.---i

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A TEMPEHATURE C('~JTROL STRATEGY FOR

STELCO McMASTER WORKS

by

Angelo M. Grandillo

A 'rhesis Submitted to the E'aculty of Graduate Studies and ReSearch in Partial Fulfillment of the Requirements

for the Degree of Master of Engineering

Department of Mining and Metallurgical Engineering Mc Gill Uni ver s i t Y Montreal, Canada

JANUARY 1988

t'

AC KNOW LED GEME NTS

1 wish to extend my sincere gratitude to Professor

F. Mucciardi, supervisor of this work and valued friend, for aIl

his efforts and help. l also wish to acknowledge the financial

contribution made by NSERC by way of an Industrial Postgraduate

Scholarship and the support recieved from management and

employees of Stelco, McMaster Works. Finally, l wish to thank

my mother for having given up her upstairs k i tchen table for the

past two years and having put up with me.

1

T A BLE o F CONTENTS

P':HJt'

Abstract.......... ..........•.......................... v Resume .••••••.••••..••••..••••.••••••••.••...••••.••••. Vil

List of Figures...................................... ... x Lis t 0 f Ta b 1 es. • • • . . • • • • . . • • • • . • . • • • • . . • • • • . . . • • • . • • • •• X '1 i

CHAPTER 1:

CHAPTER 2:

CHAPTER 3:

1.1

1.2

2.1

In trad u ct ion. . ......................•..

Ste1co McMa s ter Wor ks Steelmak i ng Facilities -Overview .•••••...•••.•.••..

Scope of Present Work •••••••.•••••••••.

The Continuous Casting Process ••••••••.

History of Continuous Casting ........ ..

L

l

4

7

9

2.2 Components of a Continuous Casting Machine •••••••••••••••••••• ., •• ~ •••••••• 15

2.2.1 The Tunùish ••....••.....•...........•.. 15 2.2.2 Tlle Mauld ..••....••....•••...•••...••.. 20 2.2.3 The Spray Zone......................... 24

2.3 Operating Aspects of Continuous Castillg .....•..... o ••••••••••••••• 2B

solidification and Structure of Con tin U 0 U S 1 y Ca s t Ste el B i 11 e t s • .. • • • • • 3 1

3.1 The Ct1ill Zone ..•••••...•••. , ••....••.. 33

3.2 InternaI Structure of Continuously Cast BIllets........................... 37

3.2.1 Mechanica1 Propertles of Steel at High Temperature....................... 37

3.2.2 Axial Segregation and "Mini-lngot" Formation •••••••••••••••••••••••••••••• 39

1 l

CHAPT ER 4:

CHAPT ER 5:

CHAPTER 6:

3.3 VarIables Influencing Cast Structure... 43

3.3.1 Machine DesIgn ....•.•..•.•..•••••...••• 45 3.3.2 Influence of SectIon Slze .••••.•....••• 46 3.3.3 Effect of Steel Composition ••••••••.••• 48 3.3.4 Influence of Castlng Temperature and

Flow Conditions in the Liquid Pool ••••• 50

Controlling Steel Temperature in the Ladle.............................. 59

4.1 Theoretical Simulations of Thermal Cycling of Ladies ••.••••••••••••••••••• 74

4.1.1 Procedure •••••••••••••••••••••••••••••• 75 4.1.2 Discussion and Results ••••••••••••••••• 79

4.2 Ladle Refining Furnaces ................ 104

Contro11ing Steel Temperature in the Tu nd i s h •••..... lit • • • • • • • • • • • • • • • • • •• 110

5.1 Flow Control Deviees ••••••••••••••••••• 112

5.2 Heating Steel in the Tundish Dur 1 ng Cast i ng ••.••..•.••...••..•..•••. 118

5.3 Scrap Additions to the Tundish to Reduce Superheat ..•••••••••••••••••. 127

5.3.1 Thermal and Klnet1c Conslderations ••••• 128 5.3.2 Plant Trials Conducted at

McMaster Works •.••••..••••••••••••••••• 133

Controll1ng Steel Temperature ln the Mould ...••....•.................••. 148

6.1 previous Work ••••••.••••••••••••••••••• 149

6.1.1 Work From USSR •.••••••••••••••••••••••• 151 6.1.2 Work From Ita1y •••••••••••••••••••••••• 155

6.2 Trial Work Performed at McMaster Works ..••••••••••••••••••••••• 159

6.2.1 Preparatory Stage •..•.••••••••.•••••••• 159 6.2.2 Experimental Stage ••••••••••••••••••••• 168

l l l

-i CHAPTER 7: Summary and Conclusions •••••••••••••••• 174

Appendix 1 ............................................. . 181

References ..•.•.••.•.•...••••..•••.•...............•...

lV

.­.

l\.BSTRACT

In thlS stuày, it was shown that Improvements in

the quality of continuously cast steel billets, similar to those

which can be achieved by e1ectromagnetic stirring (EMS) of the

liquid pool durlng solidifIcation, can be obtained if casting

superhe~ts can be consistently controlled at low levels. A

lack of casting temperature control is not only detrimentai ta

the quality of the cast prùduct, but also ta a shop's overall

productlvity. The important variables for temperature control

in the ladle, tundish and mould were quantified and possIble

methods of controlling these variables were proposed.

The thermal state of the ladle lining is one of

tbp major contributors ta the variabllity ln casting

temperature. This was shawn by performlng theoretical ladle

cycle slmulatlons of the Stelco McMaster Works 80 tonne ladIes

using a general heat and mass transfer computer software package

developed at McGill University, known as FASTP (Facility for the

Ana 1 ys ISO f Sys tem::; in Tr anspor t Phenomena). It was shown tha t

by reducing energy losses from the refractory ladie 1ining, by

way of Incorporatlng an Insulating refractory tile between the

ladle sheli and the safety linlng and by using a ladle lid

throughout the cycle of the lad le, te~perature lasses from the

liquid steel can be substantlally decreased. This can translate

lnto a decrease of heat ta heat temperature variabi1ity and the

v

1

..

....

model predicts that electric arc furnace tap temperatLlres can bl'

decreased by about 33 oC. Thus a significant improvement in

furnace prod uct i vi ty can be ach i eved •

Since, from an operating point of view, it is

preferable to cast steel on the hot side, it was determined in

th i s study that fu r ther temper a ture contro 1 can be ach i eved by

scrap cooling in the tundish. A theoretical kinetic analysis of

the McMaster Works tundish was performed, again using FJ\STP, and

it was determined that sufficient turbulence exists in the

tundish pourbox to melt scrap, l02mm X l02mm billet crops, or

sma11er bar mill crops, in order to further decrease superheat

and at the same time use excess energy which would have

otherwise been wasted to increase production. Trial heats

performed at McMaster Works showed that scrap additions to the

tundish are feasible and serve their purpose.

Liquid steel entering the mould always requires d

minimum of about lS oC superheat in arder for it to have

sufficient f1uidity to transfer from the tundish to the mould.

In order to help dissipate this superleat more quickly, this

study investlgated the feasibillty of feeding a high purity iron

powder into the mould during casting. Additions in the order of

1% were made and results supported work performed by other

researchers. Significant improvements in internal quality of

the cast product were observed •

VI

1 Darl'-, cetle étude, il (" .. ,t démontr<? p01Jr des billettes

pr()dlli tes par co',16e contin ;l', qlle des améliorations semblables

(\ Cl' 1 h·s r6alic;écs par bra~sage r-lectromagnétiqlle du liquide

en ~olidification peuvent être obten\les, si la slJrChallffe de

l'acit'r liquide' est contrôlée a bas niveau. Un manque de

contrôle dl' la température de Caillée affecte non seulement

la q11alité d" prOÙllit mais allssi la prodl1ctivité globale d'une

aciérie. Les variables importantes pOlIr le contrôle de la

tempéra turc: dans la poche de cO\ll~e, le panier réparti teur

L't le mOllle ont été quantifiées et des méthodes pour contrôler

ces variab les on t été proposées.

<1 La variabi lité de la tempéra ture de coulée es t dû en

"" majellre partie à l'état thermique de la brique réfractaire

dans la poche de co: lée. Cee i a é té démon tré par des

simulations théoriqlles basées sur les poches de 80 tonnes

de l'usine McHaster de la compagnie Stelco. Les sinlulations

on tété e ffec r-u6e s à l'aide dll logic ie l FAS TP (Faci li ty for

the Analysis of Systems in Transport Phenornena) développé

à l'l'niversiré }1cGill pOlir la simulation de transfert de

chaleur et de masse. Les résultats ont démontré que les

pertes de tempéra ture de l'acier liquide pe\lvent être

rédllites en diminuant les pertes thermiques du réfractaire

de poche à l'aide d'une tuile isolante entre la coquille

V 1 l

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1

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de la poche et le revêtement de' sécurilé. et d'un cOllvl'rch'"

posé sur la poche pendant le cyclL'. Il en n~sllltl' \me

ba is se de la va riab il i té de la tcmp0ra turc, dl' cOlllée L'Il

coulée. Les calculs théoriques prvdisent que la tcmp(>raturc

de vidange pOlir un four à arc plectriquc peut ~tre réduill'

d'environ 33°c. La conclusion révèle donc une augmentatioll

de la producti vi té du four à arc.

Puisque du point de vue opérationnel il est préférable

de couler l'acier à des températures plutôt chaudes, l'ptllde

a révélé que dans le contrôle de la température de coul('c,

des gains additionnels peuvent être obtenus par l'addition

de ferraille dans le panier répartiteur, pour refroidir

l'acier liquide. Une ana lyse théorique des paniers

répartiteurs de l'usine Nct-1aster fOt réalisé avec le

logiciel FASTP. Les résultats ont démontré qu'il exis te

suffisanunent de turbulence dans le panier répartiteur

pour fondre des morceaux de bi Ile t tes de 102 rrnn x 102 mm

ou des morceaux de barres provenant du laminoir. Cette

technique perme t de réduire la tempÉ'ra turc de surchau f f(>

en utilisant l'énergie excédentaire qui autrement aurait

été perdue. Des expériences rÂalisées à l'usine McMaster

ont révélé que l'addition de ferraille dans le panier

répartiteur est faisable.

