COOLING CHARACTERISTICS AND THERMAL PROPERTIES OF KRAET RECOVERY BOILER SMELT

Geng Tan

A thesis submitted in conformity with the re~uirements for the degree of Master of Applkd Science

Graduate Department of Chedcai Engineering and Appiied Chemistry University of Toronto

O Copyright by Geng Tan 2ûûû

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COOLING CEARACTERISTICS AND TEERMAL PRûPERTIES

OF KRAFT RECOVERY B O m R SMELT

Master of AppIied Science 2ûûû

Geng Tan

Department of Chernid Engineering and AppUed Chemistry

University of Toronto

ABSTRACT

An accurate prediction of smelt thermal state is crucial in the cooling of kraft recovery boiler

smelt in order to prevent smelt-water explosion and shorten the cooling time. The smelt

cooling process was systematically investigated in this study. A laboratory-sale

experimental apparatus was constructed and used to conduct field experiments in a kraft mil1

with a mnning decanting floor recovery boiler. The molten smelt samples were collected

from a smelt spout, and cooled in the experimental apparatus. By measuring the intemal

temperature histories, the cooldown characteristics were obtained and analyzed. The results

showed that smelt cooling involves a slow heat transfer pmcess and is mainly controlled by

the rate of heat conduction within the smelt interior due to the Iow thermal conductivity of

smelt. A transient 3-dimensional heat transfer mode1 was also developed to analyze and

predict the smelt cooling process. The results showed reasonably goad agreement with the

experimental data. Both experimental and simulation results indicated that the thermal

conductivity of fiozen smelt is well around 0.6 W/m°C.

1 would Like to express my s k r e thanks to Professors H. N. Tran and M. Kawaji for theu

invaluable guidance, supervision and encouragement throughout the course of this study.

I would like to thank Mr. A. Tavares for his help with my field experiments. Special thanks

are given to Drs. V. Agranat and A. Gofinan for their advice on the modelling setup and help

in my field expenments. 1 am grateful to Dr. S. Kochesfahani for his valuable suggestions

and help in my tests. I also thank Mr. D. Tomchyshyn for his assistance in the setup of the

data acquisition system.

The financial support frorn Amencan Forest & Paper Association and the memben of the

"Irnproving Recovery Boiler Performance, Emission and Safety" research consortium. is

gratefully acknowledged.

And finally, 1 am indebted to everyone in my family for their endless encouragement and

support.

TABLE OF CONTENTS

ABSTRACT,

ACICNOWLEDGEMENTS

TABLE OF CONTENTS

LIST OF TABLES

LIST OF FIGURES

NOMENCLATURE

CEAPTER 1 INTRODUCTION

1. I Bac kground

1.2 Objectives and Approaches

CHAPTER 2 LrmalrrqE REVLEW

2.1 Char Bed process in Recovery Boilers

2.1.1 Design of recovety boiler lower fumace water tubes

2.1.2 Formation and characteristics of char bed

2.1.3 Srnelt-water explosion and initiation of emergency shutdown

procedure (ESP)

2.2 Char Bed Cooling process Following an ESP

2.2.1 Bed conditions

2-2.2 Mechunism of char bed heut tronsfer

2.2.3 Monitoring char bed co&g

2.2-4 Decision making on carrying out water-washing

2.3 Tests and Modelling of Char Bed Cooldown Process in Recovery Boikrs

iv

viii

ix

xii

1

1

4

5

5

5

6

CHAPTER 3 EXPERIMENTAL STUDY

3.1 Experirnental Apparatus

3.1.1 Smelt cooling vesse1

3.1.2 Watercwlingmit

3.1.3 Air cooling unit

3.1.4 Imh~.mentotion and data acquisition system

3 2 Experirnental Procedures

3.3 First Field Test

3.3.1 Procedure

3.3.2 Results and discussion

3.4 Second Field Test

3.4.1 Procedure

3.4.2 Results and discussion

3.5 Measuring Thermal Conductivity of Frozen Smelt

3.5. l Nature of thermal conductivity

3.5.2 Principle of measuremertt method

3.5.3 Procedure

3.5.4 Results and discussion

CMFI'ER 4 NUMERICAL SIMULATlONS

4.1 Description of Simulation Modei

4.2 Geometry of the Computational Domain

4.3 Heat Tmsfer Equations

4.4 Determination of Boundary Conditions

4.4.1 Forced air convection on ajlat plate

4.4.2 Forced convection to water Ui cooling tubes

Consideration on Input Smelt Thermal Propenies

Resuits and Discussion

Pararnetric Studies

Validation of nwnerical calculution

Emissiviîy ut the smelt top surface

Heat capacity of smelt

Latent heat of smelt

Impact of t h e m l conductiviiy of smelt

Effect of boundary conditions

Summary of parumetric studies

CHAPTERS' SUMMARY

CEAPTER 6 RECOMMENDATIONS

6.1 Remaining Uncertainties

6.2 Recommendations

REFERENCES

APPENDICES

Appendix A Emergency S hutdown Procedure (ESP)

Appendix B Raw Experimentai Data

Appendix C Smelt Temperature Distribution at Different Depths

Appendix D Properties of Castable Powder

Appendix E PHOENICS Input Data File

Appendix F Redicted Srnelt Temperature Distribution at Different Depths

after 3 Hours of Cooling

Appendix G Effect of Parameters on the Smelt Cooling Process

vii

Table No,

LIST OF TABLES

Descriotion

Physical and thermal properties of black liquor char and smelt.

Water sources for smelt-water explosions.

Mass and heat content of char bed in a lOmx 1Om furnace cross-section boiler.

Average smelt temperature in the last one hour of the experiment.

Properties and temperature rise of cooling water during the experiment.

Approximate values of convection heat-transfer coetficients.

Reynolds number in the water tube under different cooiing conditions.

Principal input data in numerical simulation.

Summary of the effect of parameters on the predicted smelt cwling process.

LIST OF FIGURES

Fimire No.

1.1.

2.1.

2.2,

2.3.

2.4.

2.5.

3.1.

3 -2.

3.3 .

3.4.

3.5.

3.6.

Description

- Schematic diagram of a kraft recovery biler.

Lower furnace wall construction.

Surface shapes of char bed.

Char bed shapes.

Onedimensional heat-flow mode1

Temperature vs. time at given depths.

Smelt cooling experimental apparatus.

Water cooling coil.

Page

2

Location of cooling coils placed into the smelt cooling vessel. 28

Smelt spout with mnning molten smelt.

Smelt sampling.

Air cooling operation. 3 1

Smeh temperahue at various depths. 33

Top-view of the smelt top surface after 3 houa of cooling. 34

Effect of watercooled side wail at 15 cm (6')below the smelt top surface. 36

Smelt sarnpling (second test). 39

Top-view of molten smelt pool. 40

Thermal insulation. 40

Smelt temperature profiles at various depths within the smelt pool. 42

Smelt cooling in the first 30 minutes. 42

Fiare No. .

3.15.

3.16.

Descri~tion Page

Measurement of thermal conductivity of frozen smelt. 47

Smelt temperatures in the middle of the experimental vesse1

(O. 1Sm into the smelt pool) approaching constant values. 49

Smelt temperatures at the location 0.07Sm into the experimentd pool

approaching constant values. 49

Schematic of the computational domain. 55

Sketch showing diffennt boundary-layer flow regions on a flat plate. 59

Comparison of predicted and measured temperatures at various depths. 65

Comparison of predicted and mensured temperatures at various distances

from the water-cooled side wall. 67

Predicted temperature distribution after 6 hours of cooling . 68

Cornparison of various thermal conductivities for each phase with the

standard case, 72

Effect of heat transfer coefficient at the water-cooled walls, hi. 74

Smelt temperature distri bution at 5 cm (2") below the top surface. C-2

Smelt temperature distribution at 25 cm (lû") below the top surface. C-2

Temperature factor plot, D-3

Predicted smelt temperature distribution at the top surface. F- L

Predicted smelt temperature distribution at 5 cm (2") below the top

surfiace.

Predicted smelt temperatwe distribution at 15 cm (6") below the top

surface.

Figure No. Descri~tion

G.I. Identical nsults with constant p C , but different p and C,

G.2. Effect of emissivity.

(3.3. Effect of heat capacity.

G.4. Effect of latent heat.

(3.5. Effect of iow thermal conductivity.

G.6. Effect of high thermal conductivity.

G.7. Effect of heat transfer coefficient at the uncooled walls. 4.

Page

G-2

G-2

G-3

G-3

G-4

G-4

G-5

NOMENCLATURE

cross-sectional area of heat iransfer, m2

heat capacity of material, Jkg°C

inner tube diameter, m

rate of temperature change, 'Us

heat-tram fer coefficient, w/mZ O C

thickness of material, m

thermal conductivity, Wlm OC

length of plane, m

mass of material, kg

calculation exponent

Nusselt number

Prandtl number

heat flux, wlm'

heat generation (removal) rate

Reynolds number

duration time, s

temperature, OC (K)

velocity, mls

volume flow rate, m3/s

K distance, m

temperature ciifference, O C

ernissivity

latent heat (heat of hision). J/kg

viscosity, kg/rns

density, k@m3

2 4 Stefan-Boltzmann constant. = 5.67 x lu8 Wlm K

heat flux vector

temperature gradient vector

at the water-cooted walls

at the uncooled walls

arnbient atmosphere

smelt close to the boundary of the experimental apparatus

flow in a tube

frozen smelt

initial condition

at water inlet

mean value

at water outlet

smelt at the middle of the experïmentai apparatus

xiii

S

SUC

t

W

X

smelt at the melting point

at the smelt top surface

floor tube

water

local pmperty

xiv

1.1 Background

Paper and paper products play a very important role in our society. The usage of paper

can be found in every m a of modem life. Almost ail hard copy, writing and printing is

accomplished.on paper. It is also an exceiient wrapping and packing material and is important

for structural applications. Canadian pulp & paper industry's 1998 sales totaled 52.6 billion

CAD (Williamson et al., 2000). With the help of modem techniques and technologies the

pulp and paper mills are now highly automated to produce low cost and high quality

products. which is crucial for survival in the worldwide market cornpetition. Therefore, the

overall economics ofhm favors large-scale units with high productivity. As a nsult, the

construction cost for building a modem mil1 is extremely high. A new 1ûûû tondd kraft pulp

mil1 may cost more than I billion US dollars to build.

Paper products are made from wood pulp. Currently, more than 70% of pulp is

produced by the kraft pulping process.

A recovery boiler is one of the key units in the production of kraft pulp. Its stable

operation detennines the production efficiency and productivity. A schematic diagram of a

typical recovery boiler is shown in Figue 1.1. There are more than 3ûû operating recovery

Mers in North Arnerica Smook (1994) summarized that the essential hinctions of a

recovery boiler are to burn organic chernicals to recover heat energy, and to recover inorganic

chernicals which are used in pulping process.

Figure 1.1. Schematic dlagram O€ a kraft recovery boiier ( A h et ai., 1997).

b recovery boilers, the floor is covered with molten smelt, mainly sodium carbonate

and sodium sulfide, at 700-800°C during operation. For the purpose of energy recovery, the

floor and side walls as well as other heat exchanging components on the upper fumace of a

recovery biler are al1 made of water tubes. If a tube leakage occurs due to corrosion or

mechanical damages. the water in the tubes will corne in contact with the molten smelt and

evaporate rapidly. As a consequence, a smelt-water explosion may occur. There have been

mon than 140 recovery boiier explosions in North America because of the smelt-water

contact (Green et al., 1992). The damages caused by the explosion have ranged h m minor to

disastrous. Some explosions almost completely destxoyed the lower himace and resulted in

death of site personnel.

When a tube leakage occurs, the boiler is shutdown immediately for an emergency

shutdown procedure (ESP). Water is rapidly drained and depnssurized, while the char k d is

let to cool. Once al1 the smelt has solidifïed, the char bed is washed with water completely so

that the leakage can be found and repainxi.

The ESP often leaves a hot char bed on the Lower furnace of the boiler, which consists

of char, frozen and molten smelt. There is a potential that during the washing, water may be

in contact with molten smelt and cause smelt-water explosions. Thus, the water-washing

operation must be camied out with extreme caution. A proper judgement of the thermal state

of the smelt before water-washing is crucial.

Unfortunately, it is often difficult to make the decision on whether the char bed has

cooled down sufficiently so that water-washing can be safely carried out. For safety Rasons,

boiler operators often let the smelt cool until they feel that based on experience the smelt

temperature is cold enough to perform water-washing. The waiting pend is usualiy

conservative, which may mean a few days of unnecessary cooling and therefore a significant

loss of pulp production. Although there is a substantial amount of experience associated with

cooling of char beds following ESP's. a widely accepted guideline is not avaiiable. This

problem has been a concem to the industry. A nliable method to preâict the char bed cooling

process is nquired to help ensure the safety of initiating water-washing as well as shomning

the shutdown tirne, so as to minirnize the loss of production.

1 Objectives

The objective of this snidy is to obtain an understanding of the char k d cooling pmctss

so that the themal state of the char bed can be nliably predicted. The scope of the study

includes:

(1 ) Designing and consüucting a srnelt cooling apparatus;

(2) Conducting field tests using the apparatus;

(3) Analyzing smelt cooling data;

(4) Developing a heat transfer mode1 to predict the smelt cooling process;

(5 ) Examining the impact of critical heat transfer parameters on the smelt cooling rate.

2.1 Char Bed Process in Recovery Boilers

2.1.1 Design of movery bi ler lower f'urnace water tubes

The lower fumace walls of a recovery boiler are constnicted of vertical tubes set in a

row. A general configuration of the wails is shown in Figure 2.1. These tubes are typically

from 6.4 to 7.6 cm (2.5 to 3 inches) in diameter. In modem recovery boilers the tubes are

spaced between 1.25 and 2.5 cm (0.5 to 1 inch) apart, connected by a flat membrane. The

membrane is continuously and fully welded dong the adjacent tube lines to make a gas- and

smelt-tight enclosure.

The rows of tubes, so called waterwalls, form the four fumace walls. Water flowing in

the waterwaiis receives heat by radiation and by conduction from the char bed, as well as

h m names in the funiace. The waterwalls can supply almost half of the heat transfer area for

obtaining high-pressure stem (Adams et al., 1997).

In the lower fumace, the watewall tubes are exposed to hostile reducing conditions.

Sulfidation by reduced sulfur gases is the main reason for tube corrosion. In old designs,

tubes are made of carbon steel. At temperatures above 3 lS°C (600°F) the carbon steel tubing

is subject to accelerated wastage. The most common protection method for waterwall tubes in

the lower h a c e is to fabricate the tube with a 1.65 mm stainless steel sheath on the outside

of a carbon steel tube. Composite tubes are maintenance-fke for a nlatively long-tem in

nomai operation, except for the problems in some specific local areas (Green et al., 1992).

Figure 2.1. Lower furnace w d l co~tmction.

2.1.2 Formation and characteristics of char bed

Black liquor is sprayed into recovery boiler fumace and then burned. Char forms from

devolatilization of the black liquor. It consists of residual organic carbon (fixed) dong with

inorganic materials from the pulping chemicals. Some hydrogen and oxygen are always

bound to the fixed carbon. The char pnsents a porous structure provided by the carbon. High

temperature from the black liquor combustion is enough to bum the char, and to convert the

inorganic materials into molten smelt which is predominantly sodium-based salts. As the char

is consumed, molten smelt accumulates on the fumace Ooor covered by the burning char. The

char bed can form various shapes depending on the boiler size and operation conditions. Hat

beds and coaicai beds are aii often observed as illusttated in Figure 2.2.

Char bed consists of carbon, partially pyrolyzed black liquor solids, molten and

solidified smelt (sodium salts and sulfuric compounds). which covers the entire floor area.

The char bed size is determined by operating conditions. The average bed height is about 1-

2m (3-6 feet) high. In some cases, over 4m (13 feet) high char bed may be present in at least a

part of the furnace. The char bed height is not fixed in the fumace. At some local areas and

corners the char bed may be considerably higher than the rest. At extreme situations a

decanting bottom himace is operated with only a molten smelt pool. and a slant floor furnace

with a negligible char bed on the floor.

The char bed CM be mainly divided into two layers. A hot. aaive buming layer at the

bed surface supportcd by a colder, unceactive bed ôelow i t The active layer is typicaliy 15-

2Ocm (6-8 inches) thick. The temperature decrcases dramatically, from lûûû°C to 1200°C at

the bed surface to approximately 760°C at the bottom of the active char layer. The bed is

celatively imperneable to combustion. The sharp temperature gradients fiom the bed surface

inward are characteristic of char beds in recovery boilers. Below this active layer is a more

dense, chernicaily inactive core of bed which is below the inorganic melting point.

The active layer presents the same characteristics in al1 recovery boilers. On the other

hand, the characteristics of inactive layer Vary greatly in diffennt boiler types. Figure 2.3

shows typical shapes of char bed in diffennt boiler designs.

In slruited floor boilers the char bed is a single-mound with a relatively fiat top. The

inactive part of the char bed is far denser and less penneable than the active layer. Molten

smelt is formed in the active layer and flows in smelt channels on the char bed. However, in

decanting-hearth boilers the char bed is nlatively low. The char is very porous and mobile,

and molten smelt penetrates the char bed. The bed sits in a pool of molten smelt over the

whole floor, on top of a frozen smelt layer which is against the flmr tubes.

Some of the physical and thermal properties of char and smelt cumntly used in

industry are listed in Table 2.1.

Table 2.1. Physicai and thermal properües of black Uquor char and

smelt (Aàams et ai., 1997).

Active Zone inactive Zone Molten SoUàiîied Smelt Smelt

Density, kg/m3 290460 480- 1330 1923 2 163 Heat capacity, kJ/kg/K 0.25 1.25 1.34 1 -42 Thermal conductivity, WlmK 0.28-0.38 0.078 0.45 0.88 Thermal diffusivity, m21s 0.54 .0x lob 0.5-0.7Sx 10" 1.8~10-' 2 . 8 ~ IO*?

Heat of fusion. W/kg - - 142 -

\ ait

Figure 2.3. Caor bed sbapes (Aduao, et ai., 1997).

2.1.3 Smelt-water explosion and initiation of emergency shutdown pmceàure (ESP)

Recovery boilers have most of the safety problems existing in other industrial boilers.

Moreover, they are also subject to the unique hazards of smelt-water explosions due to the

pnsence of molten smelt within the lower himace.

