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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
The author has granted a non- exclusive licence allowing the National Libracy of Canada to reproduce, Ioan, distriiute or dl copies of this thesis in mkofonn, paper or electronic fomuds.
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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.