Vl11

1

L'acil'r liqlIide sortant dt! panier répartiteur nécessite

'Inl' <""lrcha1!ffp minimum de lS"C pOlIr maintenir lme fluidité

'-il!ffisanLe pOlIr passer dlI panier au moule. Pour aider à

dis<.,ipcr cettt' slJrchauffe pJus rapidement, une étude de

[ai~ahiliU~ a évaluer la possihilité d'ajouter une poudre

de ft' r plJr dans le mOllie dl Iran t la c01l1ée. Des additions

de l'ordre dt' 1010 ont ét0 faites pendant la coulée continue.

Ll'S résultats confirment le travail d'autres chercheurs et des

améliorations considérables de la qualité interne des billettes

ont été ob'-iervées.

lX

•

..

LIST OF FIGURES (reE) -----------------------

Figure 1.1: Fishbone diagram of variables afEecting quality and productivity.

Figure 2.1: Basic elements of a continuous casting machine. (4)

Figure 2.2: Diagram of Bessemer process [or continuous casting. (5)

Figure 2.3: Four basic caster designs. (4)

Figure 2.4: Wor1dwide production of crude steel and share of continuous casting. (7)

Figure 2.5: Continuous casting billet machines in operation. (7)

Figure 2.6: Bloom machines / Strands in operation. (7)

Figure 2.7: Slab machines 1 Strands in operation. (7)

Figure 2.8: Relation between mean number of inclusions (>100 microns) in slabs and flow control. (8)

Figure 2.9: Distribution of metallurgical Eunctions between furnace, ladle and tundish. (17.1)

Figure 2.10: Transverse section of (a) billet and (h) slab showing formation of air gap. (18.2)

Figure 2.11: Billet defects. (18.3)

Figure 3.1: InternaI structure of continuously cast steel billets.

Figure 3.2: Transverse cut of AISI Grade 5160 H (low casting superheat) •

x

Figure 3.3:

Figure 3.4:

Figure 3.5:

Figure 3.6:

Figure 3.7:

Figure 3.8:

Figure 3.9:

Trandverse cut of AISI Grade 5160 H (high casting sl.lperheat) •

% Equiaxed vs. Index of Center Segregation for several shapes and sizes. (25.1)

Mini-in90ts formation model. (24)

Core segregation in high-carbon spring wire resulting in breakage during drawing operation.

Quarter point inclusions.

Relation between segregation index of C and segregation index of P, S, Mn. (26)

Morpho10gy of solid/liquid interface as a function of temperature grad ient (G), rate of solidifica ... ion (R) and undercooling (AT). (2J)

Figure 3.10: Histogram of center segregation ratio of C for various casting conditions. (26)

Flgure 3.11: Effect of steel superheat on the relative size of the equiaxed zone with and without EMS. (From IRSID)

Figure 3.12: Effect of casting temperature on the size of the equiaxed zone and on the severity of axial segregation. (27)

Figure 4.1: Distribution of average steel temperature in the tundish for AISI Grade 5160 H.

Figure 4.2: Distribution of steel temperature 1055 between the ladle and the tundish for AISI Gr~de 5160 H.

Figure 4.3: Distribution of first steel temperature measured in the ladle for AISI Grade 5160 H.

Figure 4.4: Regression analysis of ladle vs. tundish temperature.

><1

Figure 4.5:

Figure 4.6:

Figure 4.7:

Figure 4.8:

Figure 4.9:

Figure 4.10:

Effect of argon bubbling in the ladle on steel temperature profile in the tundish. (32)

Diagram of energy flow in the ladle for steel casting. (33)

The effect of different refractory types on the ladle cooling behavior with a preheat of 400 0 C and a furnace tap temperature of 1650 oC. (33)

Lining configuration and thermal properties used in the simulation. l-Regular Lining 2-Insulated Li n i ng.

Simulated ladle cycle.

Energy balance equating temperature lost by liquid steel to energy gained by lining and energy lost to the surroundings.

Figure 4.11: predicted progression of cold-face temperature rise for regular and insulated ladIes during preheating of newly lined ladIes.

Figure 4.12: predicted progression of hot-face temperature rise for regular and insulated ladIes during preheating of newly lined ladIes.

Figure 4.13: Energy distribution after preheating of newly lined ladIes.

Figure 4.14: Increase in lining energy content during the first 5 heats on a regular ladle. (liquid steel/lining contact time = 130 minutes)

Figure 4.15: Increase in lining energy content during the first 5 heats on an insulated ladle. (liquid steel/lining contact time = 130 minutes)

Figure 4.16: Comparison of energy content of regular and insulated linings after working temperature has been reached, as a function of contact time t.

Xl l

Figure 4.17: Temperature profile through a regular lining at the end of each step in a cycle after the ladle has reached working temperature.

Figure 4.18:

Figure 4.19:

Figure 4.20:

Figure 4.21:

Figure 4.22:

Temperature profile through an insulated lining at the end of each step in a cycle after the ladle has reached working temperature.

simulation results of equivalent liquid steel temperature loss (oC) for a regular ladle versus 5n insulated ladle.

simulation results of equivalent liquid steel temperature loss (oC) for an insulated ladle having a captive lido

Final manganese distribution for LF practice at U.S. Steel, Fairfield Works.

Final Carbon distribution for LF practice at U.S. Steel, Fairfield Works.

Figure 4.23: Within-heat temperature 1055 distribution in the tundish with LF practice at U.S. Steel, Fairfield Works.

Figure 5.1:

Figure 5.2:

Figure 5.3:

Figure 5.4:

Figure 5.5:

Example of a typical temperature profile in the McMaster Works Tundish.

McMaster Works tundish.

Fuil-scale water model of the McMaster Works tundish.

Flow patterns in a tundish with no flow control devices incorporated.

Flow patterns in a tundish with optimum flow control configuration.

Xlll

Figure 5.6: Schematic of heating system for molten steel in the tundish during continuous casting. (42)

Figure 5.7: Temperature variations of steel in the tundish for conventional casting practice. (42)

Figure 5.8: Temperature variations of steel in the tundish for experimental heats using the heating system. (42)

Figure 5.9: Effect of using the heating system on first and last slab reject frequency. (42)

Figure 5.10: Enthalpy balance for determining amount of scrap addition required to drop liquid steel superheat by the desired amount.

Figure 5.11: Oimensionless relationship for melting kinetics of steel cylinders immersed into liquid steel at 1570 0 C. Bath liquid steel and cylinders are assumed to have a liquidus temperature of l53S oC and a solidus temperature of 1490 oC.

Figure 5.12: Regression analysis for determining the relationship between Bi and Fo fOl total time required ta completely melt the cylinders.

Figure 5.13: Test cylinder after a 5 second immersion.

Figure 5.14: Experimental data points as they appear on the diroensionless relationship established in Figure 5.11.

Figure 5.15: Tundish temperature profile for trial heat #23875.

Figure 5.16: Tundish temperature profile for trial heat #23982.

Figure 5.17: Tundish temperature profile for trial heat #24329.

Figure 6.1: Sample 23-Transverse section, no iron powder (46) Sample 24-Longitudinal section, no iron powder Sample 25-Transverse section, with iron powder Sample 26-Longitudinal section, with iron powder

XlV

Figure 6.2:

Figure 6.3:

Figure 6.4:

Figure 6.5:

Figure 6.6:

Figure 6.7:

Figure 6.8:

Iron powder addition method using a carrier gas. (46 )

Schematic drawing of metal powder feeding apparatus. (49)

Schema tic of powder feeding system developed for trials at McMaster Works.

Heat #33241 control billet sample macro-etch.

Heat #33241 trial billet sample macro-etch.

Heat #33291 control billet sample macro-etch.

Heat #33291 trial billet sample macro-etch.

xv

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Table 1.1:

Table 1.2:

Table 3.1:

Table 3.2:

Table 4.1:

Table 4.2:

Table 4.3:

Table 4.4:

Table 5.1:

Table 5.2:

Table 5.3:

LIS T 0 F T A BLE S

McMaster Works electric arc furnace characteristics.

McMaster Works caster characteristics.

Effect of carbon content on various aspects of solidification and cast structure.

Chemical composition and casting conditions of test billets.

Implications associated with the lack of control of liquid steel superheat in the tundish when using metering nozzles as a means of steel (low control.

Parameter table and ANOVA table for regression analysis.

EAF temperature compensation required for heats prior to attaining working temperature of the lining.

Comparison of maximum sheli temperature (oC) for regular and insulated ladIes; predicted results and actual plant data.

Results of water model analysis showing retention time and volume fraction comparisons for tundish without flow control devices versus tundish with optimum flow control configuration.

Composition of test cylinders, liquid steel and bath conditions.

Results of cylinder immersion tests used for determination of h. (d(initial)=23.9mm)

XVl

Table 5.4:

Table 6.1:

Table 6.2:

Table 6.3:

Condition of trial heats for scrap additions in the tundish.

Increase in billet production cost ($ per tonne) as a function of amount of powder added (shown in weight %) and price of iron powder.

properties of ATOMET 28 and ATOMET 602 iron powders (as per QMP catalogue) •

Chemistry and casting parameters for iron powder addition trial heats.

XV 11

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PM

CHA PTE R 1

Introducllon

1.1 Stelco McMaster Works Steelmaking Facilities Overview

The work presented in this thesis was carried out

at Stelco McMaster Works, Steelmaking Department. McMaster

Works is located in Contrecoeur, Quebec; about 40 km north-east

of Montreal. Steelmaking facilities were commissioned in 1974 by

the consulting firm Hatch Associates. Originally deslgned for n

nominal capacity of 175,000 tonnes per year (100% merchant

quality) , continuous improvements to procedures and equipment,

allowed production to top 300,000 tonnes ln 1985, of which, 80%

was special bar quality (SBQ) mainly for the automotive

inèustry.

The present installation consists of an electrlc

arc furnace and a 4-strand billet caster. The furnace melt

capacity is 80 tonnes. Power is delivered by a three phase

transformer of 50 MVA capacity. Three oxy-fuel burners rated at

4.5 MW each are also used during melting. This combination

gives an average tap-to-tap time of 120 minutes (including

maintenance and breakdown delays). Table 1.1 shows sorne

characteristics of the furnace.