During operation, a char bed covers the floor of the lower furnace. The char bed

contains a considerable amount of molten smelt. For any nason when there is water coming

into contact with the molten smelt, a smelt-water explosion may occur. Adams et al. (1997)

summ~zed the sources of water for the smelt-water explosions that occumd fiom 1970

through 1995. The results are tabulated in Table 2.2.

Then have ken mon than 140 recovery boiler explosions in North America due to

smelt-water contact (Green et al., 1992). Some explosions nearly completely destroyed the

boilers and some of more severe explosions even caused casualties. niete have been several

hundred emergency shutdowns performed due to possible explosions.

Shick and Grace (198 1) began to investigate the nature and features of smelt-water

explosions. The smelt-water was explaincd as a kind of stem explosions containing huge

amounts of mechanical energy. Consequently, Grace (1986, 1992, 1997 and 1999) presented

a number of articles to introduce the development on the understanding of smelt-waier

explosions.

Grace (1999) nviewed the recovery biler explosion experiences in the US and

Canada over the last 35 years and conclucied that the smelt-water explosions remain the most

common type of recovery boiler explosions, and the primary source of water for these

explosions hm k e n boiler pressure part faüures. He summarized that smelt-water explosions

genedy nsult h m failures that introduce large amounts of water ïnto the h a c e cavity.

The exceptions are flmr tube leaks and, to a lesser extent. spout leaks and smaii wall tube

Ieaks near the hearth.

Grace believes that smelt-water explosions are caused by extremely rapid generation

of vapor (stem). As a result of this vaporization, the rapid volume expansion increases the

local pressure dramaticaily. A blast energy is accumulated and then released as a shock wave

nther thui a simple overpnssurization of the fumace. The interaction of molten smelt with

water is violent. It has been estimated that the mechanical energy released frorn vaporization

of one pound of water in 0.00 1 second can be equal to an explosion of 0.25kg (0.5 pound) of

TNT dynamite (Green et al., 1992). Typically from 2.5 to 12.5kg (5 to 25 pounds) of water is

involved in an explosion. In some extreme accidents nearly 50kg (100 pounds) of water is

involved. The amount of water coming into contact with molten determines the extent of

himace explosion and damage.

Tabk 2.2. Water sources for smelt-water explosions (1970 - 1995) (Adam et ai, lsw).

Magnitude Explosion Water Source Amount of Water Major Moderate Minor

12 Wall Tubes Law 3 5 O Srnall ieaks O O 4

8 Screen Tubes 4 2 2 2 Boiler Bank Tubes Large O 1 1 1 Roof Tube Large 1 O O 1 Superheater Tube Large (Boiler Refilled) 1 O O 7 Floor Tubes Large (Two Ruptures) O 1 O

Relatively Srnall 3 3 O 5 Spouts h w Pressure O 2 2

Pressurized I O O 8 Black Liquor System Black Liquor 3 2 O

Wash Watet 2 O 1 4 Wash Water Miscellaneous Large 3 O 1

A common m o n for water coming into hünace is tube leaks (pressure-part failtues).

Floor tubes and waterwall tubes are attributed to bc the most iikely locations where the le&

take place. If the leaks can not be stopped irnmediately and the situation continues, an

emergency shutdown must be carried out. The detail of the emergency shutdown procedure

(ESP) is described in Appendix A.

The npid dnining step is unique to recovery boilen. Experience has show that the

rapid draining has effectively reduced the amount of water that enters the unit and contacts

with molten smelt so as to minimize the possibility and violence of smelt-water explosions.

The 2m (8 feet) of water is left to protect the boiler from the heat of char bed. Several

hundred emergency shuidowns following an ESP have been performed. Most of them

avoided any darnage successfully.

2.2 Char Bed Cwling Proeess FoUowing an ESP

Cooling and solidification are common processes, panicularly in the metal casting

industry, when molten metals are cast into useful objects. In recovery boilers the dominant

components of smelt are sodium-based salts (Na2S and Na2C03). Therefon. the cooling of

char bed involves the cooling and freezing processes of molten salt.

2.2.1 Bed conditions

An ernergency shutdown leaves a hot char k d on the hearth of a recovery boiler. The

initial condition of the char bed is determined by the boiter design and operating conditions

pcior to the ESP. Its size and shape Vary widely. Molten smelt can be prescnt in different

locations and amount depending on the design of the lower h a c e . Chernicd ceactions and

local combustion may continue for a short time. The frozen smelt falling to the char bed €rom

other components in the upper furnace of the boiler will also affect the char bed conditions.

These variables make it impossible to set a fixed waiting time pend for the char bed cooling

process.

It is normally believed that smelt solidifies at about 700°C (1 300°F). The goal of char

bed cooling process is to solidify al1 the smelt within the char bed so that the risk of smelt-

water explosion can be eliminated when canying out water-washing. Heat contained within

the char bed must be transfemd either upward from the top surface or downward to the flwr

tubes. Before the adoption of an ESP the char bums in the lower furnace. The top surface of

the char bed is hot so that the heat is transfemd into the kd . The temperature ranges from

800°C to 12û0°C (1500v to 22OOOF) at the top surface. Near the bottom the bed temperatun

is close to the saturated temperature of steam within the watenvalls which is about 270°C

(5200F). Following an emergency shutdown, the water within the waterwalls is drained io &

foot high. Natural and forced convection of gases mostly cools the top surface of the kd. The

floor temperature decreases to near the normal water boiling point of lOO0C. Al1 of these

changes can occur in a nlatively short time period.

2.2.2 Mechanism of char bed heat t m e r

The char bed contains a large amount of heat. For a typicai ncovery boiler with a

fumace cross-section of 10m by 10m (32.8 feet by 32.8 feet), the estimated values of the char

bed mass and heat content are iisted in Table 2.3.

The heat value of carbon oxidation is 32,850 kJ&g (14,120 Btu/lb°C) as pcesented by

Richardson and Memam (1977 and 1978). The potential heat celease fmm oxidation of

sulfide is based on complete oxidation to sulfate with a heat of reaction of 12,900 W/kg

(5,550 Btu/lb Na2S), with a smelt sulfidity of 30%. and a reductioa eficiency of 90%.

Table 2.3. Mass and heat content of char bed in a 10m x 10m hrnace cross-section

boUer (Kandi et ai., 1999).

I Bed Heieht. m 1 2 I l l 1 f 1 Porosity 1 0.5 1 0.5 1 0.75 1

Table 2.3 shows that the bed height and porosity can affect the amount of heat

required to be removed. The potential heat release cornes from continued combustion of

sulfide and carbon in the char bed, which could account for aimost 80% of the total heat

release. Therefore, stopping al1 combustion in the char bed becornes a very important factor

For the bed cooling pmcess after an emergency shutdown. The sensible heat of smelt is

another major contribution. On the other hand, the latent heat of smelt solidification is only a

small fraction compared to the sensible heat of cooling the bed from 81S°C to s a . Even

with a relatively small porous bed this fraction is only Iess than 20%.

The heat transfer within a char bed can be internai and extemal. The internai heat

tmnsfer is the heu content of the char bed flowing out to the boundaries. The combustion of

char stops shortly after an ESP. The smelt stops flowing, and hence there is little convective

Mas, metric tons 1 212 106 112 72 53

0.75 3.8 4.1 173 69- 274

Sensible heat at 8 HOC (15000F), GJ Sensible heat at S40°C ( 10009), GJ

53 56 36 26

0.38 1

1.9 l

4.1

87 -- -~ 35 137

223 145

Sensible heat at 400°C (7500F), GI Heat of fusion (5% molten smelt), GJ Heat of fusion (25% molten smelt), GJ Heat of fusion (6 in. smelt pool), GJ Heat of Carbon Oxidation (5% C), GJ Heat of Carbon (2% C), GJ Heat of Sulfide Oxidation, GJ

105 1.5 7.5 4.1 347 139 548

heat minsfer. Most of internai heat is released by thermal conduction in the smelt. Since the

char bed is porous, heat c m also be released by radiation through pores at high temperatun.

The heat of char bed is finally released by extemai heat transfer, which includes

convection to the gases above the bed, radiation fiom the bed surface to the mrrounding

wails and upper fumace, and conduction to the flwr (and side walls), possibly assisted by

boiling, and recondensation of the residual water remaining in the floor tubes.

In a typical recovery boiler, the bottom cross-section is about IOm by IOm. After an

ESP the char bed height is normally less than lm. The side wall-bed contact ana is much less

than the anas of the top and bottom. Heat transfer from the side walls only accounts for a

very small fraction compared to that from the top and bottom. Thenfore this part of heat

transfer is usually neglected in the practical heat transfer analysis. But in the cornen the heat

uansfer from the walls is still important.

Dunng the char bed cooling process, the nlease of latent heat by solidification of the

smelt is nlatively small compared to the sensible heat of the bed. Grace (1998) suggested that

when neglecting the heat loss from the side walls the average heat flux fiom the bed can be

calculated from the bed mass and the rate of temperature decrease. The equation can be

described as:

q = M C, (dTldt) @A) 12- 11

where q is the average heat flux from the k d w/m2], M is the mass of the bed materiai kg],

C, is the heat capacity of the bed materid [J/kg°C], (dT1dt) is the rate of tempemm decrease

[OUs], and A is the cross-sectional arui of fumace [m2].

In decanting bottom recovery boilers the molten smelt foms a pool on the hearth over

a frozen smelt layer which sits directiy on the floor tubes. After an ESP the temperature of the

molten smelt can be consideced to be at the rnelting point. The heat nlease rate (heat flux)

through the floor is then calculated by the following equation (Gri~ce, 1998),

q = kf (Ts-TJrnf, P.21

where kr is the thermal conductivity of fiozen smelt, Ts is the smelt melting temperature, Tt is

the floor tube temperahire, and Hf is the thickness of the fiozen smelt layer.

If it is assumed, as in most cases, that the temperature profiles across the frozen smelt

layer are linear. the amount of heat removed from the frozen smelt cm be calculated by the

equation,

Q = p A Hr [Cpr (TrTtY2 + LI, 12-33

where Q is the arnount of heat removed. pf is the density of frozen smelt, Cpf is the heat

capacity of frozen smelt, and Â. is the heat of fusion of smelt.

Using dQ/dt = qA, combining equations 12.21 and [2.3], and after integration, the

following new equation can be used to calculate the time needed to solidify the molten smelt

layer

ts = ~f [H?-H~'] [cPr (TsœTJ/2+ iC] 1 [2kf (Ts-Tt)], P-41

where t, is the time needed to solidify the molten smelt layer, Hf is the final frozen smelt layer

thickness, and Hi is the initiai frozen smelt layer thickness.

The char bed heat can also be removed from the bed surface by convection and

radiation. The speed of heat removal is controlled by the bed surface temperature.

Heat convection is due to the naniral and forceci gas flow. Heat transfer coefficients

for gases in a naturai convection are typically about 6 w/m2 OC. In forced convection this

coefficient cm be severai times pater. Radiation to the surroundings is an important heat

mlease mechanism from the top surface. In die absence of back dation, a very high black-

body radiant heat flux of 500 w/m2 can be achieved even when the bed surface temperature

is as low as lûû°C (Grace, 1998). Thus, the bed temperature does not need to be very high to

provide high heat removal rate. The high heat flux on the bed surface even when at a

relatively low temperature aiso suggests the influence of bed disruption on bed cooling rate.

Bed dismption may expose the hot intenor matenal of the char bed. This will

dramatically increase the convective and radiant heat transfer rates from the bed surface.

Kawaji et al. (1999) reported that if the exposed bed material is at 650°C (1200%). the

convective heat flux may increase by 5 to 10 times. The black-body radiative heat flux wouid

be 40,900 w/m2 (13,000 Btu/hr h2), which is about 20 to 40 times of normal heat flux from

the k d surface. Thereforr, a bed disruption accompanied by suppression of combustion of

the hot materiai can lead tu very high heat fluxes and high bed cooling rates.

Cooling from both the bed surface and the floor tubes are both important. When

cooling an exposed smeit pool in a decanting bottom recovery biler, heat is removed mainly

from the top. On the other hand, for a large bed, cooling from the bottom is superior to

cooling from the top surface.

2.2.3. Monitoring char bed codiag

There are severai ways to monitor the char bed cooling process. Visual observation is

certainly the easiest and simplest method. Glowing spots or buming anas on the bed are

direct indications of a very hot bed, and the bed netds to be cooled huther. But a dark bed

does not mean that the intenor bed matenal has aiready cooled enough.

The use of thermocouple probes is a common method of monitoring the bed cooling

pnmss. It is the only acceptable technique to collect quantitative information on char bed

themial conditions. The end of a probe can be inserted into the char bed, which gives a direct

reading of the bed temperature annind the location of the probe. The practical problem is that

the thermocouple probe can nach oniy a limited area of the char bed. But other areas,

especially the area in the centrai region of a large bed, where the molten smelt is most Likely

present or some hot spots have been seen before. may not be accessible. The temperanin in

this region would remain unknown.

The infrared video technology has been used effectively to monitor char bed operation

(Richardson and Memarn, 1977 and 1978; Harrison and Ariessohn, 1985). Other optical

devices can also be used to monitor the temperatun distributions on the char bed surface.

Local areas where the temperature is apparentiy higher than other areas clearly indicate the

presence of hot spots. Themocouple probes can then k used to penetrate the bed surface to

measure the temperature of the potential hot bed material. However, this method is not

suitable to measure the end of bed cooling process. Some smail hot spots may still hide in the

bed interior even after a quite low bed surface temperature has already achieved.

23.4 Decision d n g on carrying out wator-washing

A char bed is cooled until the bed temperature at every measuring point reaches or is

lower than some prescribed value. This value is set much k l o w the melting point of smelt.

Diffennt milis adopt different value such as 430°C or Sm (8000F or l,ûûû"F,. Once this

critical value is met, the smelt in the char k d is considered well below the melting point

everywhere and the water-washing can be perfomed safely.

2 Tests and ModeUing of C h r Bcd Cooiâown Process in Recovery BoUus

Char bed cooldown process is an important issue on the operation of recovery boilers.

However. due to the complexity of its heat transfer phenomena, very little work has been

done on conducting tests and modelling of char bed cooling. The only significant

publications that have dealt with test and modelling of char bed cooling were Arthur D.

Little, Inc. (ADL) reports (Richardson and Memam. 1977 and 1978). as well as Jones and

Lefebvre's ABB report.

The fmt systematic study on smelt cooling process was conducted in a bed cooling

study project by ADL. The project was carried out to evaluate the physical and thermal

conditions of the char beds during bed cooldown. The main objectives of the study included

determination of the mechanisms controlling cooling in beds, and the rates of cooling and

solidification that may occur in typical chadsmelt bed configurations.

ADL observed bed cooldowns at various mills, including one that followed a

simulated ESP. Physical and thermal properties of char and srnelt were obtained by

conducting measurements during mil1 visits. and from the literature at boiler manufacturers

and operators. Information relating to emergency shutdown procedures, flow of molten smelt

within the bed, smelt cooldown and solidification was also gathered. Some techniques were

developed for measunng thermal properties of char beds such as thermal conductivity and

specific heat. A thermal analysis of the cmldown process in the bed was carried out for a

range of typicai bed conditions. and the impact of various factors that influence the bed

cooidown rate was exarnined.

A one-dimensional, transient heat-flow model was &veIoped by ADL for the char

bed. Figure 2.4 shows the heat-flow phenomena considered in the model. The primary modcs

of heat transfer considered were conduction, radiation, and smelt solidification. At the bed

surface, combustion might occur locally generating additional heat, and the heat exchange

with the surmundings was assumed to be by radiation and convection at the top surface. The

heat of smelt was also removed by conduction to the boiler hearth.

The thermal mode1 was experimentally verified for bed conditions following an

emergency shutdown. A sensitivity analysis of the mode1 was conducted in order to

detemine which parameter had significant influences on the predicted cwldown. The

parameters examined included bed lieight, thermd and physical propettics of bed, initial and

cooling conditions of bed, and the cooling rate in the presence of smelt pockets within the

char bed.

The main findings and conclusions of ADL study included:

Bailei Harth

Figure 2.4. Omdmensionai heat-Uow mode1 (ilrcbardson and Merriam, 1977).

1) There is very linle that can be done to promote rapid cooling of the bed interior aftet

initiation of shutdown. The most effective way for shortening the cooldown period is

to operate with a nlatively low bed. Depending on the bed conditions the cooldown

time of the bed may require from 1 day to more than 5 days.

2) The thermal model developed gives a good prediction of cooldown time of the bed

under normal conditions-

3) High cooling rates can be achieved when the char bed has low height. high-porosity

and large pore size in the upper part of the bed, and high concentration of solidified

smelt in the lower part of the kd .

4) The cooling and solidification of smelt are slow in the presence of smelt pockets. A

longer cooldown period is required with low-porosity above the pockets and a Iow

solid smelt content below the pockets- However, the presence of smelt pockets in the

bed interior is not easily detectable.

5 ) The bed surface is covered and flattened by slags 5-15 cm (2-6 inches) in thickness

falling from the walls and other upper components of the iùmace a few hours afkr

shutdown, which decreases the bed cooling rate significantly.

The ADL study provided a sinrting point for understanding the mechanism of char

k d cooldown and smelt solidification. Some of its findings have been proven by field

experience. However, the design and operation of recovery boilers have undergone

signirlcant improvement and innovation since then, especiaily during the 1980's. Some

conclusions drawn by ADL such as wai1 firing of black liquor, and effective ways to speed up

the smelt cooldown are evidentiy no longer applicable. Besides. some values of thermal and

physical properties of the char bed and smelt reported by ADL wen obtained h m

caiculations using pure element properties, or from measurements in laboratories but

involving the change of physical properties of the samples after taken out from the boilers.

The validity of these values nmains questionable.

Jones and Lefebvre (1996) conducted a smelt pool cooldown test after a scheduled

recovery boiler shutdown. The test unit was a decanting bottom kraft chemical recovery

boiler rated at 750 tons (1.65 million pounds) of dry solids per day.

The char bed was bumed down after the shutdown. A probe was inserted immediately

into the smelt pool and continuously measured the smelt pool temperature during the

cooldown period. Temperature readings were taken at ten-minute intervals. The plot of

temperature distribution at each thermocouple from their test is shown with solid curves in

Figure 2.5. Jones and Lefebvre compared their data with the result from ADL nsearch. They

concluded that both results were similar showing that the cooldown time for the smelt pool

was between three and four hours.

