N, X, and Y have problematic deposits result from the high

chloride content of their dusts and the high gas temperatures in their boiler banks. Both are both high

enough so that fume deposits sinter rapidly. The high values of σ2hr for these three boilers support this

conclusion.

It is likely that boiler C has problematic deposits because of the relatively high gas temperature at its boiler

bank entrance (644oC). This boiler has the highest boiler bank entrance temperature of the 22 boilers in

Table IV. ∆T15,f for this boiler is 66oC. Fume deposits in its boiler bank probably densify by viscous

sintering [20] of the partially molten deposits.

The chloride content of the fume in boiler F fume (xCl,f = 2.3%) was the only one of the four variables that

would have indicated problematic boiler bank deposits.

These results suggest that multiple indicators should be considered when predicting whether a recovery

boiler will have problematic fume deposits its boiler bank, or when developing strategies to reduce boiler

bank deposit problems by modifying the composition of the black liquor. Multiple indicators are necessary

because there is more than one cause of problematic boiler bank deposits. A viable strategy would be to

consider three variables: xCl,f, ∆T15, and σ2hr. Fume composition is readily measurable for existing recovery

boilers; all three variables can be calculated easily from black liquor composition, recovery boiler thermal

load, and excess air.

Using xCl,f, ∆T15,f,, and σ2h as predictor variables, the predictions were correct and unambiguous 80% of the

time for boilers both with and without problematic boiler bank plugging problems.

The remaining 15 boilers in Table IV did not report consistent, problematic boiler bank deposits. For 12 of

these boilers, all of the possible indicator variables indicated that they would not have problematic boiler

bank fume deposits. For the other three non-problem boilers, the indicators were mixed:

• Boiler I met all criteria except T0 < 537oC

• Boiler M met criteria of Tg,BB >> T15,f and high xCl,f, but Tg,BB was low (575oC; 4th lowest of the 22

boilers)

The dusts in both boilers I and M had higher chloride content than the dust in boiler F. It is not clear why

2

boiler F would have problematic boiler bank fume deposits than boilers I and M.

Boiler W met only the T0<537oC, and only by 1oC. The chloride content of its fume was relatively low

(0.9%).

A second comparison was made of the four different variables (∆T15,f,, T0, xCl,fume, and σ2hr) used as

predictors of problematic fume deposits in boiler. This time, the fume particle compositions used

were calculated from the black liquor compositions, not the measured values. In some cases, different

boilers were predicted to encounter boiler bank plugging problems. The boilers for which at least one

or more different predictions were obtained and the variables that predicted differently are shown in

Table V. There were some differences in the predictions for boiler-predictor combinations out of 80

(19%). Among all four variables, the prediction of boiler bank fouling problems was improved for

boilers N and Y, both of which encounter fouling problems at normal loads. For boilers A, B, K, L, M,

P, Q, W, and Z that do not encounter boiler bank problems, the predictors that changed all indicated

that the boilers would have problems. These results suggest that accuracy is sacrificed when using

fume compositions calculated from black liquor composition.

We examined the differences in the calculated versus measured T0 and xCl,f to determine why the

predicted fume deposit characteristics differed when based on measured versus calculated fume

composition. In the two cases where predictions based on xCl,f differed, the calculated xCl,f exceeded

the measured value. The same trend in calculated versus measured chloride was also noted for the five

cases where predictions based on σ2hr differed. All of these boilers had low chloride (0.1-1.1 wt-%) in

their ESP catch. The differences in chloride content, between calculated and measured, were from

0.5% to 1.3%. Four of these boilers were firing black liquor at 67-73% solids content, a range where

the SO2 concentration in the combustion gases in the upper furnace and superheater region might be

high enough to convert NaCl to Na2SO4 and HCl. This suggests that SO2 concentration in the

combustion gases may need to be included in the fume composition model to improve its accuracy.

2

Table IV. Comparison of measured fume parameters with boiler bank fouling tendency.

Boiler Tg,BBa, oC T0,f

a, oC

T15,fa, oC

Tg,BB – T15,f, oC

Clfume, wt-%

Bending strength

(σ2hr)a,b MPa

Problematic boiler bank

fouling? A 616 559 722 -106 0.8 1.1 no B 622 586 712 -91 1.1 2.3 no C 644 571 578 66 2.4 30.7 yes D 587 580 684 -97 1.3 0.7 no E 594 582 776 -182 0.5 0.4 no F 568 577 586 -18 2.3 0.8 yes G 573 574 764 -191 0.5 0.3 no H 587 513 539 48 6.3 30.7 no I 619 565 572 47 2.9 15.0 no J 611 557 564 47 3.4 14.7 no K 585 567 703 -118 1.0 0.6 no L 578 558 793 -216 0.1 0.2 no M 575 542 553 22 3.8 2.0 no N 580 610 591 -11 5.1 5.8 yes O 579 551 613 -33 1.7 0.8 no P 594 577 611 -17 2.0 2.0 no Q 587 634 779 -193 0.8 0.5 no W 597 536 684 -87 0.9 0.7 no X 566 584 589 -23 7.9 5.0 yes Y 608 536 547 61 7.2 30.7 yes Z 575 578 805 -230 0.2 0.2 no Å 566 591 760 -194 0.7 0.3 no

aCalculated based on measured composition of electrostatic precipitator catch. bBased on sintering of the deposited material at an initial porosity of 0.11 for two hours at Tg,BB.