When a heat has reached the required temperature

t and carbon level in the furnace, it is tapped into a pre-heated,

dolomite lined ladle where carbon and ferro-alloys are added as

necessary. Ladle treatment of the steel consists of argon or

nitrogen stirring through a porous plug and if required,

elements such as silicon, manganese and carbon can be injected

in the form of a cored wire.

The heat is then delivered to the continuous

casting department. The casting machine, de1ivered by

'Concast', is of the curved type with a radius of 7.92 m. From

the ladle, steel is teemed into a tundish of Il.0 tonne

capacity. From the tundish, steel is distributed to 4 strands.

Table 1.2 gives sorne characteristics of the casting machine.

~bout one-third of the steel produced at McMaster

Works is AISI GRADE 5160 used in the fabrication of spring flats

for the automotive industry. The rernaining two-thirds of the

production has a variety of other applications ranging from

rounds for automotive forgings to wire for nail production, to

reinforcing bars.

2

Table 1.1: McMaster Works E1ectric ~rc Futnil~e Characteristics

1. She11 Diameter (m)

2. Transformer Capacity (MVA)

3. Maximum Power Input (MW) rncluding Oxy-Fuel Burners

4. TOp or Door Charged

5. Electrode Size (mm)

6. Shell and roof

7. product Mix: % Heats - Carbon % Heats - Alloy

Table 1.2: McMaster Works Caster Characteristics

Il. Manufacturer

Type 2.

3 • Curve radius (m)

4 • Number of strands

5. sequence Casting

16

• Tundish Capacity ( tonnes)

7. Sizes of Squares Cast (mm)

8. Mou1d length (mm)

9. Spray Zone Length (m)

10. Typica1 Casting speeds (m/m i n) 89 mm

102 mm 152 mm

----------~-_ .... _.- --

3

50

40 53.5

top

508

water-coo1ed

65 35

CONCAST

Curved

7.92

4

Yes

11.0

89, 102,

813

1.5

3.3 2.8 1.3

152

1.2 Scope of Present Work

For any continuous casting operation to be

competitive in today's markets it must be able to meet the

meticulous quality demands imposed by customers while keeping

production costs down. It used to be that quality and

productivity were thought to be two counteracting forces and one

could only be achieved at the expense of the other. In today's

environment, steelmakers are beginning to reject this

philosophy. They now realize that quality and productivity

really mean the same thing and one can't be obtained without the

other. By adopting concepts such as "Statistical Process

Control", steelmakers realize that in order to consistently

obtain high quality, procebses must operate in statistical

control. 1 ,2,3

~s can be seen in Figure 1.1, the three main

parameters characterizing the condition of steel during

steelmaking are temperature, composition, and fluid flow. The

extent to which these three parameters can be controlled wi:l

ultimately deterrnjne the consistency of the response parametersi

namely quality and productivity.

4

Ul

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TEMPERATURE FLUID FLOW

COlvlPOSITION

Fl.gure 1.1: Fl.shbone dl.agrarn of varlables affectlng qua1ity and productl.vl.ty.

QUALITY PRODUCTIVITY

The focus of the work presented in the chapters to

follow will be on steel temperature and flow in the liquid pool

of the solidifying strand. The main objectives of the work can

be summarized as follows:

1) to presen t a comprehens ive litera ture rev iew in

order to show how the quality of continuously cast

billets is affected by steel superheat and other

operating variables. This will, whenever possible, be

supported by actual plant data from McMaster Works.

2) to show, by means 0 f a compu ter s irnula t ion of the

thermal cycles of McMaster Works ladles, what action

can be taken to reduce overall temperature variability,

hence reduce tap temperatures in the EAF.

3) to present viable methods of further reducing

temperature variability in the tundish during

continuous casting.

4) to discuss and present results of plant trials

performed in order to determine the effects of adàing a

high purity iron powder to the mould during continuous

casting as a means of providing supplemental cooling to

the liquid pool of the solidifying strand.

6

•

CHA PTE R 2

The Continuous Casting Process

The development of continuous casting provides the

stee1maker with an economically attractive alternative ta

conventiona1 ingot casting. Since the end product of the

continuous casting operation i5 a 5emi-fini5hed shape, the cost

intensive prirnary r011ing stage required for in00t processing i5

e1iminated. Improvement5 in yield, surface and internaI qu~lity

are among sorne of the advantages continuous casting offers over

ingot casting.

Figure 2.1 shows the basic e1ernents of a

continuous casting machine. Steel supplied from the melt shop

via a lad1e, i5 teem0d into a tundish which in turn feeds one or

several water-cooled copper moulds. As the solidifying strand

exits the mould it enters a spray cooling zone. The cast

section is then cut to 1ength thus giving a semi-finished

product ready for subsequent processing.

In this chapter a brief history of the development

of continuous casting will be given and the basic components of

the casting machine and the casting process will be described.

7

OJ

l - Ladle 2 - TundJ.sh 3 - Mould 4 - Spray Cool~ng zone 5 - Stralghtener 6 - Cuttlng Torches 7 - Electro-Magnetic Stirring

~ ... w .... ~ ... w

::.::.:.~ :.~':":::':.::"::}f.'::::'::~::~:-:;':::'~.{~;':~';':~;'$;':~;:i:;':5;C''':~;':~:?5:t:f'::';':)i:,};:.JJ}J IJ.~~Ht~.:~:: !-,,'.;. '-': .:~ •. ~:-••• ;' ••• '.' '.' .....: .. :

Figure 2.1: Basic e1ements of a continuous casting machine. (4)

." l 1

•

j 2.1 History of Continuous Casting

Continuous casting of metals has a history going

bac\{ to before 1840 when a patent was granted to George Sellers

who developed a machine for continuously casting lead pipe. 5

Sir Henry Bessemer patented his idea of continuous casting of

steel plates in 1846 and ran plant trials in the 1890' s.

Figure 2.2 shows a diagram of the Bessemer process.

The first commercial casting machine was developed

for continuous casting of brass billets in the early 1930'5 by

s. Junghans in Germany. He also introduced the oscillating

mould system des igned for the purpose of avo id i ng st i ck ing 0 f

the casting to the mould.

In the 1930's Al1egheny-Ludlum began development

of a continuous casting machine for steel using Junghans'

oscillating mould. It wasn't until the early 1950'5 though that

Allegheny-Ludlum and others were successful on a pi lot plant

scale. By 1960, there were two commercial production machines

in operation. It' s only been in the past 15 years that

cont i nuous cast i ng of steel has und erg one i ts mos t rap id 9 rowth.

The development of mater ials and processes and the advent of

automation and computer control has given continuous casting

significant advantages over conventional ingot casting; namely

superior quality, higher yield and higher production rates. The

9

r

.,.,.,A' ROLLS

/ ';,. éf~

~ o

. , , '~\I '~

SOL.IU ME1AL--'

1 . Il' , r'f'l '

cJ~

\ o

WATE" SPRAYS

Figure 2.2: Diagram of Bessemer process for continuous casting. (5)

t 0

• l

overall reduction in steel production costs associated with

continuous casting, when compared ta ingot casting, makes it

extremely difficult for ingot casting operations to remain

viable in the near future.

Figure 2.3 shows the four basic caster designs.

The first machines built were of the vertical type. Curved

machines were then developed to minimize the height of

installations and to allow existing plants to incorporate

continuous casting without having to modify crane height. This

trend to cornpactness and reduction in height is today being

taken to an extreme with the develaprnent of horizontal

con t i nuous cast i ng. 6

Figure 2.4 shows how continuous casting has grown

over the years. More than 40% of today' s total steel production

is continuously cast. Figures 2.5 to 2.7 show how drarnatically

the number of casting machines has increased in recent years.

1 t

-~ ct 20 .... .... ct :;; 15 z LI.

o la l­x g 5 ::t:

Il

Il

$$­-$$-

-.-, $ $- Il

l}~ 4-1/ $­Il ;.

;/' ~-$ O~~ ____ ~~~_~~ __ ~ __ __

VERTICAL ST. w /BENDING CURVED

Figure 2.3: Four basic caster designs. (4)

\2

BOWED

•

1.000 :E: 900 WORLD StEEL PROIJUCliOIi :::> ~ 800 Z WOAlD CII\JOf IIlIl PMX>UCI"'" c:( ,

700 \ , , 0::: w

600 .-' a.. (%) ./ t/) 500 z

SIIARE ΠCC-PRODUCtiON ..-liilRLn C~-:= qOO 1ii'0' DI' ec--..cllON 1%1 ,," PROUUC 1 ON z

\ /" ~D C('POOOUCIION 0 300 "" , ...J 200 .,.~ ..... ' 1 ..-'-...J .......... 1 _._ ...... -:E: 100 .' .J~ • .,. .... -

._----.::----_ .. --------0 ........ -.- --_ ... '65 '70 ']'; '80 '85

(ALE~DfIR )fAR CONCAST UNION A.G. WORLD (ASTER S RVEY 198Q.1.

WORLDWIDE PRODUCllON or (RUVE SJEI:l AND FIGuP~ 2.4

1800

~1600 z: ~ 1QOO l-V) 1200 ."

1/) 1000 UJ z: ;: 800 u ~ 600

~ ~OO

~ 200 z:

o

",.ettlMfS

".AHDS

rrfufu if "f

SUARE OF CONTINUOUS CASTING (7)

--f- f- f- t-

- -- - - - -r- t- I- :--- - - -

f- - - - - - - - -

1-- - - - r- f- - - -- r-

t- - f-

1 r if " ,

1 ,

1968 69 70 71 72 73 7Q 75 76 77 78 79 80 81 82 83 814 YEAR

CONCAST SERVICE UNION A,G •• WORLD CASIER SURVEY 198~-1

'IC:C~r: 2.5 COtHIHUOUS CASlIHG BillET MACHINEs

IN OPERATION (7)

13

50 q5 qO :55 1-:50 z

lLJ

25 u ~ UJ D-

20 J5 10 5 0

1

2

900

BOO

700 -r-en c:J z: 600 --t--r-CI; 0:: ~

1- t- :- r-Vl 500 0lil

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400 - 1- t- r- t- r- f--z: - 300 ::r: f- f- i- - 1- -l-L) M'CHINtl

~ ""AHol

200 1- !- 1- - r-~

n If ~ ..; 100 t- c- t- 1 II ~ rf

- r-0

1 fJ 1 r r r r z:: 0 1968 69 70 71 72 73 14 75 76 77 78 79 80 81 82 83 84

YEAR CONCAST SERVICE 11H10 A.G WORI D CASTERSlIRYllIQR4-1

BLOOH t\ACHtNES 1 ST RANDS IK OPERM ION 3 fo'ICùRL' 2.6 (7)

450

Vl 400 Cl

~ 350 0::

t;; 300 0lil 250 Vl

~ 200 ::r: "ltHlH(~ '-' 150 "1 '''IAHOS

r t-~ ~

100 ~ 1 fl . 50 r r ..., ri 1111111 1 0 z:: 0

1968 69 70 71 72 73 74 75 76 71 78 79 80 81 82 83 84 YEAR

COHCAST SERVICE UNiOn. WORlD CASTER SURVEY. 1984-1.