Jones and Lefebvre developed a one dimensional transient heat uansfer model to

predict the smelt cooldown. The smelt cooldown process was modelled with an upper

boundary condition of convective heat uansfer to air and a lower boundary condition of

conductive heat transfer to the floor tubes. A finite difierence model was consmcted using

fifteen nodes. The mode1 predictions were compared with expecîmental &ta. The

experimental and computational results demonstrated ceasonable agreement, but the

properties of smelt adopted in the mode1 were not inuoduced in the article.

Jones and Lefebvre also used their model to evaluate the effect of various parameters on the

cooldown rate of the smelt pool. They concluded that the most h a t i c effect on cooldown

time was that of the smelt pool depth. For the model they used, an additional 2 cm of molten

smelt increased cooldown time by 30% and an additional 6 cm (increasing the molten layer

by 43%) doubled the cooldown time.

Jones and Lefebvre obtained real smelt cooldown data and did an important work on

modelling the smelt cooling process. However, their test was of a limited case. in a recovery

boiler it is often difficult to find the hottest spot. and the hot spots are sometimes unreachable

by thermocouple probes due to the large cross-section size of the recovery boiler. The

thermocouple readings from their test might not be the same as the temperatun data that

would be obtained from the hottest place. At the early stage of the smelt cooldown, the smelt

top surface keeps a relatively high temperature so that radiation could be more important than

thermal convection, which should be taken into account in their modelling.

TEMPERA W R E V S T l M E 7 Û ! V E N D E P T H

Figure 25. Temperature vs. üme at given deptbs (Jones .ml Lclcbm, lm.

--

A laboratory-scale experimental apparatus has k e n designed and constructed. Field

tests of this apparatus were conducted at a mil1 site. Molten smelt samples were obtained

fiom smelt spouts of an operating recovery boiler and placed in the experimental apparatus.

Thermal data on smelt cooling process were then collected and analyzed.

3.1 Ekperirnentai Appamais

A picture of the smelt cooling apparatus is shown in Figure 3.1. It consisted of a smelt

cooling vessel equipped with thennocouple probes placed in the vessel to monitor smelt

temperature, a water-cooling unit, an air cooling unit, and a data acquisition system.

3.1.1 Smelt codlag vessel

The vessel was made of 6.3 mm (W") thick stainless steel plates, and had an overall

size of 0.4rn x OAm x 1. lm ( 1 6 " ~ l6"~43") (LxWXH). It consisted of two parts and a top lid.

Each part had a height of 0.33m (13"). The lower part of the vessel, used as a lower furnace

to hold molten smelt, had 36 themocouple probes inserted from two of the side walls into the

smelt pool to monitor the temperature variations at different locations during the smelt

cooling process. The other part of the vessel was used as an upper h a c e . There were four

spacers used for adjusting the total height of the vessel to accommodate various smelt

heights. One of the spacea had ten 2 cm-diameter holes on each side, used as air ports. A

number of holes were a h located on the top lid of the vesse1 so that the thennocouple probes

could be inserted to measure the smelt temperature at the top surface. nie smelt cooting

vesse1 was thermally insulated with firebricks and fire blanket on the outside.

3.1.2 Water codiag unit

This unit included two water coils and piping connected to a water supply. Each

cooling coil was cornposed of two 6.3 mm ( W ) OD thick-wall copper tubes concentncally

wound to fom a two-tube coil assembly with an outer size of 0.3m x 0.3m (1Yx12") as

shown in Figure 3.2.

Figure 3.1. S d t amkg expecbntai appontan

Cooling water was passed in opposite directions within these two copper tubes, which

rninirnized the temperature diffennces ihroughout the coil. These two cooling coils were

placed separately, one at the bottom and the other against one side wall of the experimental

vessel (Figure 3.3).

3.13 Air cding unit

This unit was composed of a vacuum pump, with a maximum capacity of 0.01 m3/s

(22 SCFM), connected to a nozzle on the top lid of the smelt cooling vessel. The capacity of

the vacuum pump was adjustable so that the expiments would be nui at different air flow

rates. Cooling air was drawn into the experimental vessel from the air ports just above the

smelt top surface and then discharged through the nozzle to remove the heat of smelt nleased

from the smelt top surface.

3.1.4 Instrumentation and dota acquisition system

This system consisted of more than 40 themocouple probes and a PC-based data

acquisition board. Thermocouple probes, type K. were used to measure the temperatures of

smelt, cooling water, and cooling air duting the cooling process. The software used to collect

experimentai data was Labtech Notebook pro 10.1.1.

Figure 3.2. Watcr cooling coi1 (with diagram showing the tlow directions of coolhg

watcr in the tubes).

Figure 3.3. 3.tions of cooiing COUS piaœd M d e the smelt miing vessel.

3.2 ExperimentalProcedures

Two field tests were conducted at a kraft mil1 with a decanting flmr recovery boiler.

nie lower funiace of the recovery boiler had a Iûm x IOm cross-section size. Then were

totally eight smelt spouts on the side walls of the boiler. The temperature of the molten smelt

flowing out of the furnace was about 780°C. A picture of a smelt spout with molten smelt

running out is show in Figure 34.

Figure 3.4. Smelt Spout with running d t e n smlt.

Molten smelt was collected from the smelt spout, and poured into the experimental

vessel where it was ailowed to cool. Tap water was used as cooiing water running in the

cooling mils to nmove heat from the smelt. 18 thermocouples were inserted 16.5 cm (6.5")

within the experimental vessel (in the mîddie of the vessel) while other 18 thermocouples

inserted 5.2 cm (2.S'). The thermocouple probes were placed in 6 layers into the vessel. The

lowest layer was very close to the bottom of the vessel (1.25 cm above) while other layen

wen located above this layer at a 5 cm (2") distance respectively so that the temperature

variations of the smelt at different locations dunng smelt cooling process could be monitond.

3.3 Fît Field Test

33.1 Procedure

The lower part of the experimental vessel was thermally insulated with firebricks.

Two water-cooling coils wen used, which were placed on the bottom and against one inner

side wall of the vessel. Cooling water was supplied to the water coils at ambient temperature.

Molten smelt was scooped out directly from the operating recovery boiler and filled the

experimental vessel up to a height of about 26.5 cm (10.6") above the bottom. Figure 3.5

shows this sampling process.

In order to avoid the potential darnage to the copper coils caused by the direct contact

with molten smelt, the water-cooling coils were covered with stainless steel plates so that the

coils were isolated from molten srnelt during the smelt cooling process.

Cooling air was drawn into the experimental vessel from air ports above the smelt

surface and then vented from the vessel. The vacuum pump was used to draw air into the

vessel and was operated at its maximum capacity of 0.01 m3/s (22 SCFM.). A brief schematic

of the air cwling operation is shown in Figure 3.6.

Cooled wall

Figure 3.5. Smelt sampüng.

h ~ i r out

Air in

Uncooled waii Themocouples

The temperatun variations of smelt, cooling air, and cooling water during the smelt

cooling process were monitored by thennocouple probes. The probe signais were collected

and amplifed by EXP-32 data expansion boards, and then ncorded by the data acquisition

system at a sampling frequency of 0.2 Hz (one reading for dl thermocouples every five

seconds). The test was stopped after 6 houn of smelt cooling when al1 the temperatun

readings from the thennocouple probes inserted into the smelt were below 550°C.

3.3.2 Results and discussion

The field test was mn for 6 hours. The raw experimental data collected by the data

acquisition system are presented in Appendix B.

The smelt temperature data, obtained by thennocouple probes inserted in the middle

of the experimental apparatus ( 16.5 cm into the smelt cooling vessel) at various depths during

the smelt cooling process, an shown in Figure 3.7.

The cooling curves of the smelt clearly indicate the dependence of smelt cooling rates

on location. The curve labeled 25 cm in the figure indicates the smelt temperature very close

to the bottom of the experimental vessel (25 cm below the smelt top surface), while the other

curves refer to the smelt temperature data at different depths (0.3 cm, 5 cm, 15 cm, and 20

cm) from the smelt top surface.

TlME (MIN)

Figure 3.7. Smlt temperature at various depths.

The temperature profiles at various depths pnsent different characteristics. At the top

surface and very close to the bottom of the smelt pool, the smelt cooled down rapidly.

Temperatuns of smelt in these areas decreased sharply and nached 720°C within a few

minutes, the temperature at which the molten smelt begins to solidify. Obviously the smelt in

these areas remained in a completely molten phase only for a few minutes. Then the smelt

temperatures decreased smoothly and steadily. It can k found that the release of smelt latent

heat from phase change did not have a large effect on the cooling process. Near the top

surface (0.3 cm depth) the smelt temperature decreased much faster than that close to the

bottom. Dunng the test 3 thennocouples close to the smelt top surface prowded from the

smelt surface after a few hours of cwling due to the volume shrinkage of smelt caused by

solidification. As a result, these thennocouples actually measured the air temperatun above

the smelt top surface. A picain taken aftcr mnning the test for 3 hours is show in Figurr 3.8

in which the separation of themwrouple probes From the smelt top surface can be clearly

observed.

Figure 3.8. Top-view of the snelt top surface akr 3 boucs of codiag.

In contrast with the sucface region. the cooling of smelt in the central region was

slow. The smelt temperature at the layer lS cm (6") below the smelt top surface decreased by

only about 20°C and was still above 720°C after about 2 hours of cooling. Since the fnezing

temperature of smelt is typically about 720°C, the persistent high temperature indicates that

the smelt in ihis region was still molten at that moment. This slow cooling is attributed to the

low thermal conductivity of fmen smelt and slow removal of the latent heat of molten smelt

released during solidification. A k r this pend the srnelt temperatun decreased steedily until

it nached below 550°C at the end of the test Recent expenences and research show that the

melting temperature of smelt in recovery boiiers can be expected to range from 54û°C (fmt

melting temperature) to 740°C (complete melting tempenitun). Therefore, the results shown

clearly indicate that in the central region of the experimental vessel the smelt was still

partially molten even after 6 hours of cooling. This is consistent with numerous field

observations which show that because of the slow heat removal hot spots are still pcesent in

the char bed 12 hours after the initiation of an ESP.

The srnelt temperatures at 5 cm (2") and 20 cm (8") depths al1 remained above 720°C

for about 1.5 hom. The temperature profiles of these two curves remained almost identical

during this time pend. This can happen only when the temperatures in these two regions are

well within the melting temperature range (720°C to 730°C as generally refemd). After this

period the smelt temperatures decreased steadily. It can be seen that the smelt temperature ai

20 cm (8") depth (closer to the bottom of the experimntal vessel) had a higher cooling rate

than the smelt at 5 cm (2") depth (closer to the smelt top surface). Since both temperature

sampling points roughly had the same distance to the smelt surfaces (about 5 cm to the

bottom or top surface, respectively), this difference in cwling rate indicates that cwling of

smelt from the bottom by mainly thermal conduction is more effective than from the top

surface by radiation and convection, which is also confimed by field observations from

s hutdown prac tices on recovery boilers.

The smelt within the experimental vessel displayed high temperature gradients in each

layer. For the layea 15 cm, 20 cm and 25 cm (V, 8" and 10") in depths, the distances

between the layers were al1 5 cm (2'3. But the temperature diffennces were about lûû°C and

120°C respectively rifter 6 hours of cooling. The delay of heat removal caused the high

temperature gradient. The heat nlease within the smlt interior via thermal conduction

becomes the bottieneck for smelt cooling pmcess. The large temperature diffennces between

these two very close layea ülustrate the difficulty of using themacouple probes to assess the

overall char bed thermal state in a recovery boiler following an ESP. The confidence range of

a thermocouple temperature reading detennines the number of thermocouple probes q u i n d

to detect char bed condition and the usefulness of this measuîing method for diagnosing the

overali thermal state of aie char bed in a recovery boiler.

100: :

r I I 1 r 1 I I I 1 I

60 120 180 240 300 360 i

O TlME (MIN)

Figure 3.9. Effect of water-moled side w d l at 15 cm (6") below the smelt top surface.

The effect of a water-cooled side wall on the smelt temperature distributions at the

layer 15 cm (6") klow the smelt top surface is shown in Figure 3.9. The temperature

sampling points (thermocouple probe) at this depth were 12.5 cm (5") apart, and located at

the middle, close to the water-cwled side wall and fa- h m the water-cooled side wall. The

pcesence of the water-cooIed side wall had a clear impact on the smelt temperature

distribution. The temperatun of the smelt close to the water-cooled waii decreased much

faster than the temperature f i from the cmkd wall, while in the middle the temperature

decreased at the slowest rate. The high temperature gradients within the smelt can also be

seen clearly in the figure. This is again due to the low thermal conductivity of smelt and the

release of latent heat of molten srnelt during solidification. At the end of the test the

temperature difference between the location close to the water-cooled side wall and far from

the wall was as high as 130°C, which indicates the effect of an enhanced cooling condition at

the boundary on the smelt cooling rate. Smelt temperature distributions at other depths in the

smelt pool can be found in Appendix C.

3.4 Second Field Test

The main reason for conducting the second field test was to siudy and analyze the smelt

cooling process under better controlled heat transfer situations and clearer boundary

conditions.

in the first field test, the smelt was cooled by a bottom watercmling coil as well as a

side water-cooling coil, through the heat transfer at the smelt top surface (radiation and

convection), and also through uncw led side walls. During the smelt cooling process, the

watercooled side wall had a uniform and constant temperature. However, the smelt

temperature was not uniform in the vertical direction within the smelt pool. Therefore, the

flux of heat removal through the water-cooled side waii was not uniform at different

locations. Ail of these heat transfer phenomena made the smelt cwüng a 3dimensional

process with multi-mechanisms of heat transfer involved. The effect of some mechanisms is

very Micult to specifu and evaluate, which complexes the analysis on the experimental data.

There was also a flaw in the design of the experimental apparatus in practicai application. In

the f i t field test, the bottom water-cooling coil was covend with a stainless steel plate to

avoid the direct contact of coil with the mlten srnelt. which left some gaps between the coil

and the plate and therefore increased the heat resistance of cooling from the bottom of the

experimental apparatus.

Thenfore, a second field test was planed to be conducted. The experimental apparatus

was improved. The heat transfer situation was simplified and the uncertainties presented in

the first field test were intended to be eliminated in the second field test.

3.4.1 Procedure

The second field test was focused on measuring cooling rate of smelt when cooled

only at the bottom of the experimental vessel. The experimental vessel was under good

thermal insulation For eliminating heat loss via side walls. The side water-cooling coil and air

cooling unit were not used in this test. The upper part of the experimental vessel was

removed to diminish the ana of stainless steel walls exposed to thermal radiation so as to

d u c e the arnount of heat loss via conduction in stainless steel wails. High conductivity,

castable powder (with higher thermal conductivity than stainless steel) was used to fil1 the

gaps between the bottom water cooling coil and the stainless steel plate covering the bottom

coil so that the heat transfer resistance caused by the gap could be eliminated. The properties

of this castable powder are shown in Appendix D. Molten smelt samples were scooped and

pound into the experimental vessel as show in Figure 3.10. The filling process was finished

within 4 minutes. A topview of the molten smelt pool is show in Figure 3.1 1. Then the top

of the experimental vessel was covend with a stainless steel plate irnmediately- AU side

walls and the top plate of the experimental vesse1 were thermally insulated with fire bricks

and a fire blanket as shown in Figure 3.12. The test was continuously monitond using a data

acquisition system at a sampling fnquency of 0.2 Hz until al1 the thermocouple readings

within the smelt pool approached 250°C.

Figure 3.10. Smlt srmpliag (A boikt operator was pouriag the molten smelt

into the smelt cwbg experime~~tal vessel).

Figure 3.11. Topview of molten smelt pool.

3.4.2 Results and discussion

The smelt temperature during the smelt cooling process is shown in Figure 3.13. The

sampling points were located at the middle of the smelt cooling vessel (16.5 cm into the

vessel) at various depths within dic smelt pool. The numbers showing in the figure represent

the depth of that sampling point below the top surface of smelt pool.

The experimental data present the sarne trends as in the first field test. With better

thermal insuiation and only cooled by the bottom water-cooling coil, the smelt cooling

process was much slowcr than in the first field test. It can be seen from Figure 3.13 that the

smelt was cooled very fast at the beginning stage in which the smelt temperanite everywhere

nached the cange of 720 - 730°C and the smelt began to solidify within 20 minutes of

cooling.

A more detailed graph showing the smelt cooling in the first 30 minutes is illustrated

in Figure 3.14.

Figure 3.13. Smeit tempenature pmfiies at various depths witbia the smlt pool.

. Figure 3.14. Smeit oooling in the finit 30 minutes.

The smelt cooled from 780°C to 720 - 730°C in 15 minutes. This was remarkably

fast. For the layers close to the top surface and bonom, it is not difficult to realize this rapid

cooling. But within the smelt interior, 12 cm (4.5") away €rom the surface, this cooling stage

was also finished at the sarne time scale. It appears that the thermal conductivity of molten

smelt is quite high, which is in conflict with the literatun nsults which showed that the

thermal conductivity of molten smelt is even lowet than that of frozen smelt. In fact, in the

beginning the molten smelt pool was not punly stagnant. Because of large temperature

ciifferences between the smelt and boundaries, there could have been natural convection

within the smelt pool, and this could transport heat from the smelt interior to the boundaries

and thus intensified the heat nmoval rate. Therefore, the heat transfer in smelt at this stage

was likely controlled by thermal conduction and convection in the molten smelt pool. An

effective thermal conductivity but not a thermal conductivity of "pure" molten smeit should

be used for calculating the heat transfer rate of smelt when in molten state. A relevant

discussion about effective thermal conductivity can be seen in Section 2.2. The measurernent

and determination of this effective thermal conductivity for molten smelt needs more

specifically designed equipment and is therefon beyond the scop of this study.

Following the initial steady decline, the smelt temperature stayed above 700°C and

the cooling curves decreased slowly in a very flat dope for about 2 hours, the similar time

interval as in the first field test. The latent heat of smelt is only a very mal1 part of the heat

load required to be removed and should have a very lirnited influence on the smelt cooling

process. However, since the release of latent heat takes place within a very n m w

temperature range and the removai of this heat is slow due to low thermal conductivity of

smelt, the smelt would remain at this temperatun level until this latent heat nleased h m

solidification is conducted to the ambient atmosphere.

After this stage the smelt temperature decreased steadily. The temperature reached

550°C everywhere after 10 hours of cooling, and approached 250°C after 10 hours of cooling.