3

Table V. Boilers for which different predictions of problematic fume deposits in boiler banks were obtained, and the variables for which the predictions differed.

Boiler Tg,BB - T15,f T0 XCl,f σ2hr

A +

B + + +

K + +

L +

M +

N + +

P +

Q +

W +

Y +

Z +

Deposition of Partially Molten Intermediate Size Particles

So far, we have considered only boiler bank fouling by deposits of fume particles. However, boiler

bank fouling can also be caused by deposition of partially molten intermediate particles. The criterion

for problematic intermediate deposits in the boiler bank is that the particles arrive partially molten

(T15,ip < Tparticle) and cool to below the lower sticky temperature (T15,ip) after they have deposited. The

compositions of the intermediate size particles were calculated from Equations 1-4 and 19, using

enrichment factors for potassium and chloride of 0.68 and 0.49, respectively. Their lower sticky

temperatures (T15,ip) were calculated from their compositions using Equations 22 and 23. The

calculated lower sticky temperatures were compared with the estimated gas temperatures at the boiler

bank entrance from Table IV. The results, in Table VI, are that for all of these boilers the predicted

gas temperature entering the boiler bank (Tg,BB) exceed substantially the lower sticky temperature

(T15,ip). The smallest difference was 92oC, and on average, the difference exceeded 200oC. Sticky

intermediate particles could not have caused deposition problems in any of these boilers.

4

Table VI. Comparison of Tg,BB with T15,ip.

Boiler Kip, wt-% Clip, wt-% CO3,ip, wt-% Tg,BB, oC T15,ip, oC T15,ip – Tg,BB, oC

A 3.4 0.3 32.3 823 616 -207 B 2.0 0.5 35.2 814 622 -192 C 2.4 0.5 31.9 804 644 -160 D 1.8 0.2 34.9 836 587 -249 E 2.3 0.2 33.4 838 594 -244 F 1.7 0.4 32.8 824 568 -256 G 2.0 0.1 31.7 849 573 -276 H 5.1 1.2 39.4 733 587 -146 I 2.4 0.5 37.0 813 619 -194 J 4.4 0.6 31.3 780 611 -169 K 4.2 0.3 32.7 810 585 -225 L 4.3 0.2 36.8 827 578 -249 M 2.9 0.5 37.8 801 575 -226 N 8.7 1.6 35.0 672 580 -92 O 2.8 0.3 34.0 823 579 -244 P 2.7 0.6 27.7 797 594 -203 Q 0.1 0.4 38.0 832 587 -245 W 3.1 0.3 32.2 826 597 -229 X 1.7 1.5 32.4 724 566 -158 Y 2.7 1.4 35.1 725 609 -116 Z 3.7 0.1 39.0 842 575 -267 Å 1.1 0.2 35.4 847 566 -281

CONCLUSIONS

A new method was presented for predicting whether fume deposits in the boiler banks of kraft

recovery boilers will be problematic. The method is based on the sintering behavior of sub-micron,

alkali metal salt aerosol particles (fume) that deposit on the heat transfer surfaces of these boilers.

Application of the method presented to 22 recovery boilers for which extensive data was available

showed that problematic deposit formation in their boiler banks correlated strongly with the chloride

content of the fume particles (xCl,f) and the strength expected to develop upon sintering (σ2hr). The

expected strength development is a unique function of the first melting temperature of the deposited

particles, their chloride content, and the temperature at the location of the deposit. The temperature of

the gas entering the boiler bank, when it exceeded the lower sticky temperature of the fume particles,

5

may have led to viscous sintering of the resulting deposits. The SO2 content of the combustion gases

may be important since it can influence the chloride content of the particles and the resulting deposits;

however, insufficient data was available to determine its importance.

Deposition of sticky intermediate size particles in boiler banks did not appear to be problematic in any

of the recovery boilers included in this study, since the estimated gas temperature at the boiler bank

entrance was always substantially less than their lower sticky temperature.

The method presented here, for predicting whether fume deposits in the boiler banks of kraft recovery

boilers will be problematic, predicted correctly and unambiguously 80% of the time, both for boilers

that had these problems, and for boilers that did not. A combination of predictor variables, using xCl,f,

∆T15,f, and σ2hr, needs to be used because of the several mechanisms by which fume deposits can

densify and harden in boiler banks.

ACKNOWLEDGEMENTS

Funding for this research was provided by the Office of Industrial Technologies of the U.S. Depart-

ment of Energy, and by the Recovery Boiler Research Consortium at the University of Toronto.

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