SLAB MACHINES 1 ST RANDS IN OPERATION ~ICCRL 2.7

( 7) q

\4

2.2 Components of a Continuous Casting Machine

Casting machines may differ in size, capacity,

shape, sizes and shapes of sections cast, number of strands and

other design characteristics but they aIl have the following

components which serve a common purpose.

2.2.1 The Tund i sh

In a continuous casting operation the primary

function of the tundish is to distribute a controlled amount of

mol ten steel from the ladle to the mould or moulds. 8 The

tundish is also counted upon to perform a variety of other

functions which are now summarized.

1) Casting Speed Control

Casting speed i5 determined by the rate at which

steel is supplied from the tundish to the moulds. In practice

this is done in one of two ways. The traditional method makes

use of metering nozzles and steel level in the tundish to

control speed. Thus this method is based on operating with a

constant nazzle aperture and varying the steel level to yield

the required steel supply ta the moulds. This method is

preferred due to the relatively low costs in both nozzle

installation and maintenance when compared to other available

15

methods. The major disadvantages of this system are in its

ineffectiveness for accurate speed control due to the long

response time of the system and the high risk of slag

entrainment caused by vortexing, especial1y when casting at

lower speeds (i. e. lower tund ish levels) •

The more modern casters are equipped with

slidegates, stopper rods or sometimes both. A1though these

systems are much more expensive to instal1 and maintain, they

allow a full tundish level to be malntained throughout a cast.

Casting speed can be changed instantaneous1y by changing the

effective nozz1e aperture.

2) Buffer for Sequence Casting

Sequence casting (i.e. casting heats one after

another wi thout interrupti ng the steel supp1y to the mou1ds) is

essential in order to reduce production costs and improve steel

quality. Production cost reductions are achieved by the

resu1ting increase in billet yield, saving on tundish board

refractory by having more than one heat cast in the same tundish

and a reduction in operating personnel since less tundishes have

to be prepared and less caster set-ups are required. Qua1ity is

improved because the casting machine is allowed to run at

"steady-state" for longer periods of time.

16

3) Removal of Non-Metallic Inclusion

Much work has been done and is being done on

optimizing tundish design i.e. size, shape, etc. and

incorporating flow control devices aimed at increasing sleel

residence time in the tundish 50 as to enhance inclusion

flotation. 8 ,9,lO,11 A good example of the results which can be

achieved by flow control devices is seen in Figure 2.8.

4) Chemistry and Inclusion Shape Control

Traditionally steel chemistry control has been

achieved by alloy additions (trimming) and inclusion shape

control by additions of rare earths or calcium in the ladle (ie:

ladle metallurgy) .12,13,14 The developers of ladle metallurgy

technology have extended this to tundish metallurgy. As

tundishes are becoming larger and larger, adjustments ta steel

chemistry and inclusion morphology through tundish additions are

becoming more common. 1S ,16 Figure 2.9 shows how the different

metallurgical treatments are distributed between the furnace,

ladle dnd tundish.

17

- - -----------

Tor VI Eh' SI CE VInl

~ Il' ~ ( ...

0 ~ , A

' ...... _--

il ~ ----: 0 B ---1

U>S § --- /J 0 c --- '--

... :-

UJ ..... 2.0 c

a ..... Ul :J

r-1

U c ..... 1.5

\4-

a c... ID

1.0 Ll E :J C

C ro ru 0.5 ~

r-

0.0 nr-None A 8 C

Types of dams

Figure 2.8: Relation between mean number of inclusions (> 100 m iCrOn5) in 51 abs and flow control. (8)

18

fr"~

,

<0 neltin<::;

OXldatlon

:'1<1 Jor 1\1101' l ng

Figure 2.9:

Denassli1C:;

1'.11 O~."l ne)

DeOXlàa tlon

Temp. AdJu~tment

Desulfurization

HOl'1oqenlzatlon

Graln Slze

Cleanllness

Distribution of metallurgical functions between furnace, ladle and tundish. (17.1)

......... ~

Deox. .1\.d lUS t

]\110:." Trl['1;,lln rj

Desul"urlzatlC'n

Cleanllnes~

C:rain ~lze

".

2.2.2 The Mould

The primary function of the mould in a continuous

casting operation is to extract sufficient heat from the steel

50 tha t the sol id if ied shell a t the mould ed t i s thick enough

and strong enough to hold the liquid steel without bulging or

breaking. The design and operation of the mould and rnould

assembly have a profound effect on the heat transfer rnechanisrns

within the rnould. This in turn deterrnines the quality of the

bill et sur face and inter i or.

The mould, usually made of pure copper or copper

alloy, is cooled from the outside by water flowing through a

channel. The major concern in this method of heat extraction is

that thermal cycling of lhe mould due to nucleate boiling, can

develop locally just below the meniscus (i.e. the steel level in

the rnould) as described hy Brimacombe ane Samarasekera. l8 • 1

This phenomenon causes a drastic decrease in the local heat

transfer coefficient thus allowing the mould to heat up to

temperatures hot enough to soften the copper. This gives rise

ta permanent local distortion of the rnould. There is evidence

which shows that this severely affects billet quality~ namely in

the occurrence of corner cracking and rhomboidity of billets.

The variables which have been found to have the most affect on

mould wall temperature are:

20

•

1 l)cooling water velocity

2)scale formation on the water/mouid interface due to the deposition vf mineraI salts

3)mould wall thickness

4)steel carbon content

On the hot surface of the mould (i.e. the

mould/steel interface) the heat transfer rate is primarily

determined by the air gap formed due to the volumetric

contraction of steel as it undergoes solidification and other

phase changes. The air gap and its evolution, as described by

Samarasekera and Brimacombe,l8.2 provides the largest

resistance to heat flow (about 84% of the total) in the system.

Air gap formation is the subject of much investigation and has

proven to be extremely difficult to model mathematically. Much

work is being done in the area of mould design, specifically in

the design of mould taper. Taper in the mould is incorporated

to compensate for the shrinkage gap so as to allow a more

intimate mould/steel contact. Consequently, more heat

extraction in the mould is possible. Figure 2.10 shows how the

air gap evolves. Since the corners of the solidifying sections

have heat extracted from two directions, due to the nature of

the geometry, they will be the first to shrink and lose contact

with the mould. Even though the faces solidify, the ferrostatic

pressure from above is enough to bu1ge the faces 50 that contact

with the rnould is maintained for a longer period of time. This

condition is referred to as reentrant corners. The use of

multiple taper moùlds, having a strong taper at the meniscus

21

<f

io

(a) mould 1--- air gap

solld stQQI }

Ilquld staal

1 ~ -J

(b)

Jf

--4- --

"L J

Figure 2.10: Transverse section of (a) billet and (b) slab showing formation of air gap. (18.2)

22

•

level, has been found to alleviate this condition. These types

of moulds allow for more intimate contact between the steel

shell and mou Id by compensating for the shrinkage of the shell.

One other important aspect related to the

operation of the mould is lubrication. There are two media for

lubrication used in industry to insure that sticking between the

mould and the steel is minimized. They are oils and powders.

Synthetic oils or rapeseed oil are used in billet casting. They

wet the mould. In casting large sections, mould powders are

used. These powders melt at steelmaking temperatures and wet

the steel. The advantage of using powders is that submerged

nozzles can be used thus p~otecting the stream and exposed steel

surface from re-oxidation. Lubricants and oscillation of the

mould l9 ,20 have made it possible to continuously cast steel at

rates high enough to make the operation more profitable than

conventional ingot casting.

23

2.2.3 The Spray Zone

~ The sprdy zone is located irnrnediately below the

mould. It consists of a series of nozzles attached to vertical

headers or risers delivering water ta the individual nozzles.

The function of the spray zone is to continue to extract heat

from the solidifying sectIon after it exits the mould. At the

mould exit, the thickness of the solidified shell is typically

in the range of 9 to 12.5 mm for billets and 16 to over 20 mm

for blooms and slabs. The thickness basically depends on the

casting speed (i.e. the residence time of the cast section

inside the mould) •

The thermal properties of steel make it such that

once a shell in the mentioned size range is formed, the shell

acts as an important thermal resistance. If the cooling rate at

the surface of the solidifying section is too high, the

solidification rate doesn't increase. Instead the surface is

overcooled. This produces a high temperature gradient in the

solidified shell resulting in the presence of thermal stresses

within the shell. Thermal stresses also corne about when the

spray zone is physically too short, when there's a sudden

reduction in heat transfer coefticient, as is the case when

nozzles are clogged, and/or the steel superheat is too high.

According to Brimacombe et al.,IS.3 these thermal stresses

cause the steel surface to reheat from the inside allowing the

24

•

i she1l to expand relative to the interior. Thus tensile strains

at the solidification front are created. Cracking will result

if local strains exceed the critical 0.2 to 0.3%. It has been

suggested that lOOoe should be the maximum surface reheat

acceptable. In billets, these cracks, known as midway cracks,

are the most commonly found defects.

Another billet defect, whose origin is believed ta

be in the mould, but is accentuated in the spray zone, is

rhomboidity, a1so known as off-squaredness. Rhomboidity arises

when the billet is nol symmetrically cooled. If for sorne redson

(ex. plugged nozzle(s) near the mould exit) the billet is

unsymmetrical1y coo1ed, diagonal tensi1e strains can be

generated yie1ding a billet with two diagona1ly opposed obtuse

corners and the two other diagonal corners being acute.