At the layer very close to the smelt top surface (1.3 cm below), the smelt temperatun

is highly affected by heat transfer at the surface and so the surface region was cwled much

faster than the smelt intenor in the fint few hours. Because the top of the expenmental

apparatus was sealed in the test, there was only radiation from the smelt surface to remove

heat from the smelt. This fast cooling near the top surface shows that radiation heat transfer

had some influence on the cooling rate of smelt very close to the top surface. However, this

influence was reduced with the decrease in smelt tempetatun and with the incnase of

distance below the smelt top surface.

The limits of this influence caused by radiation on the cooling rate can also be seen

from another phenornenon. During the cooling process, the hottest core of the smelt was not

in the rniddle region of the smelt pool, but moved up towards some location close to the layer

6.3 cm (2.5") below the top surface. In the test, heat was mainly removed €rom smelt by both

radiation on the top surface and water cooling at the bottom. Therefon, this upward

displacement of the hot core indicates that cooling from bottom was dominant in the second

test.

3.5 Meunving Thetmal Conductivity of Fmzen Smelt

nie thermal conductivity of frozen smelt is probably the most important factor

influencing the smelt cooling rate. The value cumntly used in industry was recommended by

ADL in the 1970's. However, the determination of this value is questionable as discussed in

Chapter 2. A more diable value is needed for more accurate pndiction of the smelt cooling

process.

3.5.1 Nature of thermal conductivity

A material at a given temperature contains energetic free electrons if this material is

metallic or semi-conducting, plus a concentration of lattice phonons (Love11 et al., 1976;

Grimvdl, 1986). These electrons and phonons move in random directions and hence

transport euergy. But in equilibrium thece is no flow in any particular direction. This energy

is eventually lost by interactions between phonons and electrons, phonons and phonons, or

electrons and electrons. Therefore, the equilibrium is a dynamic situation with the excitation

and deexitation of electrons as well as the citation and destruction of phonons.

If a temperature gradient is imposed on the material, the hot and cold areas will have

different phonon concentrations and different mean electron energies. The phonons and

electrons will move towards the lower concentration and energy ana by a natural diffusion

process which results in a net energy transfer. If the temperature gradient is continuously

imposed upon the material, then the energy must be transferred to the lower temperature end.

Heat is thus being continuously conducted.

The ability of a material to conduct heat is detemined by its thermal conductivity, k,

which is defined by Fourier's law of heat conduction:

q = - k ( r n

Equation 3.1 States that the heat flux vector q is proportional to the temperanice

gradient vector, PT. The positive constant k is called the themal conductivity. The aegative

sign is due to the fact that heat flows down the temperature gradient (Bird et al., 1966).

35.2 Principle of measurement method

There are severai ways to rneasure thermal conductivity of solid matcrials. ASTM has

developed many standard methods for the rneasurement in specific materiais. The meihod of

steady-state linear heat flow at high temperature is the most common approach (Laubitz,

L969; Finck, 1937).

When a constant heat source is applied on the surface of a material and if the heat can

be removed only by heat conduction from one direction, the heat transfer process can be

sirnplified to a one-dimensional steady-state thermal conduction problem. Equation 3.1 c m

be simplified to:

Q/A = k AT& [3 4

where Q is the heat generation rate at the surface [Wl, A is the cross-sectional area of heat

fiow [mt], AT is the temperature ciifference between any two measuring points, and Ax is the

distance in heat flow direction between these two measuring points.

3.53 Procedure

The smelt was frozen af'ter the second test. A heating plate (45V, 45W) was used as a

constant heat source to heat the smelt at the top surface. A stainiess steel plate was placed

below the heating plate to achieve a unifom temperature distribution on the smelt top

surfafe. Fiiblanket was used to cover the top of the heating plate and to seai the gaps

between the stainless steel plate and the experimental vessel side walls, to eliminate heat loss.

A steady temperature profile was gradually established within the smelt sample by cooling

the smelt with the bottom water-cooling coil. Smelt temperature gradients were monitored by

the thennocouple probes remaining in the frozen smelt and then recorded by a data

acquisition system.

I Fire Blanket Heating Plate SS Plate Frozen Smelt

Thermal lnsulation

Themiocouples

Cooling Coil f

Figure 3.15. Measurement of thermal conductivity of fmzen smlt.

3.5.4 Results and discussion

Smelt temperature curves dunng the expriment are shown in Figure 3.15. The

sampling points are in the rniddle of the experimental vessel (O. 15m into the smelt pool). It is

obvious that the smelt temperatures at diffcnnt layers were gradually approaching constant

value as the heat transfer within the smelt reached steady-state at the end of the experiment.

A more detaiied graph illustrating the smelt temperature history in the last one hour of the

experiment is shown in Figure 3.16. Aimost no change in the smelt temperatures cm be seen

in this figure. At the location closer to the side boundary (0.075111 into the smelt pool), the

smelt temperature curves presented in Figure 3.17 showed the same tendency as in Figure

3.14. Heat transfer in the smelt ai this location was also approaching steady-state.

The average temperature data in the last one hour of the expriment at the locations

0.15m and 0.075m into the smelt pool are tabulated in Table 3.1.

Tabk 3.1. Average smelt temperature in the hast one hour of the urperinient.

I O.ISm into the smelt 0.07Sm into the I 1 Temperature difference

Deph within I pool (middle of the

the smelt pool

1.3 cm (0.5")

6.3 cm (2.5")

11.3cm(4.5")

16.3 cm (6.5")

21.3 cm (8.5")

26.3 cm (10.5")

smelt pool (cioser to bctween different locations

smelt pool)

Tp (OC)

239.5

135.3

116.1

IO 1.2

88.1

78.2

in the same Iayer

AT ( O C )

55.0

6.8

6.2

6.0

3 -4

2.8

ATp ( O C )

- 104.2

19.2

14.9

13.0

9.9

boundary)

Tb (OC)

1 84.5

128.5

109.9

95.2

84.5

75 .4

ATb (OC)

-

56.0

18.6

14.7

10.7

9.1

TlME (MIN) Figure 3.16. Smelt temperatures in the middk of the eirperimatd v d

(O.LSm into the smelt pool) approaching constant values.

- 6.3cm 1 1 . 3 ~ m 16.3cm 21.3cm "

v 1 i - 1 1

L I - 1

1 9 - 26.3 cm . 1

1 1 i I I I I 1 i

O 80 120 180 240 300 360 420 480 540 600 660

TlME (MIN)

Figure 3.17. Smdt temperatures at the location 0.WSm into the smelt pad

a p p c 0 1 l ~ g constant values.

It cm be seen €tom the table that the smelt tempera- ciifferences between adjacent

layers, ATp (or ATb), were not q u a i aithough the distances between these two adjacent laycrs

wen al1 0.05m (2"). For the smelt at the location OMm into the smelt pool. the average

temperature dnerence, AT,, decnased from 1W0C between the layers of 1.3 cm and 6.3 cm

(0.5" and 2.5") below the smelt top surface to lg°C between the layers of 6.3 cm and 1 1.3 cm

(2.5" and 4.59, and finally to 10°C between the layen of 21.3 cm and 26.3 cm (8.5" and

10.5'). ATp decreased in the direction towuds the bottom of the smelt. For the smelt at

location 0.075m into the smelt pool, the temperatun difference, ATb. s howed the same

tendency.

The difference in AT was probably caused by heat conduction dong the side

directions. During the expriment, the heat source was placed on the top of the smelt, and

heat released from the heat source was conducted in the smelt towards the bottom. but also

towards the side boundaries. This heat loss to the sides reduced the amount of heat reaching

the bottom of the smelt. This heat transfer process can not thus be simply regarded as a one-

dimensional problem.

It should be noticed that the temperature diffennce. AT, (or ATb). was lû4 OC (or 56

OC) between the top two layers. This value was much larger than the temperature difference

between other layers. It appeared that the thermal conductivity of smelt at this region was

quite iow. Since the distance between each layea was only 5 cm (2"). it was unlikely that the

high temperature difference was due to the heat loss to side boundaries. Because the heat loss

was also via conduction, if the thermal conductivity of smelt was low, the rate of heat loss by

thermal conduction shouid also be low. One possible nason for the p ~ e n c e of this high

temperature difference in the region close to top SUfface was due to the effect of pomsity of

smelt. After fiozen, the smelt at the location close to top surface is always more porous than

the smelt interior. When the smelt temperature is relatively low, the porous s t~cture of smelt

may hinder the heat transfer into the smelt and result in a low effective thermal conductivity

at this region, Another possibility was that the thermal conductivity of smelt is dependent on

temperature, and if so the dependence of the themal conductivity on temperature would not

be linear but should be complicated. However, a good explanation for the presence of a large

temperature difference near the top region can not be pmvided at this moment.

The magnitude of this heat loss to the side walls can be qualitatively evaluated from

the smelt temperature difference at the sarne depth but different horizontal locations within

the smelt pool. AT in Table 3.1 shows the temperature differences of smelt at the location

0.15m and 0.075m into the smelt pool. At the layer close to the top surface, since the smelt

temperature (240°C) was much higher han the temperature close to the bottom (78'C), the

heat loss was much pater. For the layers of 1.3 cm (OS"), 6.3 cm (2.5") and 1 1.3 cm (4.5").

AT was 55 O C . 7OC and 6OC. respcctively. This means that the horizontal temperature gradient

was relatively high and heat loss was therefore significant. However, at the layer of 26.3 cm

(10.SW), only 1.3 cm (W) above the bottom, AT was only 3OC, which indicates that the

arnount of heat loss was reduced quickly in the direction from the top surface to the bottom of

the smelt and the heat loss was no longer important at this layer.

It is not straightforward to calculate the value of thermal conductivity for frozen smelt

using Equation 3.2 because the heat flow was strictly one-dimensionai. However, the= is still

a way to make an approximate calcuiation.

During the expriment, the smelt was mainly cooled by the cooling water especially at

the Iayers close to the bottom. The temperature rise of the cooting water was known and the

amount of heat nmoved by the cooling water could be easily calculated. Since the layers of

21.3 cm (8.5") and 26.3 cm (10.5") of the srnelt were very close to the water-cooled bottom

and the horizontal heat Ioss was very limited at these layers, it is nasonable to assume that

the amount of heat reaching these two layen was equal to the heat removed by the cooling

water. Therefore, an approximate thermal conductivity of frozen smelt can be obtained.

The relevant properties and thermal situation of the cooling water during the

experiment can be found in Table 3.2.

The rate of heat nrnoval by cooling water can be determined by equation

Q = f i v CpwATw, P-31

where p, is the density of water, v is the volume flow rate of water, C,, is the heat capacity

of water, and AT,,, is the temperature difference between the water outiet and inlet.

Combining with Equation 3.2,

kfAATp/h = f i v CwATw. f 3 -41

After manipulation of Equation 3.4,

kt = (fi v CPw& /A). (ATW/ATP). W I

Using the data from Tables 3.1 and 3.2, and Ax = O.OSm (Y) and A = 0.09 m2, the thermal

conductivity of frozen smelt, kf = 0.60 W/m°C, is obtained. It is important to notice that this

kr value calculated from Equation 3.5 is 3 1% less than the 0.88 W/m°C value recommended

by ADL.

Table 3 3 Roperties and temperature rise of codiag water during the experhwnt,

Pmperties and t h e d situation

~ensity, hv

Unit

Volume flow me, v

hour of the expriment, ATw

Vaiue

W m 3

f Heat capacity, C,,

Average temperature rise dunng the last one

1 , r n

rn3/s 1.917~10'~

Jkg°C

OC

4,186

O. 133

Mathematical modelling based on CFD techniques and CFD codes is a well

established method for simulating and analyzing heat trsnsfer problems of materials, and

predicting temperatun histones. In this study a generai-purpose CFD code, PHOENICS, was

used as a framework for modelling of the smelt cooling process in the Wrst field test.

4.1 Description of Slmdation M d e l

The smelt cooling process within the experimental apparatus was modelled as a

transient 3-dimenssional heat transfer problem, using a computational fluid dynamics code,

PHOENICS. This mode1 solved a transient thermal conduction equation within the smelt

pool, subject to radiation at the top boundary, heat removai by forced convection to cooling

wnter at the bottom and side water-cooling coils, and convection from other thennally

insulated walls to the ambient atmosphere. A transient solution gave the smelt temperature

distribution in the smelt pool throughout the smelt cmling process.

4.2 Geometry of the Computationnl Donuin

The computational domain of this heat transfer mode1 is shown in Figure 4.1. At the

begi~ing of the cooling process, the entire smelt pool was molten and considered to be

stagnant. The location of the moltenlfrozen srnelt interface was treated as an unknown

parameter and was calculated by the code. The pool geometry was defined as a rectanplar

pool with a flat top surface. The smelt pool was sumunded by steel walls except for the

smelt top surface, which was open to the ambient atmosphere. The side walls were 3.2 mm

('/B") in thickness whiie the bottom wall was 6.4 mm (W') thick. The bottom wall and one

side wall were water-cooled. Other uncooled walls were thermally insulated by finbricks. A

body fitted coordinate system was employed to implement the geometry in the numerical

model.

Convection from water Radiation from top surface cooled side wall (front) /

Convection from uncooled walls

Convection from water-cooled bottom

LX 0-45 m D

Figure 4.1. Schematic of the computationP1 domnia.

The computational domain was specified by choosing appropriate grids in X, Y, and

Z directions. PHOENICS requires that if any heat transfer occurs in an object and l i s

process needs to be taken into account in the numericai cdculation. at least two Mds must be

assigned to this object (Kundsen. 1998). During the smelt cmling process heat conduction

took place within the stainless steel walls and finbricks, which should be calcdated in the

numecical simulation. Seventeen grids were specified in X direztion, out of which two grids

were assigned to the left wall, eleven @ds to smelt pool, two grids to right waii, and the

remaining two grids to a fmbrick. Twentysne grids were specified in Y direction, out of

which fou. grids were assigned to the two walls. thirteen gri& to the smelt pool, and four

grids to the two firebrick layers which were placed on both sides. The smelt temperature

distributions at different depth in the smelt pool are of interest to this study. Therefore, thcm

were thirty gr& specified in Z direction. out of which two grids were assigned to the bottom

wall, and twenty eight gr& to smelt pool. Grid distributions of the smelt pool in al1

directions were not even. Since higher temperature gradients wen present near the smelt pool

boundaries during smelt cooling, particularly near the water-cooled bottom and side walls,

grids were assigned to have finer spacing (doser) near the smelt pool boundaries so that the

smelt temperature variations in these areas could be calculated more precisely.

There were 36 temperature proôes monitoring points within the domain at the sarne

locations as the thennocouple probes used in the field test so that the smelt temperature data

at these sarnpling points from the expriment and numerical simulation cm be compared.

There were also other temperature probes assigned within the bottom and side walls of the

domain to monitor temperature variations of these walls during the simulated smelt cooling

process.

4.3 Heat Tramfer Equatiom

Numerical simulation of smelt cooling process is a complicated task because of the 3-

D geometry and a combination of heat transfer phenornena Heat conduction within the smelt

bed as well as through stainless steel walls and thermal insulation layers, convection through

the boundaries, and radiation from the smelt top surface are al1 associateci with the smelt

cooling process. Heat conduction took place within the smelt bed, stainless steel walls, and

themal insulation layers. The fundamental equation is Fourier's law of heat conduction as

described by Equation 3.1

q = -k (m. 13.11

The energy equation after combining Equations 2.1 and 3.1 becomes:

p C, aT/& = (VokVT) [4- 1 1

If the themai conductivity is independent of the temperature or position and for a

Cartesian coordinate system, Equation 4.1 kcomes (Bird et. al., 1966; Bennett and Myers,

1962)

pcpa/lat = k [aWax2 + a'r/ay2 + a%az2]. [4.4

This is the transient, 3-D heat conduction equation solved in the pnsent numencal

simulation. Slab by slab and whole-field methods of solution of the PHOEMCS code were

applied.

Radiant energy was mainly emitted from the smelt pool top surface. The net radiation

heat flux can be calculated by Stefan-Boltzmann equation,

q = & a (Tm4 -TA [4-31

where q is the net heat flux emitted from the surface, E is the emissivity of the smelt surface,

a is Stefan-Boltzmann constant which is equal to 5.67 x 10" ~ l r n ' l ~ ~ , T, is the absolute

temperature of the smelt top surface, and Ta is the absolute temperature of the ambient

atmosp here.

Heat was also transfened by convection through boundaries of the computational

domain. The principal equation for convection is descnbed by,

Q=hAAT,

where Q is the heat flow into the air [Wl, h is heat-transfer coefficient [w/rn2~], A is the

characteristic area [m21, and AT is a characteristic temperature difference [KI.

At the boundaries, the transient heat conduction problem is coupled with the

convection. The heat removed by the ambient atmosphen by convection is equal to the heat

reaching the boundary by conduction.

4.4 Deternination of Bounàary Conditions

Heat transfer processes at boundaries involved forced convection of cooling water in

the water coil, forced convection of cooling air and radiation on the smelt top surface, and

convection to air at other thermal insulation layer surfaces. The radiation at the smelt top

surface was alnady discussed in Chapter 3, and therefore no detail is given in this section.

4.4.1 Forced air convection on a flat plate

When a fluid flows over a flat plate as illustrated in Figure 4.2, beginning at the

leading edge of the plate, a region develops where the influence of viscous force is felt This

region of flow is called the boundary layer.

uiitially, the boundary-layer development is laminar, but at some critical distance

from the ledge edge, depending on the flow field and fluid properties, small disturbances in

the flow begin to become arnplified, and a transition process takes place untii the flow

becomes turbulent. A dimensionless number, Reynolds number, is used to determine how

turbulent the flow is. The transition from laminar to turbulent flow occurs when

~e ,=pur /p>5 x los, [451

where Re, is the local Reynolds number, p is the density of the fluid, u is free-stream

velocity, x is the distance h m leading edge, and p is dynamic viscosity of the fluid.

Figure 4.2. Sketch showing difterent boundary-layer fiow regions on 8 h t plate

(Holman,l991).

In the field test, the temperature of the ambient air was measured, thenfore, the

density and dynamic viscosity of air could be obtained. The velocity of air was not uniform

over the exposed vesse1 walls, but cm be assumed to be less than 5 mls. When air at room

temperature (25 - 3S°C) flows over the entin length of the thermal insulation layer, Say 0.4

m. the Reynolds number can be calculated as:

Re=(O.4x5 x 1.177) / 1.85 x lo% 1.27~ lo4

Since the Reynolds number is Iess than 5 x ld. the boundary layer is stiii laminar.

Theoreticaiiy, the local heat transfer coefficient, h, can be calculated fiom the comlation

NU, = 0.332 Pr1" ~ e : ~ , w-61

where Nu, = hx /k is the Nusselt number and Pr = C, p /k is the frandtl number.

For the entire length of the plate (L), the average heat-transfer coefficient can be

obtained fiom the following equation:

IR 1B NUL = h U k = 0.664 ReL Pr 14.11

In practice, air flowing over a flat plate with a laminar boundary layer always has a

relatively low value of heat-transfer coefficient. Table 4.1 gives the approximate values and

ranges of convection heat-transfer coefficient. The data cited in Table 4.1 can be used as a

reference to determine the range of the heat-transfer data to be imported to the numerical

simulation.

Both radiation and convection took place on the smelt top surface. At the primary

stage of smelt cooling process, the top surface was relatively hot. When the smelt

temperature on the top surface is ûûû°C, the arnount of heat removed via radiation is 30,860

w/m2 (assurning emissivity is 0.95 and the arnbient ternperature is 2S°C). From Table 4.1.

the value of the heat-transfer coefficient on the smelt top surface during the cwling process

is about 12 W I ~ ~ ' ' C , therefon the amount of heat removed via convection is 6,900 w/m2.

The ratio of heat removal from the smelt top surface by convation to the removal by

radiation is only 22%. When the smelt temperatun on the top surface is 500°C, this ratio is

30%. Obviously, the high temperature favors radiation so that the heat transfer rate on the top

surface was dominated by radiation. At the surface, because the heat nmoved by convection

was equal to the heat reaching the surface by conduction, and the thermal conductivity of

water is much higher than that of air (more than 20 times), the smelt was dominantly cooled

by forced convection to cooling water at the bonom of the smelt cooling apparatus, and the

effect of convection on the smelt top surface was not important.

Table 4,1,

Appmximate vaiues of convection heat-trader caffldents (Hoiman, 1997).

Air u 2 atm Uowing in 2.54- tub i t 10 m/s

4.42 Forced convection to water in coolfng tubes

For flow in a tube, the Reynolds number is again used as a critecion for Iaminar and

nubulent flow. For

Rea = p u - d I p 230, [4-81

the flow is usuaily distinguished to be turbulent. Here, u,, is the mean velocity in the tube

and d is the inner tube diameter,

The mean velocity of cooling water can be cdculaied from the total heat load. cwling

time, and the properties of water as follows. Consider the smelt pool in the expenmental

apparatus had a size of 0.33 x 0.33 x 0.28m ( 1 3 " ~ 1 3 " ~ 11"). the smelt was cwled from

7 m to 550°C, water temperature was 2S°C at the inlet and 3S°C at the outlet, and the

themal property data of smelt are as given in Table 4.3. The heat capacity of the smelt (C,)

is 1,420 J/kg°C, the density (p) is 2000 kglm3, and the latent heat (k) is 142,000 Jkg. With

the above parameter values, the total heat load can be calculated.

Q = M (CdT + k) = (200)(0.33 x 0.33 x 0.28)[1420 X (740 - 550) + 1420001

= 2.47 x 10' [JI

The mean velocity of cooling water and hrthermore the Reynolds numkr can be

obtained from,

Q = pwu-~*cpw (Tout - Ti3&

where At is the cooldown duration time, Tout and Ti,, are the water temperature at outlet and

inlet, with the density of water (pu) is 995 ICg/rn3, viscosity of water (p) is 7.65 x 104 kg/ms,

and heat capacity (C,,,,) is 4,174 Jlkg. The results an tabulated in Table 4.2.

Since the Reynolds number is higher than 2300, the water flow in the tube was

turbulent. A cornrnon expression for calculation of heat transfer in turbulent f b w in smooth

tubes is

Nud = 0.023 ~ e ~ ~ . ~ w (0.6 c Pr c 10) .

where the expoaent a has the following values:

n = 0.4 for cooüng of the fluid and n = 0.3 for heating of the fluid.

Table 4.2. Reynolds number in the water tube mder dieteront cooiing conditions.

4 5 Consideration on Input Smelt T h e d Roperties

Phase change took place during the smelt cooling process. The smelt was completely

molten at the beginning, and then the molten smelt cooled and froze gradually until all the

smelt within the smelt pool became solidified at the end of cwling pmcess.

According to industrial experience, there may be a slight diffennce in the density and

specific heat of the smelt when in different phases, and thermal conductivity could k quite

diffennt in each phase. In this study. constant density and specific heat of smelt for molten

and frozen phases were specified in the numerical calculation, and diffennt thermal

conductivity in each phase was examined. However, the best prediction of the experimental

data was achieved when the sarne themal conductivity for both phases was assumed.

Some input data used in the numerical simulation are tabulated in Table 4.3. An input

data file for the numerical calculation is shown in Appendix E.

Cooldown duntion (hour) Mean velodty (mfs) Reynolds number I

The numencal simulation was peîformed using the heat transfer mode1 just described.

The whole cooling process lasted for 6 hours. A calculation time step was determined by

decreasing the step size until there was no change in the result, which was one minute. The

effect of latent heat was included in the calculation by changing the heat capacity of the smelt

when the smelt temperature reached a range very close to the pceset melting temperature

(730°C). This range was from O.S°C above to OS°C below the melting temperature.

Table 4.3. Rincipai input dota in numericai simuhtion.

M d n g

Density of smelt

Specific heat of smelt

Thermal conductivity of molten smelt

Thermal conductivity of frozen smelt

Initial temperature of smelt

Smelt complete melting point

Smelt latent heat

Emissivity of smelt

Heat-transfer coefficient at uncoded walls

Heat-transfer coefficient at water-cooled walls

Ambient temperatun at smelt top surface

Ambient air temperature at other sides

Value

2 , r n

1,420

0.60

0.60

740

730

Unit

W m J

Jkg°C

W/m°C

W/m°C

OC

O C

142,000

0.95

6

13

37

25

Jkg

wlm2a0c

~ l r n ~ a ~ c

O C

O C

The experimental data were used to adjust the computational model parameters. The

model was modified by varying the input data until the best match with the experimentai data

was obtained, mainly by adjusting thermal conductivities of the molten and fiozen smelt, and

the heat-transfer coefficients at the water-cooled and uncooled walls. The simulation results

showed the best prediction is obtained when the value of thermal conductivity of smelt was

assumed to be equal to 0.60 W/m°C for both the molten and solid phases.

Cornparisons between the experimental and pndicted results at various depths of the

smelt pool are shown in Figure 4.3. The temperature probes were positioned in the middle of

the smelt pool.

- Experimental - Simulation

O 60 1 20 180 240 300 360 TlME (MIN)

Figure 4.3. Comparison of prrdicted and masuml temperatures

New the bottom wall [25 cm (lû") below the smelt top surfaca both experirnental

and predicted results exhibited very good agreement. The only obvious difference appeared

at the beginning stage of the cooling process, with the predicted smelt tempetanire decreasing

slightly €aster. But the difference was Iess than 30°C at any time.

At 5 cm (2") below the smelt top surface, the predicted temperature remained at the

initial temperature of 74û°C for almost 20 minutes, then it speeded up to decrease. This was

very different from the exprimental nsult which showed that the smelt at this depth began

to solidify as soon as the experiment started and becarne completely frozen within a few

minutes. Afier cooling for about 1 hour, the predicted temperature nached the same

temperature as the experimental data Then the predicted temperature remained at a lower

temperature level until 5.4 hours after cooldown began. After that, because of the sudden

acceleration of cooling in the expriment, the predicted temperature gradually became higher

than the experimental data. At the end of the cooling process the temperature diffennce was

about 50°C. Since the smelt at this depth was close to the smelt top surface, the predicted

temperature may be influenced by radiation as well as by convection on the top surface.

At 15 cm (6") below the smelt top surface (in the central ngion of the smelt pool), a

good prediction of the smett cooling curve was achieved especially for the fmt 5 hours. At

the end of cooling the temperature difference was less than 4û°C. Because the heat trapped in

this central region could escape only via thermal conduction of the smelt, the themal

conductivity of smelt played a major role in the heat transfer rate. Thenfore, the agreement is

an evidence that the value of the thermal conductivity used in ihis simulation was close to the

real value.

Cornparisons of the temperature curves at the depth 15 cm (6") below the top surface

of the smelt pool is shown in Figure 4-4, in which the cooled wal1. rniddle, and uncwled wall

indicate the locations 1.3 cm. 13.8 cm and 26.3 cm (OS", 5.5". and 10.5") from the water-

cooled wall respective1 y.

The reasonably gooâ agreement between the experimental and predicted data can be

seen in this figure. It is noticeable that the predicted curve repiicated the experimental curve

very well at the location very close to the water-cooled wall. Since the heat removal rate near

the water-cooled wall was sensitive to the forced convection of water, the agreement

obtained in this ngion suggests a proper boundary condition in the numerical simulation.

800 I

1 - Simulation

O 60 1 20 1 80 240 300 360

TIME (MIN)

&tances fmm the water-cooled side wall.

The pcedicted temperature distribution in the smelt pool after 6 hours of cooling is

illustrated in Figure 45, in which the numbers shown in larger font npresent the predicted

temperature data whüe the numbers in srnalier ones reprcsent the experimental data. A

relatively hot core can be observecl clearly in the central region in both the experimental and

numericai results. It cm be seen that the temperatun distribution was not symmetric when

comparing the locations close to the water-cooled wall the opposite uncooled wall. The effect

of the si& water-cwling can be seen clearly. The isothermal sphens rnoved slightly toward

the water-cooled wall. The closer to the water-cooled w u , the higher the temperature

gradients are in this figure. Other figures showing the temperature distributions at different

depths after 3 hours of cooling cm be seen in Appendix F.

TOP SURFACE

BOlTOM (COOLED)

Predicted temperature dhtdbutbn dtcr 6 hours of coolhg (numbers

shown in Ltpr font repmmting the pcedjcted temperature &ta, whae the numbers in

Srrmller ones rrpeesenting the experimentai data in the ûrst field test).

4,7 Parametric Studies

The nurnericai simulation mode1 was hirther used to evaluate the impact of some

major parameters on the smelt cooling process. The parameters examined include emissivity,

heat capacity. latent heat, thermal conductivity of smelt, and heat-transfer coefficient.

4.7J Vaiidation of numerical calculation

In numerical calculation p.C, was used together as a product. Different values of

density p and heat capacity C, but the same value of pC, should not change the result of

cakulation. This is a convenient way to examine the validity of the CFD code used in the

numerical calculation. Figure G.1 shows this promising result, in which the density was

decreased by 50% but heat capacity was increased by the same amount to kecp p C ,

constant. The calculation results were identical.

4.72 Emissivity at the smelt top surface

The effect of emissivity of the smelt surface on the cooling time was exarnined by

modifying the value of the emissivity. Figure 0.2 shows the cornparison of the predicted

temperature curves at different ernissivities. When the emissivi ty was decreased b y aimost

5096, the results showed that at 5 cm (2'') below the top surface of the smelt pool, the cooling

rate was slower. However, this temperature diffennce was nlatively small, at only 20°C afkr

6 hours of cooling. At the layer 15 cm (6") below the top surface there was almost no

temperature ciiffierence during the entire cooling process. It clearly indicates that emissivity

can ody affect the cooling rate of the srnelt near the top surface but not the interior of the

smelt pool even a few inches inside. In other words, radiation at the top surface has a minor

impact on the smelt pool cooling pmcess.

4 . 7 Heat capacity of smelt

Heat nleased in a smelt cooling process is mainly sensible heat of the smelt.

Therefore, the heat capacity has a significant impact on the smelt temperature profile. This

phenomenoa can be seen in Figure G.3. When the heat capacity was increased by 50%. very

large temperature differences were obtained everywhere in the smelt pool. After 6 houn of

cooling this diffennce was as high as LO0C at al1 depths. At the layer 15 cm (6") below the

top surface the smelt temperature remainecl above 730°C for more than 3 houa. By

extrapolation of the smelt cooling curves, the increase in heat capacity of smelt resulted in

extending the smelt cooling process by about 3 hours (more than 50% time). A high heat

capacity can dramatically increase the smelt cooling time.

4.7.4 Latent heat of smelt

The value of latent heat adopted in this study is from ADL report (Richardson and

Memam, 1977). In recent years some higher values were aiso suggested in Iiterature, but the

ADL value is still commonly accepted in industry.

The effect of latent heat of smelt was assessed by comparing the predictions with and

without the contriiiution of latent. When the value of latent heat was specified to be zero,

limited temperature changes were exhibited at every depth as shown in Figure (3.4. The

temperature ciifference was about 30°C after 6 hours of cooling. Thus, the effect of latent heat

is not very important to the cooüng process. As discussed eariiec? latent heat talces only a

small part of the heat load required to Ise removed €rom smelt. which is consistent with the

numerical calculation.

4.7.5 Impact of t b e d coductivity of smlt

Thermal conductivity has a substantial effect on the smelt cooling process. The

currently used values for diffennt phases were recommended by ADL, 0.88 Wlme°C for

rnolten smelt and M5 W/mm°C for frozen smelt. in this study, these values were originally

used as input in the numerical simulation. Then the predicted results were compared with

experimentai data The input values of thermal conductivities were then moâified until the

best match with the experimental data was achieved. It was found that for the best match a

thermal conductivity of 0.60 W/rnm°C should be selected for both phases, and the ~ s u l t

calculated with this thermal conductivity was used as a standard case for parametric studies.

The numerical nsults using the ADL values are compared with the standard case as

shown in Figure 4.6.

From the top surface to 15 cm (6") below. the smelt temperature predicted with

various thermal conductivities was always lower than that of the standard case. After 6 hours

of cooling the largest temperature difference was about 6û"C. Since the smelt temperature

was under the preset melting temperature of 730°C most of the tirne during the cooling

process, which means that k = 0.88 W/me°C was mainly used in the numerical calculation,

the nsults confirmed that a higher smelt cooling rate is obtained when with a higher thermal

conductivity. On the other hand, near the bottom of the smelt pool the temperature difference

was negligîble. Since this layer was very close to the water-cwled bottom wall. heat that

reached the bottom was nmoved quickiy. Heat transfer rate at the bottom was thus controiled

by the rate of heat conduction which removes heat from the smelt pool interior to the bottom.

iI - Km=0.45 Wlm a°C Kf=0.88

O 60 120 180 240 300 360

TlME (MIN)

Figure 4.6. Cornparison of various thermsl conductivities for eacb phase with the

standard case.

The impact of thermal conductivity on the smelt cooling pmcess was further

investigated using k a . 4 W/mm°C and k = 1 .O WIma°C to compare with the standard case.

The cornparisons are shown in Figures 0 .5 and G.6. The large temperature dierences with

the standard case indicate that the highet and lower values of the thermal conductivity used

are both far from the actual value.

4.7.6 of boiiaduy conâitions

Ultimaîely, heat is cemoved from the smelt pool by heat tmsfer to the sunoundings

ai the boundaries, which include tht smelt top surface, watercooled bottom and side wails,

and other three uncooled walls. In an earlier section, the effect of smelt emissivity at the top

surface has been investigated, and it tums out that the heat transfer from the top surface could

not appanntly affect the smelt cooling process. In this section, the effect of boundary

conditions on the smelt cooling procas outside the water-cooled and uncooled walls is

evduated.

The effect of varying the heat-transfer coefficient for the water-cooled walls is shown

in Figure 4.7, where hi means the heat-transfer coeficient at the water-cooled walls while h2

indicates the heat-transfer coefficient at the uncooled walls. When the magnitude of hi was

increased from 13 W / ~ ~ O C by IO%, heat transfer at the water-cooled walls was intensified

as shown in Figure 4.7a. The smelt was cooled down faster. The modification of hi resulted

in the change of smelt temperature distribution at al1 depths, but the effect decreased in the

direction towards the top surface of the smelt pool. Mer 6 hours of cooling, the temperature

difierence at 25 cm (10') below the top surface (close to the bottom) was as large as 90°C.

However, at the layers 15 cm (6") and 5 cm (2") below the smelt top surface the temperature

difference was smaller at about 40°C and 20°C. At the top surface, heat transfer was

dominated by radiation, and so the effect of changing hi was negligible.

The same trend was obtained at the depth 15 cm (6") below the top surface as shown

in Figure 4.n. The smelt was cooled down much faster at the location close to the water-

cooled wdl. After 6 hours of cooling the temperature ciifference was more than 80°C, but at

the locations in the middle of the smelt pool and far frorn the water-cooled side wall, the

temperature dflerence was less than 40°C.

O 60 1 20 180 240 300 360 TlME (MIN)

a. Effeet of heat trsnsfer coefficient hl at didferent depths.

O 60 t 20 180 240 300 360 TlME (MIN)

b* Elkt of heat t rader coefficient L at 15 cm (6") k k w the top surface.

Figure 4.7. Wixt of heat trruisler eodncknt at the water.cooled wriis, hl.

It is clear €tom Figure 4.7 that the modification of heat transfer coefficient at the

water-cooled bottom and side walls has an impact on the smelt cooling rate, especiaiiy at the

location close to the wails. During a smelt cooling process, the temperature of smelt in the

core region is always important. Thenfore. the focus is placed on this core. In this test, when

the heat-transfer coefficient was doubled the change in the smelt temperature in the con

ngion (15 cm below the top surface) was not very large, only about 40°C after 6 houa of

cooling when the smelt had been completely solidified.

It cm be concluded that the change in boundary conditions does affect the smelt

cooling rate, but the impact is limited especiaily in the con region. This is another proof that

the smeit cooling rate is mainly controlled by heat conduction fmm the srneit interior.

#en the heat-transfer coefficient outside the uncooled walls was changed, the effect

of this change on the smelt cooling rate cm be seen in Figure 0.7. A ciramatic change in the

heat-transfer coefficient, h2, had little impact on the entire smelt cooling process. The change

in the smelt cooling rate due to the variation of h2 is negligible even at the location far from

the watercooled wall and close to the uncooled wall when the smelt temperature was more

likely inlluenced by h2.