Diagonal cracks are often found in the opposing obtuse corners.

Figure 2.11 shows the billet defects just described •

From this discussion it is apparent that the spray

zone plays an important role in the production of quality steel

in the continuous casting process. In the early days of

continuous casting, spray zones were designed with two basic

characteristics; high heat extraction in arder ) maximize

productivity and simplicity 50 that they could be easily

maintained. As steel quality became increasingly important, a

re-thinking of spray zone design was required.

25

r

Figure 2.11: Billet defects. (18.3)

1 l' / - -- -

Billet

(1) Midway cracks (2) Rhomboldlty / diagonal

cracks

26

1

1

Today's approach in designing a spray 7.one, as

outlined by Brimacombe et al.,IS.3 is first to calculate the

optimum steel section temperature profile within the spray zone,

50 as to avoid excessive thermal stresses. In the calculation

for the design of a billet caster spray zone, usually performed

using a numerical solution to solve the unsteady heat conduction

equation, the boundary condition imposed to the billet surface

is one of fixed temperature (1140 - l150 0 C) for a spray

distance of about 3.5m. A heat transfer coefficient, as a

function of distance below the exit of the mould, required ta

maintain the fixed temperature condition i5 then calculated.

This 'required' heat transfer coefficient is then decamposed

into its radiative component and its spray component. The spray

component is then used to calculate a water flow density

(usua11y expressed in 11m2 5) with the aid of empirica1

relations. Final1y a spray nozzle assemb1y which can supply the

required water f10w density is arrived at.

The following approach is a valid one if the

caster always operates at the conditions imposed in the design

calculation. Unsteady state operations however do arise and

more often than not cause most of the quality problems. It is

27

.,

for this reason that the described approach to spray zone design

is today taken one step further; computer controlled spray

zones. with computer control, caster conditions are

continuously monitored and as conditions change, adjustments to

the spray zone are made automatically. This insures that the

prescribed boundary surface conditions are kept under the best

control possible.

2.3 Operating Aspects of Continuous Casting

The success of a casting machine in yielding a

quality product greatly depends on the consistency of

performance of each individual component. Misalignment and wear

of components, if undetected by the caster operating crew, can

make the difference between a quality steel heat and a

downgraded heat.

prior to the start of casting of a heat, several

verifications must be carried out by the operating crew. Apart

from insuring that aIl hydraulic and electrical systems are

fully functional, the following checks are usually performed

during each set-up, prior to the start of each casting sequence:

l) tund i sh noz zle al ignment - mi saI ignment could ca use breakouts or a double skin (scabby) condition

28

•

1

J

---- -------------------------

2) regular or irregular mould wear - can usually be seen as 10ca1ized wear of the mould plating material

3) 1ubricating oi1 supp1y uniformity

4) spray zone clogged or misa1igned nozzle - must be replaced and/or aligned as soon as possible so as to reduce the risk of unsymmetric cooling

5) skid pad free of sku1ls and properly a1igned

6) straightener roll pressure

7) billet length sensor

8) dummy bar positioning

9) proper working order of emergency 1aunders - used to deviate steel away from the mou1d into a scrap bucket in case of breakout or mou1d overf10w

10) mou1d level detectors

11) mould oscillating system

12) water flow and pressure to the mou1ds and spray zones

When particu1ar qua1ity prob1ems are detected

during routine billet sample inspectjon, the operating crew is

notified and action can be taken in arder to prevent the problem

from reoccurring.

In order to begin casting a heat, a dummy bar must

be positioned and chi11ers p1aced in the mould. Once the

tundish has reached the required mo1ten steel level, an oxygen

lance is used to burn the copper plug sea1ing the metering

nozz1es and steel begins to flow into the moulds. When the

steel reaches a predetermined 1eve1 in the mould, strand

withdrawa1 begins. As the dummy bar c1ears the straightening

ro11s, the torch burners are activated and the dummy bar is

separated from the cast steel section. As the cast section

29

travels down the skid, and hits the length sensor, the torches

are again activated. The sections are then sent to a cooling

bed where they can be subsequently disposed of as required.

The steel temperature in the tundish is monitored

during the cast by periodicaliy taking a temperature measurement

using expendable thermocouples. Casting speed is directly

determined by the temperature. The hotter the steel, the slower

the casting speed must be in order to reduce the risk of

breakout. Quality is aiso adversely affected by steel of high

superheat. If the temperature is too cold the resuit could be

premature termination of a cast.

In the continuous casting process, there are two

objectives; high production rates and 'perfect' qua1ity of

product. In order to have both, two requirements must be met;

1) liquid steel of consistent chemistry and temperature, for the grade, must be supplied.

2) each individual component of the casting machine must have a high degree of reliability and functiona1 consistency.

30

CHAPTER 3

Solidification and Structure of Continuously

Cast Steel Billets

The overall structure of continuously cast billets

is made up of the following three distinct regions (as seen in

Figure 3.1):

1) a chilI zone immediately near the surface consisting of very fine equiaxed grains,

2) a columnar zone consisting of dendrites extending inwards from the chilI zone in a direction perpendicular to the billet surface (i.e. parallel to the direction of heat extraction) ,

3) an equiaxed core consisting of randomly oriented dendrites.

The progress of solidification along the strand

has a direct influence on the quality and the productivity of

the continuous casting operation. In this chapter the important

process variables which influence solidification and cast

structure will be reviewed. The effects of flow in the liquid

pool, steel chemistry, steel superheat, section size, and

casting speed on the cast product will be discussed. To assess

the influence of these variables, two characteristics of the

cast product, inherited during solidification, are analyzed;

shell formation in the mould and relative size of the columnar

31

W N

-*"l..\ ,J;... ..

~III \ \ \~. - r?",L-~

~.!"X ---

\ 1 --,1\ \ 1 111-

'-- ---~-~

- CHILL ZONE

COLUMNAR

EQUIAXED

LIGURE 3.1: Internai structure of contlnuously cast steel billets

~'"",

1 and equiaxed zones in the final product. These two

characteristics can usually be corre1ated to quality problems

such as surface cracks, internaI cracks and centerline

segregation and porosity.

3.1 The ChilI Zone

The chilI zone is effectively the first part of

the cast steel to solidify. The formation of this initial

shell, taking place high in the mould, directly influences the

surface quality of the cast section. It therefore becomes

important to understand the mechanism of shell formation and

growth in order to control it.

The main function of the solidified shell is to

provide support below the mould for the solidifying liquid

core. The thickness of the shell at the mou1d exit is

determined primarily by the casting speed (i.e. residence time

of the casting in the mould). Other variables such as steel

composition, superheat, flow in the liquid pool and mould design

aiso influence shell development. These variables aIl, directIy

or indirectly, affect the heat transfer mechanism within the

~ouid.

33

1

~--

From the literature surveyed it is evident that

surface cracks occur as a result of thermal and/or mechanical

stresses placed on the solidifying shell. l8 • 4 A relatively

thin she11, being weaker, is more susceptible to cracking than a

relatively thick shell. Figures 3.2 and 3.3 show transverse

cuts from two heats of AISI Grade 5160 H, 10l.6mm billets

produced at McMaster Works. Figure 3.3 shows how nonuniform

heat transfer in the corners can give rise to a relatively thin,

weak she11. These re-entrant corners, which constitute areas of

weakness, can result in a strand breakout or in surface and/or

subsurface cracks.

34

Il

1

Figure 3.2: Transverse eut of AISI Grade 5160 H (low casting superheat) •

36

Figure 3.3: Transverse cut of AI SI Grade 5160 H (high casting superheat) •

36

•

1

3.2 l nternal Structure 0 f Con t i nuous ly Cast Bill ets

The internaI structure of cast bi llets is

characterized by the relative size of the columnar and the

egu i axed zones. The si ze of these two zones i s very impor tant

in determ i n i ng the internaI cons i s tency and soundnes s of the

final cast product. Columnar dendrites, inherently weaker than

egu i axed dendr i tes, are more suscept ible to cr ack i ng .. It has

also been found that a long columnar zone increases the severi ty

of centerline segregation and porosity.

Referring again to Figures 3.2 and 3.3, it can be

seen that if the chill zone is thin, as is the case with

re-entrant corners, the columnar zone can be relatively close to

the surface. It can also be seen that when casting with

relatively low superheat, the columnar zone as well as the

center cavity, can be virtually eliminated. Thus, a more

uniform structure, consisting of finely dispersed equiaxed

grains results. Figures 3.2 and 3.3 clearly show this.

3.2.1 Mechanical Properties of Steel at High Temperature

Three di st i net tempera ture ranges have been

identified in which steel has low strength and/or ductility, as

discussed by Brimacombe and Sorimachi 18 • 4 and Vaterlaus and

wolf. 22 The first range has been found to exist between

37

1 ,

13700C and the solidus temperature. It has been propased that

in this temperature range, a low rnelting point liquid film,

highly concentrated in positively segregated elements such as

phosphorus and sulfur, surrounds solidi fied dendrites. This is

supported by measurements which show interdendritic phosphorus

concentrations of 0.2 to 0.5%, in contrast to a matrix

concentration of 0.02%. Strong segregation of phosphorus was

also found to be associated with internaI cracks.

The second zone of low d uctil i ty is in the

intermediate temperature range of 800 to 1200 o C. The

ductility of steel in this range is strongly dependent on the

Mn/S ratio and the thermal history of the steel. An increase in

the MnlS ratio results in increased ductility. Studies aiso show

that 10ss in ductility is reduced by decreasing cooling rate and

increasing isothermal holding time.

A thi rd zone of low duct il i ty ex ists a round the

700 to 900°C temperature range. It is believed that the 10ss

of ductil ity in this zone is primarily due to the precipitation

of AIN at grain boundaries. Studies show that during cooling,

AIN precipitation doesn't proceed to any significant degree.

During heating between 700 and lOOOoe however, precipitation

occurs rapidly. These find ings thus suggest that repeated

cooling and reheating cycles, which can occur in the spray zone

if it is badly designed or not functioning properly, can

possibly be a mechanism for surface cracking.

38

•

1 appar ent.

The impor tance 0 f stee 1 chem i stry i s thus

Steels high in Sand Pare most susceptible to

sur face and inter nal cr acking a t tempera t ures nea r the sol id us.

Steels of relatively low MnlS ratios (such as resulphurized,

free-cutting grades) are susceptible to cracking at the high and

ioterrned ia te low d uctil i ty tempera tu re zones. Steels w i th h igh

level s of soluble alum i num and ni t rogen are more susceptib 1 e to

surface cracking.

3 .2.2 Axial Segregation and "Mi ni-Ingot ll Formation

Macro-segregation of elements such as C, S, P and

Mn, having a distribution coefficient (k) less than one, arises

as a result of solute rejection at the solidificdtion front. 23

Liquid, rich in these s01ute elements, is swept away from the

solid ification front by convectional forces developed from

temperature differences within the liquid or by gravitational

forces acting on free growing crystals inside the liquid pool.

AxiaJ_ segregation (also refer red to as cen terl 1. ne seg rega t ion)

becomes severe when col umnar g rowth predorn inates. Fig ure 3.4

shows how the size of the equiaxed zone influences the index of

center segregation for several section sizes. The general trend

observed is that as t1le relative size of the equiaxed zone

increases, segreg a tion becomes less severe.

39

~ a -ri f.J

~7 dl H 6 bl)

dl 5 U)

H 4 dl +J c: 3 dl

U 2 4-4 a 1 >< 0 dl "j J::

H

(

Figure 3.4:

0 10 20 30 40 50 60 70 Equiaxed Crystal Ratio (%)

~ -:-161)( 161. (87)( 181. )( · ··210 x 210. 0 .. ·250 x 250. 0-··244)( lX)

(Dimensions in nun)

% Equiaxed vs. Index of Center Segregation for severa! shapes and sizes. (25.1)

40

•

1

...

When examining longitudinally cut sections, either

by sulphur print or by macro-etch, macro-segregation is seen dS

severe nonuniformity along the centerline. This nonuniformity,

as described by Moore l8 • 5 and Alberny et al. ,24 is explained

by the "mini-ingot" solidification model. Figure 3.5 shows the

four stages in the "rn ini-ingot" formation process. At (irst a

regular growth of columnar grains develops. The instability of

col umnar gr owth then leads to sorne reg ions grow i ng f as ter than

others. Th is subsequently results in the joini ng of colurnnar

regions growing from different directions thus forming a

bridge. The br idges i solate sorne 0 f the solid i fying 1 iqu id •

While this liquid continues to solidify, the volumetrie

contraction due to solidification cannot be compensated by other

liquid above since the bridges choke off any liquid supply.

Hence the t rapped liqu id is 1 eft to sol id i fy in much the same

way as conventional ingots. Coarse equiaxed grains, high in

solute concentration are found near the head of the mini-ingot

where the colurnnar zone is longest and most of the void

shrinkage cavity is also found. At the base nf the mini-ingot,

small equiaxed grains, low in solute elernents can be found.

Aiso found at the base is a significant concentration of

inclusions entrained during the descent of the solidifying

equiaxed crystals.

Axial segregation must be minimized since it is

detr imenta l to the qua l ity 0 f the cast product. Bi llets

supplied for applications such as high carbon wire, rope, tire

41

(

Figure 3.5:

l-Columnar growth

2-Some columnar dendrites grow ahead of the i r neighbours

3-A solidification bridge forms

4-The mini-ingot finishes solidifying while a shrinkage cavlty forms.

5-Actual macrostructure

Mini-ingots formation model. (24)

42

, , , J

1 mesh, etc. require uniform core properties. High solute

concentrations in the billet core, as is evidenced in

Figure 3.6, can result in breakage problems further downstream.

In today's market environment, the billet producer must overcome

the problem of segregation in order to guarantee customers a

high quality product. Failure to do so will more th an likely

result in a loss of business.

3.3 Variables Influencing Cast structure

Several process variables affect the mechanism of

solidification during continuous casting, and thus determine the

final cast structure~ The importance of uniformity of heat

transfer between the solidifying section and the mould and spray

zone has already been emphasized in the previous pages. The

other main variables which influence the cast structure (i.e.

the relative size of the equiaxed and columnar zones) are:

1) casting temperature

2) machine design

3) section size

4) flow conditions in the liquid pool

5) steel composition

43

r \

Figure 3.6: Core segregation in high-carbon spring wire resulting in breakage during drawing operation.

44

l

,-

In discussing these variables, the focus will be on the

individual and combined effects of casting temperature and flow conditions in the liquid pool (i.e. electromagnetic stirring).

3.3.1 Machine Design

The main types of billet casting machines

present1y in op:ration are:

1) the straight vertical type, where solidification begins and ends while the casting is in a vertical position,

2) the curved type, where solidification begins while the casting is near vertical and ends in a near horizontal position,

3) the straight horizontal, where solidification begins and ends while the casting is in a horizontal position.

Cast structure is dependent on the type of machine

because of the ef fect grav i ty has on the flow, in the liqll id

pool of both the liquid and the solidified crystals. For

example, in a curved machine, as discussed by Lait and

Brimacombe,18.6 the columnar zone growing from the inner

radius is usually longer than the one on the outer radius.

Crystals growing inside the liquid pool, being denser than the

liquid, settle downwards. Because of the curvature of the

strand, they settle onto the outer radius solidification front.

These randomly oriented crystals thus inhibit further growth of

the columnar zone from the outer radius. The implication of

45

1

thlS 1S that InternaI cracks will tend to preferentially forrn

along the inner radius where the colurnnar zone is extensive.

Also, non-rnetalllC incluslons, being less dense than the liquid

they displace, float upwards. The result is that a band,

concentrated in inclusions, appears about one quarter of the way

down from the top face (i.e. inner radius); hence the narne

'quarter point inclusions'. Figure 3.7 shows the inclusion

concentration band on a merchant quality billet sarnple cast wjth

no stream protection dt McMaster Wnrks.

3.3.2 Influence of Section Size

The effect of section size on the cast structure

i5 w~ll documented by Moore. 18oS It has been shown that with

increasing section size, there is a significant reduction in the

relative size of the columnar zone, accompanied by a decrease in

the width of the central segregated zone and in the severity of

segregation. Il is th us desireable, from a quality point of

V1ew, to cast large sections. From a total cost point of view

however, casting as small a section as possible, is more

beneficlal.

In casting large sections, the solidification time

lS longer. This allows for the removai of superheat in the

46

.. .

1

1

Figure 3.7: Quarter point inclusions.

1

1

'Er

liquid core long before solidification is complete. This then

results in a relatively large equiaxed core. Superheat is

therefore brought into play when considering section size. The

larger section sizes are more tolerant of high casting

superheats for the reason just given. For small section sizes

the superheat must be kept as low as possible and it is

therefore easy to see why smaller section sizes are more

susceptible to axial segregation than larger section sizes.

3.3.3 Effect of Steel Composition

As has previously been discussed, axial

segregation can be most damaging to high carbon grades which are

subsequently hot rolled and control-cooled. Elements sueh as P,

S and Mn, like C, will also segregate. Iwata et al. 26 have

quantified the extent to which these elements segregate compared

to carbon. The results are seen in Figure 3.8.

It is weIl established that carbon content

determines the solidification sequence a particular grade of

steel will undergo. Lait and Brimacombe18 •6 and Moore 18 • 5

discuss work done by several researchers on how carbon content

affects the solidification process. Table 3.1 offers a summary

of the main points.

48

~~------~-~----------_.

• 1

j

2.4

2.2

2.0

O.B·

Figure 3.8:

0 P M S' , Mn

1.0

Cleo 1.2

" 111

• o 0

1.4

Relation between segregation index of C and segregation index of P, S, Mn. (26)

49

.--------------

Table 3.1 Effect of carbon content on various aspects of solidification and cast structure.

Increasing Carbon Columnar Zone Heat Dendrite From Length Transfer Spacing

0.0 to 0.1% decrease decrease increase

0.1 to 0.6% increase increase decrease

0.6% and over decrease decrease decrease

3.3.4 Influence of Casting Temperature and Flow Conditions

in the Liquid Pool

Casting temperature and mixing in the liquid pool,

imparted by the input stream or by external means such as

electromagnetic stirring (EMS), play a major role in the

development of the cast structure. In this section, the effects

of steel superheat and EMS will be reviewed.

In the casting process, the mode of solidification

depends largely on the thermal conditions in the mushy zone.

The G:R ratio (G=temperature gradient, R=rate of

solidification), as described by Moore 25 and Lipton et

al.,2l determines the morphology of the solid/liquid

interface. Figure 3.9 shows how the morphology changes from

equiaxed to columnar to planar as G/R increases. It is thus

50

1

Equlued ColUftnar Oendrlte Cellular Planar Oendrlt.

(JI:] ~

~~ ~ ~ ~

il ~ ~ Q

Figure 3.9:

I"creas I"g G/R ... Decreaslng At

Morphology of solid/liquid interface as a function of temperature gradient (G), rate of solidification (R) and underC'ooling (61'). (25)

5 t

1

apparent tha~ by reducing the temperature gradient ahead of the

solid/liquid 1nterface, equiaxed crystallization lS promoted.

This can he achieved by operating with low steel superheat. The

application of EMS has aiso been found to promote an earlier

columnar to equiaxed transition even when casting steel with

relatively high superheat. The effect of EMS is also to reduce

the thermal gradient ahead of the solid/liquid interface. The

stirring action provided by EMS is also believed to cause the

breakage of columnar dendrites. The broken dendrites are

believed to then act as nucleation sites for equiaxed crystals.

An investigation conducted by Iwata et al. 26

showed that EMS is effective in' reducing the severity of axial

segregation in high carbon steel billets when the superheat is

relatively high. For low superheat however no clear

improvements were seen with EMS. Table 3.2 shows the test

conditions for this particular study. Figure 3.10 shows sorne

results. These results lead to two basic conclusions, which are

also corroborated by other rese~rche·s.29,30

1) The reduction in the severity of axial segregation

through the use of EMS is more apparent at higher steel

superheats. EMS has Iittle effect at low superheat.

2) When no EMS is applied, the severity of axial

segregation is significantly reJuced as the superheat

is reduced.

52

F

(Il

w

Sam pie C (%) Si (%) Mn (%) p (%> S (%) Super beat. Stirring CG) JT (ee)

H.C. ·0.15 0.28 0.58

H.E. 0.01+ 0.017 27 0

280

M.C.