In the field test, the uncooled walls were thermally insuiated with 7 cm (2'1;) thick

firebricks. Heat-transfer coefficient, hz, actually described the heat removal at the outer

surface of the firebricks. Because of the high temperature gradient which iikely existed in the

firebricks due to their low thermal conductivity. the temperature at the outer surface of the

firebrick was rnuch lowet than on the steel walls. Therefore, the heat loss fiom the uncooled

wall was diminished, and the impact was very limited in view of the huge amount of smelt

heat which bad to be removed in the pcesent experiment.

4.7.7 Summary of parametric studies

The impact of major parameters on the pndicted smelt cooling process are

summarized and tabulated in Table 4.4, in which these parameters are modified by some

extent from the standard case (as show in Table 43), and the new results are compared with

the data caiculated fkom the standard case.

Table 4.4. Summary of the effkct of parameters on the predicted smelt codiag process.

1 Parameter

Emissivity

Heat capacity

Latent heat

conductivity

p=LOaO

C, = 2840

(pCp constant)

C, =2l3O

(50% increase)

ADL nata

= 0.45

kr- = 0.88)

Result

There was no change in the

result of calculation

It only affccted the cooling

rate of smelt near the top

surface. The smelt interior

(15 cm inside) was not

stffected.

After 6 hours of cooling, the

smelt temperature difference

was as high as 100°C at al1

depths in the smelt pool.

There was no large change

on the smelt cooling rate.

The smelt was cooled faster.

S ince the smelt temperature

was unâer the present

melting temperature most of

Discussion --

p C , was combined

together when it was

used in the calculation.

Radiation at the top

surface had a minor

impact on the smelt

cooling process.

Significant effect

Minor effect

Significant effect

The smelt cooling rate

was mainly conmlled

by the heat nmoval

Thermal

conductivi ty

Boundary

condition

(heat-tram fer

coefficient)

h 1 = 26 (doubled)

for the water-

cooled walls

h2 = 12 (doubled)

for the uncooled

walls

the time during the cooling

process, k = 0.88 was mainly

used.

Large temperature difference

on the smelt cooling rate

Thcre was a large smelt

temperature difference at the

location close to the water-

cooled wdls, but the eflect

on the core region of the

smelt pool was limited

Very linle change on the

entire smelt cooling process

from the smelt interior

via heat conduction.

The value 0.88 was

high, and the values 0.4

and 1 .O were al1 far frorn

the actual value.

Very lirnited efiect on

the cooling rate of the

smelt core region

The cooling process of kraft recovery biler smelt was investigated systematically

using a laboratory-scale experimental apparatus. Smelt samples were obtained directly from a

smelt spout of an operathg kraft recovery boiler, and cooled in the experimental apparatus.

By measuring the smelt temperature, the cooling chanicîenstics were obtained. The nsults

showed that smelt cooling is a slow heat transfer process and mainly controlled by the rate of

heat conduction within the smelt interior due to the low thermal conductivity of srnelt.

There are two major factors affecthg the smelt cooling process: thermal conductivity

of fiozen smelt and the nlease of latent heat of smelt during solidification. The low thermal

conductivity of frozen smelt hinders heat removal from the smelt interior. thenfore the smelt

cooling rate is actually controlled by heat conduction in the smelt. hcreased cooling at

boundaries can not speed up the smelt cooling process substantially.

The amount of latent heat of smelt released duhg the cooling process is only a few

part of the total heat load. Its variation generally does not change the pend of required

cooling time significantly. However. since the latent heat of smelt is released within a very

narrow temperature range and the heat released can not be nmoved quickly due to the slow

hcat conduction in fiozen smelt even close to the boundacies, the smelt interior will stay at

this temperature level for a nlatively long time. In this study. the smelt stayed at about 720°C

for 2 hours in the expecimental apparatus. In a na1 recovery boiler, due to the large size of

lower tùmace and huge inventory of smelt, the smelt temperature may remah at around the

first melting point of smelt (720 - 730°C) for hours or even days.

On the other hand, the heat transfer rate is quite fast when the smelt is in completely

molten phase. This cm not be simply attributed to a higher thermal conductivity of molten

smelt. A "heat flow" in the smelt pool and radiation as well as convection may bc involved in

this cwling Gage. More detailed study on cwling of molten smelt will be nquired in the

funire to hilly understand this issue.

A transient 34mensional heat transfer model has also been developed on a

PHOENICS platfonn to simulate and predict the smelt cooling process in the experimental

apparatus. The simulation results showed good agreement with the expenrnental data

especially within the smelt pool interior.

Both experimental and simulation results indicated that the thermal conductivity of

lrozen smelt is likely about 0.60 W/m°C, which is smailer than the value of 0.88 W/mOc

previously suggested by ADL (Richardson and Memam, 1977).

The heat transfer model developed in this study was dso used to analyze and evaluate

the impact of several key parameters on the simulated smelt cooling process. The results

showed that the boundary conditions and latent heat of smelt have only a minor effect on the

cooiing of smelt, but the heat capacity of smelt can radical1 y change the smelt cooling rate.

The agreement between the experirnental and simulation results clearly shows the

availability and capability of the heat muisfer mode1 developed in this study for predicting

the smelt cooling process. But there are still some uncertainties which quire lurther

considerations.

The smelt pool investigated in this study was about 33~33x28 cm (13"~13"~11"

LxWxH). But in a real rocovery boiler, the char bed n o d l y has a cross-sectional size of

LOmxlûm, and a height of more than lm after the initiation of an ESP. The ratio of height to

length or width is only about 1110. It is very difficult to make a prototype and nach a

dynamic similarity .

In this study, since the shape of the smelt pool was nearly cubic, which means a less

ratio of top surface to the total surface area of smelt than the situation in a recovery boiler,

the effect of heat transfer from the top surface was less. However, because the ratio of the

side wall areas to the total outer area of the smelt cooling vesse1 was much bigger than that in

a recovery boiler, the smelt cooling in this study was more sensitive to the heat loss h m the

side boundaries.

Heat conduction in molten smelt appeared to be relatively very fast in this study,

which bhgs about a conflict with the nliability of the thermal conductivity value of molten

smelt recommended by ADL and used in the numerical simulation.

6.2 Reconunenàatioas

Mo= detailed investigation into the effective thermal conductivity of the molten

smelt should be further conducted.

In a real recovery biler shutdown. the smelt is generally covered with a layer of char.

A M e r study of smelt cooling under this condition should be conducted.

As discussed in Section 2.2, to expose the hot intenor material of a char bed can

dtamaticaily increase the heat transfer rate From the bed surface. Therefore, if the partially

cmsted smelt surface cm be broken and some coolants such as NaHC03 and liquefied CO2

can be used to cool the intenor material. smelt cooling will be accelerated and the cooling

time may be significantly shortened. This is a practice beginning to be adopted in industry.

The experimental and theoretical studies on this topic are highly recommended.

Adams, T.N., Fredecick, W.J., Grace, TM., Hupa, M., Lisa, K., Jones, A.K., and Tran, H.,

"Kraft Recoverv Boilers", TAPPI PRESS, Atlanta, GA, 1997.

Agranat, V., Kawaji, M., and Tm, H., "Modelling of lower Mace heat transfer in recovery

boilen: analysis of high floor tube temperature excursions", prepared for Irnproving

Recovery Boiler Performance, Emission and Safety Consortium, University of

Toronto, 1998

Agranat, V., Kawaji, M., and Tran, H., "Development of a recovery boiler lower fumace heat

transfer mode1 - part 1: analysis of floor tube temperature excursions", proceedings of

the TAPPI 1997 Engineering and Papemialcers Conference, pp 1 13 1- 1 139, Nashville,

Tennessee, 1997.

Bennett, C.O., and Myers, JE., "Momentum. Heat. and Mass Transfer", McGraw-Hill Book

Company, Inc., New York, 1962.

Bird, RB., Stewart, W.E. and Lightfoot, E.N., Tranmrt Phenornena", John Wiley & Sons,

Inc., New York, 1966.

Finck, J .L., "hproved Apparatus for Measuring Themai Conductivity of Refrac tories at

High temperatures", Journal of Arnerican Ceramic Society, 37[1], 378-382, 1937.

Grace, T.M., "Recovery Boiler Explosion", pnsented at the 1986 Kraft Recovery Operations

Seminar, TAPPI PRESS, Atlanta, G A 1986.

Grace, T., Walsh, A., Jones, A. et al., ' 1989 International Chernical Recovery Conference

Proceedings, Technicd Section". -PA, Montreal, pp. 1-8

Grace, T.M., Leopold, B., and Malcolm E.W., "PUID And Pamr Manufacture. Volume 5:

Alkaline Pul~ing", The Joint Textbook Cornmittee Of The Paper Industry, TAPPI,

CPPA, 1989.

Grace, TM., "Bed Cooling Following An ESP," 1998 International Chemical Recovery

Conference Proceeding, TAPPI PRESS, p.355-365.

Grace, TM., "1999 TAPPI Recovery Boiler Short Course, Chapter 6.5 Smelt-water

Explosions", Orlando, Fl, 1999.

Green, R.P., and Hough, G., " Chemical Recoverv In The Alkaline Pul~ine Processes", Third

Edition, prepared by the alkaline pulping cornmittee of the pulp manufacture division,

TAPPI PRESS, Atlanta, GA, 1992.

Grimvall, G., 'Thermal Pro~erties of Materials", North-Holland, Amsterdam, 1986.

Harrison, R.E., and Ariessohn, P.C., "Application Of A Smelt Bed Imaging Systemt', 1985

TAPPYCPPA International Chemical Recovery Conference Proceedings. TAPPI

PRESS, Atlanta, GA.

Holrnan, J.P., "Heat Trans fer", Eighth edi tion, McGRAW-HILL, Inc ., New York, 1997.

Jones, A.K., and Lefebvre, B.E., "Experimental And Computational Modelling Of Smelt

Pool Cooldown", ABB report prrscnted at CPPA Stem and Power Subcomrnittee

Meeting, Prince Rupert, B.C., 1996.

Jones, A.K., "1995 TAPPI Kraft Recovery Short Course, 64, CFD Modelling as a Problem

Solving TooP', Orlando, FL, 1995.

Jones, AK. and Cben, K., "CFD Modelling for Retrofit Evaluation", 1995 International

Kraft Recovery Conference, A123-13 1, Toronto. ON, 1995.

Kawaji, M., Nickfarman, H., Tan, G., Grace, T.M., and Tran, &, "Recovery Boiier Char Bed

Cooling Foilowing An Emergency Shutdown", Task 1, review and interpntation of

available information, prepared for American Forest & Paper Association's Recovery

Boiler R & D Subcommittee, University of Toronto, 1999.

Knudsen, M., "Starting with PHOENICS-VR (Version 3.1 )", Heat and Momentum Limited,

London, 1998.

Laubitz, M.J., "Measurement of the Themal Conductivity of Solids ai High Tempe rature by

using Steady-state Linear and Quasi-linear Heat Flow', Chapter 3 in 'Thermal

Conductivity", Volume 1, Academic Press, London, 1969.

Lien, S.J. and Horton, R.R., "A Review of Recovery Boiler Mode1 Applications", 1995

International Kraft Recovery Conference, A 133- 148, Toronto, ON, 1995.

Lovell, M.C., Avery, AJ., and Vernon, M.W., "Phvsical Pro~erties of Materials", Van

Nostrand Reinhold Company, Berkshire, 1976.

Mimms, A., Kocurek, MJ., Pyatte, J.F., and Weight, E.E., "Kraft Pul~in~. A Com~ilation Of

Notes", TAPPI PRESS, Atlanta, GA, 1993. - Richardson, D.L., and Memam, R.L., "Study Of Cooling And Smelt Solidification In Black

Liquor Recovery Boilers", Phase 1 report, prepared for the American Paper Institute,

Arthur D. Littie Inc., Cambridge, MA, Feb. 1977.

Richardson, D.L., and Memam, RL., "A Study Of Black Liquor Recovery Furnace Firing

Conditions, Char Bed Characteristics And Performance", Phase II report, prepared for

the Amencan Paper Institute, Arthur D. Little Inc., Cambridge, MA, Dec. 1978.

Salcudean, M. et al., "Black Liquor Combustion Vaiidated Recovery Boiler Modelling Five

Years Report, Appendix P', Prepared for the US Department of Energy, 1996.

Shick, PE., and Grace, TM., "Review Of Smelt Water Explosions", Presented at 1981

International Confennce on Recovery of Pulping, Vancouver, B .C.. 198 1.

Smook, GA., "Handbook For MD & Pawr Technolonists", 2nd Edition, Angus Wilde

Publications hc., Vancouver, BC, 1994,

Veml, CL. et al., "Recovery Fumace Simulator-Design and Modeling", 1995 International

Kraft Recovery Confennce, A 1 1 1 - 122, Toronto, ON, 1995.

Williamson, P.N. et al., "A review of World Warkets", Tappi J. 83(1):34 (2000).

Emergency Shutdown Procedure (ESP)

The Black Liquor Recovery Boiler Advisory Cornmittee (BLRBAC) has developed

and recommended an ESP. Cumntly ESP has become a standard procedure in pulp and

paper industry. An ESP can be briefly desdbed as follows:

Sound an alann to clear the recovery area of al1 unnecessary personnel.

immediately stop firing al1 fuel. Divert black liquor. Secure the unit's auxiliary fuel

system at a remote location.

Immediately shut off feedwater and al1 other water sources to the boiler except for

smelt spouts.

Shut down the air supply to p n m q air ports immediately. Provide a balanced draft

and an air supply above the char bed to purge gases from the fumace.

Drain the boiler as rapidly as possible and in accordance with manufacturer's

recommendations to a leveI2.4m (8 feet) above the low point of the himace flwr.

Reduce s t em pressure as rapidly as possible afker the boiler has been drained to the

2.4m (8 k t ) level.

Appendur B

Raw Experimental Data

The fmt field test was conducted at Espanola mill, Eddy Specialty Papers Ltd.

Totally 32 type K thermocouples were used to measure the temperatures of smelt, cooling

water and air during the snieit cooling process. Therrnocouple 21 (TC21), 9 (TC9). and

15 (TCIS) were used to measure the temperatures of cooling water at inlet, outlet, and

side cooling coi1 outlet. Thennocouple 27 (TC27) was used to measure the temperature of

cooling air ai the lid of the experimental apparatus.

TIME MIN 204 205 206 207 208 209 210 211 212 213 214 215 216 217 218 219 220

Smeit Temperature DiiMbution at DMferent Depths

O 60 120 180 240 300 360 i: O

Time (min)

O 60 ..

120 180 240 300 360

Time (min)

Figure C.2. Smeit temperature distribution at 25 cm (IV') k b w the top d a c e .

Appendix D

Properties of Castable Powder

Appendk E

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* Setobjects: x0 y 0 a) * dx dy &

name XPO= 0.000000Em; YPO= 6,667500E-02; ZPO= 6.350000E-03 XSI= 3.175000E43; YSI= 3. t86SûûE-O 1 ; B I = 2.984500E-O 1 RSET(BrnALL1 XPO= 3.03 L75OE-01; YPO= 6.667500E-02; ZPO= 6.350000E-03 XSI= 3.17SûûûE-03; YS11 3J86SoOE-O 1; ZSI= 2.984500E-0 1 R=nBm-AL) XPO= 0.000000E+00; YPû= 6.350000E-02; Zf03 3.O48ooOE-û 1 XSI= 3.063500E-OI ; YSI= 3.2~ooOoE-O 1; ZSI= O.ûûûûûûE+OO RSET(B,*RADTOP ) XPO= 0.000000EE+ûû; YPO= 6.350000E-02; ZPOz O.ooooOOE+ûû XSk 3.0636ûQE-OI; YSI= 3.250000E-01; ZSI=6350000E-O3 RSET(B,BCYIWALL ) XPO= 0.000000E+Oo; YPO= 0.000000E+oo.; P>O= 0.000000E+O

XSk 3.0635ûûE-0 1 ; YSI= 6.350000E-02; ZSI= 3.048oOOE-0 1 RSET(B&BRICK ) xeOz O,OOOOQOE+OO; Y K k 3.885OûûE-01; ZPû= 0.000000E+ûO XST= 3.063500E-01; YSI= 6.349996E-02; ZSI= 3.048ûûûE-0 1 RSET(B,B-BRICK ) XPO= 3.0635ûûE-û 1; YPO= O.ûûûûûûE&; ZPO= 0.000000E+00 XSI= 6.350000E-02; YSI= 4.S2oooOE-O 1 ; ZSI= 3.048ûûûE-û 1 RSET(B,R-BRICK ) XPû= O.O00000E+00; YPO= 6.350000E-02; ZPO= 0.000000E+OO XSI= 3.0635ûOE-0 1 ; YSk 3.2SOOOOE-0 1 ; ZSI= O . O O O E + O O RSET(B,BOTPLATE) XPO= O.ûûûûûûE+00; - 6350E-02; ZPO= 0.000000Em XSI= 0.0000ûûE+ûû; Y SI= 3 250000E-0 1 ; ZSI= 3 .OWlOOE-û 1 RSET(B,LEFïPLAT) XPO= 3698SûûE-O 1; YPO= O,ûûûûûûE+OO; ZPO= O,OûûûûûE+00 XSI= 0.000000E+00; YSI= 4520000E-01; B I = 3.0480E-0 1 RSET(B,NGHTPLT) XPO= 0.000000E+00; YPû= 0,0000ûûE+00; ZPû= O.OOûOûûE+ûO XSI= 3.6985ûûE-01; YSI= O.ûûûûûûE+oo., ZSI= 3.048000E-O 1 RSET(B ,FRONTPLT) XPO= 0.000000E+Qû; YPû= 4.52OOOOE-û 1 ; DO= 0.000000E+00 XSI= 3,698500E-01; YSI= 0.000000E+00r ZSI= 3.O48OOOE-û 1 RSET(B$ ACKPLAT) XPO= O.OûûûûûE+ûû; YPO= 6.350000E-02; 250= 6.350000E-03 XSI= 3.063500E-0 1 ; YSI= 3. I7SOOOE-03; ZN= 2.984500E-0 1 RSET(B,FN'IWALL ) XPû= O.ûOûûûûE+OO; YPO= 3.8532SOE-O 1 ; ZPO= 6.350000E-03 XSI= 3.0635ûûE-O 1 ; YSI= 3.17SûûûE-03; ZSI= 2.984SoOE-O 1 RSET(B,B ACKWALL)

* Modify defadt grid RSET(X, 1,2,1.000000E+ûû) RSET(X,2,- 1 1,1.600000E+00) RSET(X,3,2,1 .a00000E+00) RSET(XA~,1.000000E+Oû) RSET(Y, lt2,1.000000E+00) RSET(Yt2J, 1.000000E+OO) RsET(Y 3,-13; L .600000E+W) RSET(Y ,4,2,l.OOOûûûE+00) RSET(Y,S,2,l.oooooOE+00) RSET(Z,1,2,1 .ûûûûûûE+00) RSET(Z2,-28.1 .dOOOOOE+00) *******************************f**************************** Gmup 6. Body-Fitted coordinates ************************************************************ Gmup 7. Variables: STOREd,SOLVEd&AMEd ONEE'HS = T

* Non-default variable names NAME(146) =KOMI ; NAME(148) =DEN1 NAME(149) =SPHl; NAME(I50) =TEMI

* Solved variables list SOLvE(TEM1)

* Stored variables List STORE(SPH1 ,DEN1 ,KOND)

* Additional solver options SOLUTNWM 1 ,Y,Y,YflS*Y)

Group 9. Roperties RH01 =2.000E+O3 PRESS0 = l.OQOE+OS TMPI A = l.@OE+03 ;TMPlB = 1.42OE+O3 ;TMP IC = 1 .OO3E+O3 TEMfO =2.730E&2 CP1 =GRNDS PHNHlA = 1.4SOE.tOS ENUL = 1.544E-05 ;ENUT = 0.000E+00 DVOLDT = 3.410E-03 PRNDTL(TEM 1) = -6.500E-0 1 TMPl A = 1.420E+03 ************************************************************ Group LOhter-Phase Transfer Rocesses ************************************************************ Group I 1 .Initialise VadPorosity Fields FIINIT(K0NQ) = 6.OOOE-û 1 ;FIINIT(DEN 1) = 2.000E43 FIINIT(SPH1) = 1.420E+03 ;FIXNïï(TEM 1) = 7.400E+02 No PATCHes used for this Group

Group 12. Convection and diffusion adjustments No PATCHes used for this Group

************+*********************************************** Group L3. Boundary & Speciai Sources

PATCH (*RADTOP ,HIGH ,3,0,0,0,0,0,1,360) COVAL (*RADTOP ,TEM 1,5392E-O8,3.728E+û 1)

PATCH (Bû'ïPLATE,LO W ,S0,090,0,09091,360) COVAL (BOTPLATE,TEM l,l3OOE+ûI, 2.SoOE+û 1)

PATCH (LEFïPLAT,WEST ,51,0,0,0,0,0,1,360) COVAL (LEFI'PLATJEMl, l.3oOE+ûl. 2.500E+01)

PATCH (RIGHTPLT,EAST ,!Q,O.O.O,O,O. 1,360) COVAL (RTGHTPLT,TEMl, 6.ûOOE+OO, 2500E+01)

PATCH (FRONTPCT,SOUTH ,!S.O,O,O,O,O, 1,360) COVAL (FRONTPLTrTEMl, 6.000E+00,2.5ûûE+O 1)

PATCH (BACKPLAT,NORTH ,S4,O,O,O,O,O, 1,360) COVAL (BACWLAT,;IEMI, 6.000E+o. ZMOE+û 1)

EGWF = T ************************************************************ Group 14. Downstteam h u r e For PARAB ************************************************************ Group 15. Tenninate S weeps LSWEEP= 30 SELREF = T RESFAC = I.OOOE-03 ************************************************************ Group 16. Terminate Iterations ************************************************************ Group 17. Relaxation ************************************************************ Group 18. Limits ************************************************************ Group 19. W T H Calls To GROUND Station ASAP = T ************************************************************ Group 20. Reliminary Printout ECHO = T ************************************************************ Group 2 1. Print-out of Variables ************************************************************ Group 22. Monitor Pcint-Out rXMON = 8;IYMON= 11;IZMON= 15 NPRMNT = 1 TSTSWP = -1 ************************************************************ Gmup 23Iield Pnnt-Ouc & Plot Contcol No PATCHes used for this Group ************************************************************

Group 24. Dumps For Restarts NOWIPE = T

> DOM, Sm 3.698SOOE-û l,4.S2ûoE-û 1,3.048ûûûE-0 1 > DOM, MONIT, 1 .53 13ûûE-0 1, 2.26OOûûE-û 1, l.27OOûûE-û 1 > DOM, SCALE, L-OOOOQOE+00~ 1 .000000E+00, 1.000000E+00 > DOM, SNAPSIZE, 1.000000E-02 >DOM, RELAX, S.OOOOOOE41

> OBI1, NAME, LEFIWALL > OBI 1, POS~ON, O,ûûûûûûEtûû, 6,667500E-02,635ûûûûE-03 >OBJI, SIZE, 3.175000E-03,3.186500E41,2984500EQI >OBJl, CLPART, cubet > OBJ 1, ROTATION, t

> OBJl, > OBJI, > OBJl,

> OBJ2, > OBJ2, > OBJ2, > OBJ2, > OBJ2, > 0852, > 0852, > OBJ2,

> OBJ3, > OBJ3, > OBJ3, > OBJ3, > OBJ3, > OBJ3,

> OBJ4, > OBJ4, > OB J4, > OBJ4, > OBJ4, > OBJ4, > OBJ4, > OBJ4,

> OBJS, > OBJS, > OBJS, > OBJ5, > OBJ5, > OBJ5, > OBJ5, > OBJS, > OBJS,

> OBJ6, > OBJ6, > OBJ6, > OBJ6, > OBJ6, > OB J6, > OB J6, > OBJ6, > OBJ6,

> OBJ7,

TYPE, BLOCKAGE M A . , 111 HEAT* 0.000000E+00

NAME, RiGHTWAL POSITION, 3-03 l7SOE-O 1,6.667500E-02,6.350000E-03 SIZE, 3,175000E-03,3.18650QE-01,2.984500E-O1 CLIPART, default ROTATION, t TYPE, BLOCKAGE MATERIAL, 1 1 1 m T , 0.000000E+Oo

NAME, *RADTOP POSITION, O,OOOûûOE+ûû, 6.3Sa000E-02,3.0480ûûE-O1 S m , 3.0635ûûE-0 1,3.250000E-O 1,0,000000E+ûO CLIPART, default ROTATION, 1 TYPE, USER-DE-

NAME, BOTWALL POSITION, 0.000000E+00,6.350000E-02,0.000000E+00 S m , 3.0636ûûE-û 1,3.25innnnE-û1,6.350000E-03 CLIPART, default ROTATION, 1 TYPE, BLOCKAGE MATERIAL, 1 1 1 W T * 0*000000E+00

NAME, PROBE 18 POSKION, 1.298 142S-02,2.06848 1 E-û 1,l .W39 1 1 E-02 SIZE, 1.9921 1 lE-02,3.830390E-02,7.4 l7232S-03 CLIPART, cube3 ROTATION, 1 TYPE, PROBE TEMPERATURE, 2.000000E+O 1 PRESSURE, 0.000000E+00 VELoCm* 0.00000QE+o, 0.000000E+OQ, 0*000000E+00

NAME, PROBE12 POSITION, 1.3 19593E-0 1.2.06848 lE-01, 1,9039 1 tE-02 SIZE, 4.243 138E-02,3.830390E-02,7.4 17232E-03 CLIPART, cube3 ROTATION, 1 TYPE, PROBE TE.pERATüRE, 2.000000E+OI PRESSURE, 0.000000E+00 VELOClTY, 0*000000E+QO, 0.000000E+00,0*000000E+00

> OBJ7, > OBJ7, > OB J7, > OBJ7, > OBJ7, > OBJ7, > OBJ7,

> OBJ8, > OBJ8, > OBJS, > OBJ8, > 0858, > OBJS, > OBJ8, > OBJS, > OBJ8,

> OBJ9, > OBJ9, > OBJ9, > OBJ9, > OBJ9, > OBJ9, > OBJ9, > OBJ9, > OBJ9,

S m , 1.992lWE-O2,3.83039OE-O2,7.4 17232E-03 CLPART, cube3 ROTATION, 1 TYPE, PROBE TEl@ZRATURET 2.-E+O 1 PRESSURE, O.ûûûûûûE+Oo VEL,OCm, 0.000000E+00,0.000000E+00, 0*000000E+00

NAME, PROBE36 POSlTION, 1.298 142E-02,3.120563E-0 1.1.9039 1 1E-02 S m , 1.992 1 1 lE62,2,702889E62,7.4 l7232S-03 CLIPART, cube3 ROTATION, 1 TYPE, PROBE TEMPERATURE, 2.000000E+0 1 PRESSURE, O,000000E+Oo VELOCITY, O.OQOOOOE+OO, 0.000000E+00,0,000000E+00

NAME, PROBE30 POSITION, 13 19593E-01,3.120563E-0 1, 1.9039 1 LE-02 S m , 4.243 138E-O2,2.702889E-O2,7.4 17232E-03 CLIPART, cube3 ROTATION, 1 TYPE, PROBE TEMPERATURE, 2 . 0 0 O E 4 1 PRESSURE, 0.000000E+00 VELOCITY, 0.000000E+00,0.000000E+00,0*000000E+00

> OB J 10, NAME, PROBE24 > OBJ 10, POSITION, 2.734475E-01,3.120563E-û 1,1.9039 1 1E-02 > OB J IO, SIZE, 1.9921098-02,2.702889E-02,7.4 17232E-03 > OBJ 10, CLIPART, cube3 > OB J 10, ROTATION, 1 > OBJIO, TYPE, PROBE > OB J 1 O, TEMPERATURE, 2.00ûûûûE4 1 > OB J IO, PRESSURE, O.OûûûûûE+OO > OBJ 10, VELOCITY, O.ûûûûûûE+OO, O.ûOûûûûE+ûû, 0.0000ûûE+ûû

> OBJ 1 1, NAME, PROBE 17 > OBJ 1 1, POSITION, 1.298 lME-O2,2.06848 1E-0 l,6,73O lûûE-02 > OB1 1 1, SIZE, 1.9921 1 lE-02,3.830390E-û2, 1.26400lE-O2 > OBJll, CLIPART, cube3 > OB 11 I, ROTATION, 1 >OBJll, TYPE, PROBE > OB11 1, TEMPERATURE 2.00ûûûûE+0 1 > OBJ 1 1, PRESSURE, 0.000000E+ûû > OBJl L, VELCMXTY, O.ooooOOE+ûû, 0.000000E+00, 0.000000E+Oû

> OBJ 12, NAME, PROBE1 1 > OBJ12, POÛITION, 1 3 t 9593E=ûl, 2.06848lE-O L,6XNIlOOEO2 > OBJ 12, SIZE, 4.243 l38E-û2,3.83O390E--O2,l.26400lE-M

> OBJ12, > OBJ12, > OBJ12, > OBJ12, > OBJ12, > OBJ12,

> OBJl3, > OBJ13, > OBJl3, > OBJl3, > OBJ13, > OBJl3, > OBJ13, > OBJ13, > OBJl3,

> OBJ14, > OBJ 14, > OBJ 14, > OBI 14, > OBJ14, > OBJ14, > OBJ14, > OBJ14, > OBJ14,

> OBJ15, >OBJlS, > OBJ15, > OBJ15, > OBJ15, > OBJlS, > OBJ15, > OBJ15, > OBJIS,

> OBJl6, > OBJ16, > OBJ16, > OBJ16, > OBJ16, > OBJ16, > OBJ16, > OBJ16, > OBJ16,

> OBJ17, > OBJ17, > OBW, > OBJ17,

CLPARTT cube3 ROTATION, I TYPE, PROBE TEMPERATURE, 2.000000E41 PRESSURE, 0.00ûûûûEiûû VEL-9 O.OOOOOOE+QOT O.OOOOOOE+OOTO.OOOOQOE+OO

NAME, PROBES POSITION, 2.734475E-0 l,2.06848 IE-0 l,6.730 100E-02 SIZE, 1.992 109E42,3.830390E-02, 1.264001E-O2 CLIPART, cube3 ROTATION, 1 TYPE, PROBE TEMPERATURE, 2.0ûûûûûEiû1 PRESSURE, 0.000000E+OO VELOCITY, 0.000000E+00,0.000000E+00,O.OOOOQOE+00

NAME, PROBE35 POSKXON, 1.298 142E-02,3.120563E-0 l,6.730 100E-02 SIZE, 1.9921 1 lE-02,2.702889E-û2,1.26400 1E-02 CLIPART, cube3 ROTATION, 1 TYPE, PROBE TEMPERATURE, 2.000000E+0 1 PRESSURE, 0.000000E+00 VELOCITY, 0.000000E+00, 0.000000Et00,0*000000Ern

NAME, PROBE29 POSITION, 1.3 lgSWE-01, 3.120563E-O 136,730 100E-02 SIZE, 4.243 138E-û2,2.7O2889E-O2, 1.26400 1 E-02 CLPART, cube3 ROTATION, 1 TYPE, PROBE TEMPERATURE, 2.000000E+O 1 PRESSURE, 0.000000Ern VELOClTY, O~OOOOOOE+OOT 0.000000E+00, 0*000000E+00

NAME, PROBE23 POSITION, 2.734475E-0 lT3.120563E-0 1,6.730 100E-02 SIE, 1.992 lO9E-û2,2.7OB89E-O2,1.264OO 1E-02 CLIPART, cube3 ROTATION, t TYPE, PROBE TEMPERATURE, 2.000000E41 PRESSURE, O.OOûûûûE+ûû V E L m , 0.000000E+00,0.000000E+00,0,000000E+00

> OB J 17, ROTATION, 1 > OBJ17, TYPE, PROBE > OBJ17, TEMPERATURE, 2.000ûûûE+û 1 > 08517, PRESSURE, O.QOQOOQE+OO > OBJ 17, VELOCITY, O.OûOOE+ûO, 0.000000E+00,0.ûûûûûûE+00

> OBJ 18, NAME, PROBE10 > OBJ 18, POSITION, 1.3 LgS93E-O 1,2.04848 1 E-0 l,l.229576E-O 1 > OBJ 18, SI=, 4,243 138E-û2,3.83039OE-O2, 1 S93217E-02 > OBJ 18, CLIPART, cube3 > OBJ 18, ROTATION, 1 >OBJ18,TYPE, PROBE > OBJ 18, TEMPERATURE, 2.0000ûûE+û 1 > OBJ 18, PRESSURE, 0.000000E+00 > OBJ 18, VELOCITY, 0.000000E+00,0.000000E+00,0.000000E+00

iJ19,NAME, PROBE4 ;J 19, POSITION, 2.734l47SE-O 1,2.06848 1 E-0 1, l.229576E-û 1 iJ19, S m , 1.992109E-02,3.830390E-02, 1 S93217E-02 iJ 19, CLIPART, cube3 iJ 19, ROTATION, 1 iJ19, TYPE, PROBE ;J 19, TEMPERATURE, 2.000000E+O I iJ19, PRESSUMI, 0.000000Ern ;J 19, VELOCITY, 0.000000E+OO, 0.000000E+ûû, 0.000000Et00

> 08120, NAME, PROBE34 > 08520, POSITION, 1 .B8 142E-02,3.120563E-0 1, 1.229576E-û 1 > 08120, SIZE, 1.9921 1 lE-02,2.7028898-02, 1.5932 17E-O2 > 08520, CLPART, cube3 > 08520, ROTATION, 1 > OBJ20, TYPE, PROBE > OBJ20, TEMPERATURE, 2.000ûûûE+O 1 > OBJ20, PRESSURE, 0.000000E+00 > OBJZO, VELOClTY, 0.000000Em, O.OûûOûûE+ûû, 0.000000E+Oû

> 0BJ2 1, NAME, PROBE28 > OBJ21, P O S ~ O N , 1.3 19593E-ûl,3.12OS63E-O 1, l.229576E-O 1 > OBJ2 1, S m , 4.243 l38E-û2,2,702889E-O2,l S932 l7E-02 > OB J2 1, CLIPART, cube3 > OB J2 1, ROTATION, 1 > OBJ21, TYPE, PROBE > OB52 1, TEMPERATURE, 2.000000E+0 1 > OBJ21, PRESSURE, O . O E + O O > OBJ21, VELOCiTY, 0.000000E+00,0.ûûûûûûE+00,0.000000E+00

> OBJ22, NAME, PROBE22 > OBJ22, POSITION, 2.7M475S-0 1.3-1 SOS63E=Ut, 1 .ZgS76E-O 1 > OBJ22, SIZE, 1.992109E-(n, 2,702889E-û2,1 .S932lîE-O2 > OBJ22, CLIPART, cube3 > OBJ22, ROTATION, 1

> 08522, TYPE, PROBE > OB J22, TEMPERATURE, 2.0ûûûûûE+O 1 > OBJ22, PRESSURE, 0.000000E+00 > OB J22, VELOCITY, 0.0000ûOE+00,0.000000E+00,0.000000E+00

> OBJ23, NAME, PROBE15 > OBJ23, POSïTïON, 1.298 l42E-O2,S,O6848 1E-û 1, l.722603E-û 1 > OBJ23, SIZE, 1.992 1 1 1 E-O2,3.83039OE-02, 1 S93220E42 > OBJ23, CLPART, cube3 > OBJ23, ROTATION, 1 > 08523, TYPE, PROBE > OBJ23, TEMPERATURE, 2.000000E+û 1 > OB123, PRESSURE, 0.000000E+00 > OBJ23, VELûClTY, 0.000000E+00,0.000000E+ûO, 0.000000E+ûû

> OBJ24, NAME, PROBE9 > OBJ24, POSITION, 1.3 19593E-0 1,2.06848 1 E-0 1, 1 .'?22603E-O 1 > OBJ24, SIZE, 4.243 138E-û2,3,83039OE-û2, 1 593220E-02 > OBJ24, CLPART, cube3 > OBJ24, ROTATION, 1 > OBJ24, TYPE, PROBE > OB JU, TEMPERATURE, 2.000000E+O 1 > OBJ24, PRESSURE, O.ûûOûOûE+OO > OBJ24, VELOCITY, O.ûûOûûûE+ûû, O.ûûOûûûE+OO, 0.0000ûûE+00

BJ25, NAME, PROBE3 BJ25, POSITION, 2.734475E-0 1,2.06848 1 E-0 1,1.722603E-û 1 BJ25, SIZE, 1.992 lOgE-O2,3.83039OE-û2, 1 393220E-02 iBJ25, CLIPART, cube3 iBJ25, ROTATION, I 16525, TYPE, PROBE BJ25, TEMPWTURE, 2.000000E+01 1BJ25, PRESSURE, O.OûûûûûE+ûû iBJ25, VELOCITY, 0.000000E+00,0.ûOûûûûE+00,0.ûûûûûOE+00