M.E. 0.74 0.25 0.84 0.009 0.014 10

0 280

L.C. L.E.

0.76 0.28 0.85 0.011 0.01+ 2 0

280

Table 3.2 Chem~cal composition and cast~ng condition of test billets.

• ~

Casting !pœ (m/min)

2.+ 2.4

2.4-

2.4-

2.4 2.4-

-

HeC H.E

•• 39... 38 " l2CO 1: 1.174

l. ~ .,: 0.1 ~4 d'. C109Z

M.C Il.!

"1 31 •• 39 , tI.I~1 Il 1 104 d':QC89 d' : ClO71

Le LE .: 39 Il: 39 1: 1.109 ,. 1.132 tI.0074 d l o.c~9

Figure 3.10: Histogram of center segregation ratio of C for various casting conditions. (26)

54

•

1

This is also shown in Figure 3.11. As superheat decreases, th~

same equiaxed zone size can be achieved with and without EMS.

Another set of data, seen in Figure 3.12, shows that as

superheat decreases, the size of the equiaxed zone increases and

segregation becomes less severe.

EMS can offer a number of metallurgica1

advantages, depending on how and where along the strand it is

applied. Birat and Chone 25 • 1 describe the types of stirrers

available and their characteristics and applications. From the

literature surveyed, it is evident that EMS has been relatively

succes5ful in reducing columnar growth and axial segregation.

It is aiso apparent however that the technology i5 still far

from perfecto

A thorough discussion on sorne of the potentials

and shortcomings of EMS is offered by Tzavaras and Brody.2B

The main reason for applying EMS in continuous casting IS ta use

fluid flow in order to suppress columnar grain growth and ta

enhance the formation of nuclei for equiaxed crystals to grow

on. The metallurgical problems arising with EMS are basically

related to stirring intensity. One common problem associated

with EMS is "white banding". White bands represent areas of

negative segregation and are formed when there is a sudden

change in fluid flow conditions, as is the case with EMS. The

severity of the band increases with stirring intensity. If the

55

1

Superheat (OC)

J'l<JlIrL' 3.11: Effcct of steel superheat on the relatlve size of the cquiaxed zone wlth and wlthout EMS. (From IRSID)

56

~

u o~

..c Ul

-ri '0 C ::l +J

C -M

+J ru

<.TI Q) ~ ..c

H C) 0.. ::l

U)

• ........... 1 Ratinf! of se:!rezanon - --20 f- '"

1 1 • S~vere

• • • • . ~ A Slightly severe

• o Good l '.

---.,~-~~-- ---s----10 r- "

1 g

1

Q

o ,

50

1

1 1 1 1

c

0 0

0 0 0

" 0 0 o

,~ h'ldth of equlaxed crysta l zone (mm)

, 100

Figure 3.12: Effect of castIng temperature on the size of the equiaxed zone and on the severity of axial segregation. (27)

•

"""""'"

J

t

stirring intensity is insufficient, growth of columnar dendrites

is not hindered. In the case of billet casters, where casting

speeds can be relatively high, thus residence time in the EMS

region is short, this can occur. Multi-level stirring is thus

required. From an electrical engineering point of view, many

problems must be overcome. problems such as low efficiency of

induction motors available and control of stirring intensity

must be overcome in arder to give industry the expected benefits

from EMS.

It is evident from the preceeding discussion that

sorne quality benefits similar to those expected from EMS can be

achieved by controlling liquid steel superheat to low levels as

it enters the continuous casting mould. Furthermore,

temperature control can aiso yield benefits in overall shop

productivity. For these reasons, the remaining chapters in this

thesis will concentrate on methods of achieving improvements in

billet quality and shop productivity by a temperature control

approach. Methods available to achieve temperature control in

the ladle, tundish and mould will be discussed. Work performed

dt McMaster Works in these areas will be presented.

58

CHA PTE R 4

Controlling Steel Temperature in the Ladle

The primary production stage in steelmaking

requires that liquid steel be transferred between several

containers:

1) furnace to ladle

2) ladle to tundish

3) tundish to mould(s)

The molten steel must therefore carry sufficient superheat

(temperature above liquidus) in order to allow easy flow of the

steel from the beginning to the end of a cast. If at any point

during a cast the superheat in the tundish becomes too low,

steel fluidity decreases and the result can be an aborted cast.

This is especially true when flow control from the tundish to

the mould is by means of metering nozzles, as is the case at

McMaster Works. If on the other hand the superheat is too hi~h,

the effects on the cast product quality and on shop productivity

can be detrimental. Table 4.1 lists sorne of the consequences of

casting with high superheat versus casting with controlled, low

superheat.

Figures 4.1 to 4.3 show sorne temperature control

performance charts for a campaign of AISI Grade S160H obtained

59

1 Table 4.1: Impllcations associated wlth the lack

of control of liqUld steel superheat ln the tundish when us ing meter ing nozzles as a means of steel flow control.

,li 19h Superhea t

J - Steel level in the tundlsh must be lowered in order to reduce castlng speed

2 - Low steel level ln the tun­ùlsh leads to:

a) Increased rlsk of slag entrainment due to vortexlng

b) reduccd efficlency of in­clUSIon flotation due ta reduced steel residence times

c) reduced steel volume ln tundlsh allowlng less of a buffer for sequence castlng

3 - Increùsed rlsk of break­out

4 - Blilet internaI structure wIll contaln a relatlvely hlgher columnar zone fractlon and a lower equiaxed zone fractIon

~) A hlghcr dcgree ofaxlal segregatlon of elements such as C, P and S.

60

Low Superheat

1 - Tundish can operate at a full steel level thus normal castIng speed

2 - A full tundish level re­sul ts in

a) reduced risk of vor­texing

b) maXlmum steel residence times in the tundish th us ailowing for more efflcient flotatlon of lnclusl0ns

c) a large buffer thus allowlng more time for changing ladIes in se­quence castlng

3 - Reduced rlsk of breakouts

4 - Billet internaI structure will improve due ta a decrease ln columnar zone fractlon and a correspon­ding increase ln equlaxed zone fractlon

5 - AXlal segregatlon is less severe ~

•

1

ID . ~ CO

":' 1f1 . ,...; 1f1 0\ ,...;,...; ,...;

Il Il Il

C C III > <lJ Ci 1:'0

'0 +J !II

1 1

--' 1

1 ln ... o ...

(%) ,(.:luanba.J,:j

1 1

--r-"

.... ~

~

\CI

0 N 1

0 ... 0 ... 1

0 0

0

? 0 en

0 en 1

0 m

0 -m U 1 a

0 -"'-

0 Q)

" l. 1 ::J

0 +J co CO

0 l.

co Q)

1 Q 0 E III CU

t-0

~ 0

"' 0

"" 1 0 C")

0 C") 1

0 N

0 N 1

0 ... 10 ...

0

Figure 4.1: Distribution of average steel temperature in the tundish for AIS! Grade 5160 H.

61

o '"

Figure 4.2:

1 • •

000 .,. . .-Ior-..... 00 .-t

Il Il Il

C c 11) > Il) Il) E'U .

'U +J tIl

. • J_ !.

1-

-

i.-...-

-i.-...-

r-

'-

0

"' 1 0 111

0 111 1

0 C\I

1

o ... 1 o o ... o o ... 1 o

01

o 01 1 o CD

o , o .....

o ..... 1 o

ID

o

" o Il

o

" o "' o '1 o (f)

o

-lJ o -

CD CD o ...J

Q)

L. :J .4J CU L. Q) c. E Q)

~

Distribution of steel temperature 1055 between the ladle and the tundi5h for AISI Grade 5160 H.

62

i .a

•

1

r Figure 4.3:

. • T

1.-

0

\DN '<1'1"'\ . o-i\D.-i o-io-iN

Il Il Il

c: c: 113 > Q) QI E;ro .

'0 .., tI)

1 1 • 1 .. .!.

(%) ~ :Juanba'-'.:J

0

Y ~ 0 m 1

0 ID

0

1 ,

o f' 1 o co

o , o ri

o ~ o " o

" 1 o 1')

o ('Il 1 o cu

~ 1 o "" o "" 1

o

o o co "" o y o 01 ri

""

-U o -

Q)

L. :J +J CU t.. QJ a E QJ t-

Distribution of first steel temperature measured in the ladle for AIS! Grade 5160 H.

63

unJer normal operating conditions at McMaster Works. For this

pdrticular grade, the liquidus temperature was calculated ta be

approxlmately 1485 0 C. The aim temperature in the tundish was

1525 0 C (i.e. an aim superheat of 40°C). Figure 4.1 shows

however, that the actual average steel temperature in the

tundish, for 114 heats surveyed was lS5SoC (i.e. an actual

average casting superheat of 70°C) with a standard deviation

of 19.8 oC. Statlstically this means that, under normal

operating conditions, for a 99.73 confidence interval (Le. + 3

standard deviations), average casting temperatures from one heat

to another can be anywhere between 1496 0 C and 1614 oC. This

much variability in superheat cannot possibly lead to consistent

quality levels required to meet today's markets. It can also be

seen that to reduçe present levels of superheat by merely

reducing tap temperatures at the EAF, will have the effect of

shifting the distribution to the left thus increasing the risk

of aborled casts. It is therefore imperative that in order to

lower tundish superheats ta average levels closer to the aim,

variability must be reduced.

The impact of the present 've:; of high

variability (which again is responsible for the actual average

temperature in the tundish to he sorne 30 0 C above the aim) on

shop productivity is aiso evident from examining Figures 4.1 ta

4.3. Figure 4.3 shows the distribution of steel temperature in

the 1adle taken immediately after tap (within 3 minutes from end

64

of tap). Flgure 4.2 shows the dlstrlbution of llquid stt?C' 1

temperature loss between the ladle dnd the tundlsh (l.e. flrst

ladle temperature minus average tundish temper~ture for thE:'

ca st) •

The largest recorded temperaturp loss betwepn thp

ladle and the tundish was 140 oC. Thus, with the àlm

temperature in the tundish being 1525 0 C, ft would be

reasonable to assume that first ladle temperaturcs should not

exceed 166SoC. As it turns out however, about 11% of the

heats surveyed had a first ladle temperature ln excess of

166S oC. From this analysis, as a first step, .; C]uidel1ne WdS

given to opera tors that the absolute maximum first ladle

temperature for this grade should be 1665 0 C. Thls can

t r ans 1 a te t ° a p r od u ct i vit Y gai n 0 f a bau t O. 5 % • I:-: ti 9 h ter

control is to be achieved, operators will require more thdn Just

this guideline, as lS eVldenced from the followlng n~gresslon

analysis.