> OBJ26, NAME, PROBE33 > OB J26, POSITION, 1-298 l42S-02,3- l2OS63E-O 1, l.722603E-û 1 > OBJ26, SIZE, 1.9921 1 1E-O2,2.702889E-02, 1 S9322OE-02 > OBJ26, CLIPART, cube3 > OBJ26, ROTATION, 1 > OBJ26, TYPE, PROBE > OBJ26, TEMPERATURE, 2.ûûûûûûE+O 1 > 08126, PRESSURE, O.OûûûûûE+00 > OBJ26, VELûCïTY, O.ûûûûûûE~, O.ûûûûûûE+00,0.0ûûûûûE+00

> 08127, NAME, PROBE27 > OBJ27, PûSlTïON, 1.3 lgSWE-ûl, 3.120563E41,1.722603E-O 1 > OBJ27, SIZE, 4.243 1?8E-M,2.702889E-O2,1 SWZOEQ2 > OBJ27, CLIPART, cube3 > OBJ27, ROTATION, 1 > OBJ27, TYPE, PROBE

> OBJ27, TEMPERATURE, 2 . 0 0 E 4 1 > OBJ27, PRESSURE, 0.000000E+00 > OB J27, VELOCITYT O.OOOOOOE+QOT 0.000000E.t00, 0.000000E+00

BJ28, NAME, PROBE21 BJ28, POSlTION, 2.734475E-O 1,3.120563E-O 1,1.722603E-û 1 B J28, SIZE, 1.992 lWE42,2.702889E-02, 1 S93220E-02 BJ28, CLIPART, cube3 B J28, ROTATION, 1 BJ28, TYPE, PROBE B J28, TEIWERATURE, 2,OOOûûûE+O 1 IB J28, PRE!3SüRE, 0.000000E+00 1B J28, VELOCITY, 0.000000E+OO,O.O00000E+ûû, 0.000000E+00

> OBJ29, NAME, PROBE 14 > OBJ29, POSITION, 1.298 142E-02,2,06848 LE-O 1.2.176963E-0 1 > 08529, SIZE, 1.9921 1 lE-02,3.830390E-02, 1.35 1269E-02 > OB J29, CLIPART, cube3 > OB J29, ROTATION, 1 > OBJ29, TYPE, PROBE > OB 129, TEMPERATURE, 2.OûûOûûE4 I > OB 129, PRESSURE, 0.000000E+00 > OBJ29, VELOCl'IY, 0.000000E+00,0.~E+00,0.000000E+00

> OBJ30, NAME, PROBE8 > OBJ30, POSITION, 1.3 t9593E-01,2.068481E-0lV 2.176963E-01 > OBJ30, SIZE, 4.243 t 38E-O2,3,83039OE-O2, 1.35 1269E-02 > OBJ30, CLIPART, cube3 > OBJ30, ROTATION, 1 > OBJ30, TYPE, PROBE > OBJ30, TEMPERATURE, ZûûûûûOE+O t > OBJ30, PRESSURE, 0,ûûûûûûE- > OBJ30, VELOCITY, 0.000000E+009 0.000000E+00, 0~000000E+00

> OBJ3 1 , NAME, PROBE2 > OB J3 1, POSITION, 2,734475E-0 l,2.06848 1 E-0 1,2. I76963E-0 1 > OB J3 1, S E , 1.992 IO9EQ2,3.83039OE-O2, 1.35 l269E-02 > OBJ31, CLIPART, cube3 > OB J3 1, ROTATION, 1 > OBJ3 1, TYPE, PROBE > 08531, TEMPEWTURE, 2.ûûûûûûE+01 > OBJ3 1, PRESSURE, 0.000000E+ûû > 0BJ3 1, VELûCITYw 0.000000E+00,0.000000E+ûû, 0.0ûûûûûE+00

> OBJ32, NAME, PROBE32 > OBJ32, POSITION, 129% l42E-O2,3.L20563E-û 1,2. t76963E-û 1 > OBJ32, Sm, l.992lI LE= 2.702889E-û2, 1351269E-02 > OBJ32, CLPART, cube3 > OBJ32, ROTATION, 1 > OBJ32, TYPE, PROBE > OBJ32, TEMPERATURE, 2,000000Etûl

> OBJ32, PRESSURE, 0.000ûûûE+00 > OBJ32, VaOCKY, O.OûûûûûE+00,0.000000E+00,0.ûûûûûûE+00

> OBJ33, NAME, PROBE26 > OBJ33, POSITION, 1.3 19593E-01,3.120563E-0 1.2. t76%3E-0 1 > OBJ33, SIZE, 4.243 l38E-O2,2.702889E-O2, 1 .35 1269E-02 > OB J33, CLIPART, cube3 > OBJ33, ROTATION, 1 > OBJ33, TYPE, PROBE > OBJ33, TEMPERATURE, 2.ûûûûûûE-tû 1 > OBJ33, PRESSURE, 0.000000E+00 > OBJ33, VELOClTY, 0.00QOOOEtûû, 0.000000E+00,0.000000E+OO

> OBJ34, NAM., PROBE20 > OBJ34, POSITION, 2.73UTE-û 1,3.120563EQ 1,2,176963E-01 > OBJ34, SIZE, 1.992 lO9E-O2,2.702889E-O2,1.35 l269E-02 > OBJ34, CLIPART, cube3 > OBJ34, ROTATION, 1 > OBJ34, TYPE, PROBE > OBJ34, TEMPERATURE, 2.000000E+O 1 > 08534, PRESSURE, O.OûûûûûE+ûû > OBJ34, VELOCITY, 0.000000EtOO,0.000000E+00,0.~~0000E+00

> OBJ35, NAME, PROBE1 3 > 08535, POSITION, 1.298 142E-02,2.06848 1 E-O lV2.76O66SE-O 1 > OBJ35, SIZE, 1 -992 1 1 1 E-02,3.830390E-02,8.627206E-03 > OB135, CLIPART, cube3 > OBJ35, RûTATION, 1 > OBJ35, TYPE, PROBE >OBJ35, TEMPERATURE, 2.000000E+Ol > OBJ35, PRESSURE, 0.000000EtOO > OBJ35, VELOCIlTY, O,ûûOûûûE+OO, 0.000000E+00,0.000000E+ûû

> OBJ36, NAME, PROBE7 > OBJ36, POSITION, L -3 19593E-01,2.06848 1E-0 1,2.760665E-01 > OBJ36, SIZE, 4.243 138E-O2,3.83039OE-O2,8.6272WE-û3 > 08536, CLIPART, cube3 > OBJ36, ROTATION, 1 > OBJ36, TYPE, PROBE > OBJ36, TEMPERATURE, 2.000000EM 1 > OB136, PR&SURE. 0.000000E+ûû > OBJ36, VELûCiTY, O.ûûûûûûE+Oo, O.OOûûûûE+ûû, O.OOOûûûE+O

> OBJ37, NAME, PROBE1 > OBJ37, POSmONT 2-734475E-0 l,2.06848 LE4 1,2.760665E-O1 > OBJ37, SIZE, 1.992 lWE-O2,3.83O39OE-O2,8.6272OdE-O3 > OBJ37, CLPART, cube3 > OBJ37, ROTATION, 1 > OBJ37, TYPE, PROBE > OBJ37, TEMPERA'ïURE, 2.000000E+OI > OBJ37, PRESSURE, 0.000000E00

lJ38, NAME, PROBE31 J38, POSmON, 1.298 142E-O2,3.120563E-0 1 , 2,76066SE-01 iJ38, SIZE, 1.992 1 t lE-02,2.702889E-02,8.627206E-03 J38, CLTPART, cube3 iJ38, ROTATION, 1 1538, TYPE, PROBE 1J38, TF;,MPERATURE, 2.000000E41 J38, PRESSURE, O.ûûûûûûEt00 1138, VEL(XXï'Y, 0.000000E+ûO, 0.000000E+00,0.000000E+OO

J39, NAME, PROBE25 J39, POSITION, 1.3 19593E-0 1,3.120563E-O 1.2.760665E-0 1 1539, S I E , 4.243 138E-û2,2.702889E-û2,8.6272O6863 1539, CLPART, cube3 1539, ROTATION, 1 1539, TYPE, PROBE iJ39, TEMPERATURE 2.000000E+û 1 iJ39, PRESSURE, 0.000000E+ûû 1539, VELOCITY, 0.000000E+00,0.000000E+OO, O.OOûûûûE+OO

> OBJ40, NAME, PROBE19 > OBJ40, POSIITION, 2.734475E-0 1,3.120563E-0 l,2.760665E-O 1 > OBJ40, SIZE, 1,992 lO9E-O2,2.702889E-O2,8.6272O6E-O3 > OBJ40, CLIPART, cube3 > OBJ40, ROTATION, 1 > OBJ40, TYPE, PROBE > OBJ40, TEMPERATURE, 2.000000EN 1 > OBJ40, PRESSURE, 0.000000E+00 > OBJ40, VELOCITY, 0.000000E+00,0.000000E+Oû, 0.000000E+00

J4 1, NAME, R-PROBEB Ml, POSITION, 3.03 l75OE-O 1,2.06848 1 E-0 1, 1 9039 1 1 E-O2 IJ41, S E , 1587480E-03,3.830390E-O2,7.417232E43 iJ41, CLIPART, cube3 iJ4 1, ROTATION, 1 J41, TYPE, PROBE iJ4 1, TEMPERA-, 0.000000E+00 iJ41, PRESSURE, O.O00000E+OO Ml, VELOCrTY, O.ûûûûûûE+ûO, O.ûûûûûûE+ûO, O . O O O E + O û

> OBJ42, NAME, R-PROBEM > OBJ42, POSKION, 3.03 l7SOE-O 1,2,06848 1E-O l,l.229576E-O 1 > OBJ42, SIZE, 1 .58748OE-O3,3.83O39OE-O2,l.S932 lïE-02 > OBJ42, CLIPART, cube3 > OBJ42, ROTATION, 1 > 08542, TYPE, PROBE > OBM, TEMPERATURE, 0 000000Etûû > OBJ42, PRESSURE, 0.000000E+ûû > OBJ42, VELCITY, O.ûûûûûûE+OO, O.OOOOQOE+OO, 0.000000E+00

> OBJ43, NAME, R-PROBE3' > 08543, POSITION, 3 .O3 1750E-0 1,2,06848 IE-O l,2,176963EQ 1 > OBJ43, SIZE, LS8748OE-û3,3.83O3WE-O2,1.3S l269E-02 > OBJ43, CWPART, cube3 > OB J43, ROTATION, 1 > OBJ43, TYPE, PROBE > OBJ43, TEMPERATURE, 0.000000E+ûû > OBJ43, PRESSURE, 0.00ûûûûE+ûû > OBJ43, VELOCITY, 0.000000E+00,0.000000E.t00,0.000000E+00

544, NAME, FRTPROBB ,544, POSITION, 1.3 19593E-0 1,6.508750E-02, 1.9039 1 1 E-O2 1544, S I E , 4.243 138E42, 1587503E-03,7.417232E-û3 IJ44, CLIPART, cube3 J44, ROTATION, 1 M4, TYPE, PROBE iJ44, TEMPERATURE, 0.000000E+ûû iJ44, PRESSURE, 0*000000E+00 iJ44, VELOCITY, 0*000000E+00, 0.000000E+00, 0.000000E+00

'45, NAME, FRTPROBT '45, POSITION, 1.3 19593E-0 1,6.508750E-02,2.176963E-O 1 '45, S E , 4.243 138E-02, 1.587503E-03,1.35 1269E-02 '45, CLIPART, cube3 '45, ROTATION, 1 !45, TYPE, PROBE 145, TEMPERATURE, 0.000000E+00 145, PRESSURE, OmOOOOOOE+OO 145, VELOCITY, 0.000ûûûE+00,0.OûûûûûE+00,O.OOOOOOE+00

> OBJ46, NAME, BOTPROBE > OBJ46, POSITION, 1.3 19593E-û i,2.O6848 1E-0 1,3.175005E-03 > OBJ46, SUE, 4.243 138E-02,3.83039OE-û2,3- l75OOSE-03 > OBJ46, CLIPART, cube3 > 08546, ROTATION, 1 > OBJ46, TYPE, PROBE > OBJ46, TEMPERATURE, O.ûûûûûûE+OO > OBJ46, PRESSURE, O,ûûOûûûE+OO > OBJ46, VELOClTY, 0.000000E+00,0.000000E+00,O.OQOOOOE+ûû

> 08547, NAME, F-BRICK > OBJ47, POSITION, 0.000000E+00,0,000ûûûE+ûû, 0.000000E+00 > OBJ47, S E , 3 .O63SOOE-O 1,6350000E-O2,3.048000E-0 1 > OBJ47, CLI~ART, cube6 > OBJ47, ROTATION, 1 > OBJ47, TYPE, BLOCKAGE > OBJ47, MA-, 145 > OBJ47, HEAT, 0.000000E+00

> 08548, NAME, B-BRICK

> OBJ48, POSITION, 0.000000E+0OT 3.88SOOOE-0 1 , 0.000000E+00 > OBJ48, SIZE, 3.0635ûûE-01,6.3499%E-02,3.048000E-01 > OBJ48, CLPART, CU&

> OBJ48, ROTATION, 1 > OBJ48, TYPE* BLOCKAGE > OBJ48, MATERIAL, 145 > OBJ48, HEAT, 0.000000E+00

> OBJ49, NAME, R-BRICK > OBJ49, POSITION, 3.063SOOE-O 1, O.OOOOOOE+OOT 0.000000E+00 > OBJ49, SIZE, 6.350000E42*4.520000E-0 1 , 3.M8OOOE-0 1 > OBJ49, CLIPART, cube6 > OBJ49, ROTATION, 1 > OBJ49, TYPE, BLOCKAGE > OBJ49, MATERiAL, 145 > OBJ49, =TT O-ûûûûûûEM

> OBJ50, NAME, BOTPLATE > OBJ50, POSITIONT O*OOOOOOE+OQT 6.350000E42T 0.000000E+00 > OBJSO, SIZE, 3.063SOOE-O 1,3.2500QOE-O l , O . O O E + 0 0 > OBJSO, CLIPART, cubet 1 > OBJ50, ROTATION, 1 > OBJSO* TYPE, USER,DEFINED

> OBJS 1, NAME, LEFWLAT > OBf 5 1, POSKION, 0.000000E+00,6.350000E-û2,0.000000E+00 > OB55 1, SEET O*000000E+OO,3.2SOOOOE-O I,3.O48WEa 1 > OB15 1, CLIPART, cubet 1 > OBJS 1 , ROTATION, 1 > OBJS 1, TYPE, USER-DEFINED

> OBJ52, NAME, RIGHTPLT > OBJ52, POSrrrON, 3.698500E-O 1, 0.000000E+OoT O.ooooOOE+W > OBJ52, SmT 0*000000E+00* 4520000E-0 l ,3.W8WEa 1 > OBJS2, CLIPART, cubet 1 > OBJ52, ROTATION, I > 0Bf52, TYPE, USER-DEFINED

> OBJ53, NAME, FRONTPLT > 08553, PûSlTïON, O.ûûûûûûE+ûû, O.OûûûûOE+OO, 0.0ûûûûûE+00 > OBJ53, SEE, 3.698SOE-û l,O.OOûûûûE+Oû, 3.M8ûûûE-û 1 > OBJ53, CLIPART, cubet I > OBJ53, ROTATION, 1 > 08553, TYPE* USER-DEFINED

> OBJ54, NAME, BACKPLAT > OB J54, POSITION, 0-000000E+00, 4XooOoE-O 1 , O1OOOOOOE+OO > OB J54, SEE, 3.6985ûûE-û l,O1OOûûûûE+OO, 3.M8OOOE-û 1 > OBJS4, CLIPGRT, cubetl > OBJS4, ROTATION, I > OBJ54, TYPE, USER-DEFINED

> OBJ55, NAME, FNTWALL > OBJ55, POSITION, O.ûûûûûûE+Oû, 6.35ûûûûE-û2,6.35ûûûûE-03 > OB JS5, SIZE, 3 .û635ûûE-O 1,3.175000E-03,2.9845ûûE-0 1 > OBJ55, W A R T , defautt > OBJSS, ROTATION, L > OBJSS, TYPE, BLOCKAGE > OBJS5, MA-, 1 11 > OBJSS, HEAT, 0.000000Em

> OBJS6, NAME, BACKWALL > OBJS6, POSKION, O.ûûûûûûE+W, 3.853250E-û 1,6.35ûûûûE-03 > OB 356, SIZE, 3 M3SOOE-O 1,3.175OOOE-03,2.984500E-0 1 > OBJ56, CLIPART, default > OBJ56, ROTATION, 1 > OB JS6, TYPE, BLOCKAGE > OBJS6, MATERIAL, 1 1 1 > OB J56, HEAT, 0.000000E+O

M STOP

Predicted Smelt Temperature Distribution at DiiPerent Depths

af'ter 3 Hours of Coolhg

Figure El. Pcedicted smelt temperatme distribution at the top surlace.

SIDEWALL (COOLED)

SIDEWALL (UNCOOLED)

SIDEWALL (COOLED)

SIDEWALL (UNCOOLED)

Appendbr G

Effeft of Parameters on the Smelt Cooiing Pmcess

TlME (MIN)

Figure G.I. Identicai nsults with constant pC, but dîtFeret p and Cp.

O 60 1 20 180 240 300 360

TlME (MIN)

120 180 240 TlME (MIN)

Figure GA Effkct of heat capacity.

CVV ' * ' I I I I 1 k I I I I

O 60 1 20 180 240 300 360

TIME (MIN)

Figure G.4. Effkct of iatent kt.

O 60 1 20 180 240 300 360 TlME (MIN)

Figure CS. Effect of low thermai conductivity.

Figute G.6. Edlèct of hi@ tbemd eoaductiviîy.

1 20 180 240 TlME (MIN)

Effect of heat transter eodîicknt h2 at ddincrent depths.

1 20 1 80 240 TlME (MIN)

b. Effkct of bat m e r coefficient h2 at 1s cm (6'3 belon the top surfa-

Figure G.7. Effet d heat transfer coefficient at the uncooled wiUo, h2.