Figure 4.4 and Table 4.2 show the results of ù

simple regression analysis relating ladle temperé\ture ta tundish

temperature. The conclusion [rom this analysis i5 that tundlsh

temperature cannot be adequatpIy predicted by the startlng ]Rd10

temperature aione. with a standard deviation of the regression

of about lSoC and an excessively wide 95% confldence band, as

shawn, jt is evident that an understanding of the sources of

65

o 4-------~~----~------~_.'.-----r-------r------_r~

, \ , , ,

,

Cl

, , , \ , ,

, , \

o

o

o

o

\ ... \ \ \ \ \

'. \

'. \ \ . . . .

'\ 0

o

\ . .. \

'. \

.. \ \ \ \ \ \

'. , , , , \

\ ,

.... ID > t­G ~ C H

~

C a u

o al ca ...

o Cl ca -U

o \ . '.

M 10 171 ... 0

, \ , , , .

'\ , , , \ , . ,

'. ... 1=1 0

'. I:j \ .

... 0 \\ \\

\ \ \ \

'. \ \

... . . . . \

\ \

'. '. '.

\ \ \

\ \ \ \ \

\ 0 .. '. '.

\ \

o

. , . o \

.. .. \

'.

DO o

, . '.

\ \ ,

\ , \ .

\ . '. '.

o

. . ...

'. \ \ \

\ '.

\ \ ..

\ \

.. \ \

'.

o ,. ca ...

o N CD ...

o o CD ....

\ 0 ~-------4------~~------~\------~------~~------~Om o 0 g ~ ~ ~ O~ ~ ~ ~ m ~ ~ ~ ... ~ ~ - ~ ~ "

Figure 4.4: Regression analysis of ladle vs. tundish temper a ture.

66

-Ol '­:J ~ 10 '­Ol a. E Ol ..... Ol r-1 '0 10 ..J

(J)

~ o

Parameter Table

o 1 PARAME TER 2 VALUE 3 STANDARD 4 T-VALUE 5 SIG LEV. DEVIATION

------------------------------------- ------------------------1 2

1 SOURCE

INTEACEPT SLOPE

551 6700 o 6134

108 6300 0.0664

Analysls of VarIance Table

5 078 9 238

2 SUM OF 3 D.F. 4 MEAN 5 F VALUE 6 SIG LEV SQUARES SQLIARE

o 0001 a 0001

7 MULT R-SQ

-

8 STD OEV OF RE GR

--------~---------------------------------------~----- ----------------------------------

1 2

REGRESSION RESIDUAL

19180 25170

Table 4.2:

1 112

19180 0 224 7

85 34 a 0001 0.4325

Parameter table and ANOVA table for regression analysis.

14 99

•

ener(JY loss bet',;een the 1,1c1le and the tundlsh is imperative.

Thcse losses must be under tight control if temperature control

is ta be attnined.

The degree ta which steel temperature is

controlled in the ladle determines ta a great extent its

v,uiablllty in the tundish. Temperature variability in the

ladle arises in two forms; within-heat and between-heat.

Withln-heat Vëlrlability occurs as a result of stratificatIon of

the steel in the ladle. Natural convection, as described by

Szekely and Chen 3l , causes downward motion of steel near the

relatlvely cold ladle walls. As the ladle walls heat up the

effect diminishes but stratification in the form of a

relatively cold steel mass near the ladle bottom and a

relatively hot inner core within the ladle results. When

teeming of the ladle begins, the ternperature stratificatIon

present in the ladle 1S reflected in the steel ternperature

(iistribution obtained in the tundish throughout the cast. 1ner t

gas bubbl ing ln the ladle, such as argon or nitrogen, before the

stélrt of a cast has been found to alleviate sorne of the

stratification. Domchek 32 shows ho'''' argon bubbling can

significantly decrease within heat variability caused by

stratification in the ladle. This can be seen in Figuye 4.5.

68

•

1

It ,

2860 -

u.. 0 . 2840 U.I 0: ::l ~ « 2820 a: UJ Q.

~ UJ 2800 .... :t Vl -0 2780 z :J ....

2760

---Oc = (5/9) (oF - 32)

• 0 ... --. ..... -.. ~ '0 " , 1 .....

..... .

.,o~'-...,., , . "" "" 0, • , ,

e

• Argon Stirred o Not Stirred

10 20 30 40 50 60 70 00 90100110120

CASTING TIME. MIN.

Figure 4.5: Effect of argon bubbling in the ladle on steel temperature profile in the tundish. (32)

69

Between heat variabilty arises from differences in

operatlng p~r~meters from one heat to another. In order to

reduc0 between-heat varlability, it is important to understand

its source. This implies that aIl the energy inputs and energy

losses of the steel in the ladle must be quantified. Heas1ip et

13 ~1 • present a ladle energy balance though the use of an

energy flow dlagram as shown in Figure 4.6. In the case where

energy input to the steel is done strictly by the stee1making

furn~ce, (ie: no subsequent heating in the ladle), energy inputs

to the system are as fol1ows:

1) energy provided to the steel at the steelmaking furnace

2) energy provided to the lad1e refractory by burners

during preheating prlor to tapping of a heat

Energy losses by the system are as follows:

1) radiant energy losses during tap to the surroundings

2) energy 1055 from the steel to ferro-alloys and other

additions made to the ladle during or after tap

3) energy 1055 from the steel to the 1adle brick and from

the lad)e brick to the surroundings

4) radiant and convective energy lcsses, through the s1ag

layer and subsequently to a cover or to the environment

70

',)

1

L{lr,~pq to I\.mhi'"'llt 11,'I'\11t \1'1"

rreheat -= -"-1111 -('y ----. l/,----'4----..11t(/~ ~1 .... C'L..rj_s _--lIo.

J __________ ..!~!,'_ J

" Loqr,('g

~ Sur"rhc"t ln HeL1l ___ .....

---....... r"Jt~l1t U('at Clf Fur,lol1

Figure 4.6:

---,..

n",1t ln "rt,,1 ln Tur,,\l~h

\ Diagram of energy flow in the ladle for steel casting. (33)

11

Figure 4.6 shows how the energy content of the

stp01 dellvered from the ladle to the tundish is dependant on

loss(!s during the period which the steel i5 in the ladle. The

cnprrJY dlagram also emphasizes the importance of the thermal

state of the refractory at tape It is intuitively obvious that

the higher the initial energy content of the ladle refractory

is, the lower the losses from the steel will be. Figure 4.7

shows some theoretical simulations from Heaslip et al. done for

thrpe dlfferent refractary types and two heat sizes. As can be

spen, the steel temperature is predicted ta decrease

signlficantly throughout the time that the steel is in contact

with the refractory. The effect is more pronounced for small

ladIes than for large ones. It is also seen that energy losses

Increase with the higher quality lining8. This is especially

Important in today's operating environments where steelmakers

are using higher quality brick such as high alumina, dolomite

and chrome-magnesite.

In this chapter, a theoretical ladle cycle

sim\llation for the McMaster Works ladIes is presented. Also,

the use of a ladle furnace as a means of further improving

temperature control and shop productivity will be discussed.

72

----- -------------------------

1

l

IlSO )000

'~II 't'f"'" l '0 lont'

u ... • m

1 W

11

"' a: 7J :J ~ ... 0( ~SO i a: w tioldln, tum.n, , ... ~ leoo ;1 -. ~ '"

~O 'f,

"'urntlle

)1'9"

ltSO Alum,n.

• -f . - -i- • tHO 10 lO CO 10 100 JOOO

"ME Imln.t Ho tcmnll

1 ln to'I"

u • -"' 1&1 t lX

Holdm. lnm ln, ~ :J

~ IUO ~ W ~D 'r. e: ~ 14lumonl noo r.l 1&1 ~ - •

14, ." "" -AlumIne

IlSO ~ ____ ~~ ____ ~ ______ -A ______ ~ ______ ~ ______ ~

Figure 4.7: The effect of dlfferent refractory types on the ladle cooling behavior with a preheat of 400°C and a furnace tap temperature of 1650 oC.(33)

73

4. 1 Theoretical Slmulations of Thermal Cycling of LadIes

In arder to reduce temperature variability, a

basic understanding of the temperature los ses in the system is

required. With this in mind, theoretical ladle cycling

simulations were performed using a computer software package

known as FASTP (Facility for the Analysis of Systems in

Transport Phenomenal that was developed at McGi1t

University.34 This software js detailed in Appendix 1. The

object ives of the simula t ion .'ere as follows:

1) Determine the maximum hot spot temperature of the

ladle shell.

2) Determine the effect on the hot-spot temperature of

incorporating insu1ating tiles between the ladle shel1

and the safety lining. This action was necessitated

by the excessive shell temperatures inherent with

normal shop operation.

3) Quantify liquid steel temperature lasses bath with and

without the insu1ating tiles.

4) Suggest poss i b le changes to pres en t lad le cyc li ng

practice in arder ta reduce liquid steel temperature

variabi1ity.

74

---------------~--- --~~--

1

1 Wlth the significant lncrease ln available

computing power over the past five years, simulations of complpx

systems, like ladle cycling, have been rendered quick and

simple. This has made the use of computer madels an invaluable

tool for the purpose of process development. This sectIon

describes the approach caken in simulating ladle cycling and

presents the results of the simulations, comparing these reRults

ta sorne actual plant data. 35

4.1.1 Procedure

Prior to executing FASTP, which is ~n IntprdctlvP

program, the user must clearly define the problem in order to bp

able ta answer the program prompts when execution is inItiated.

Among other things, the program requires the input of the

dimensions of the system, the thermal properties of the

materials in the system, the boundary conditions and the actual

duration of the simulatIon. As IS the case with aIl models

based on the explicit finite difference method, assumptions must

be made; sorne due to the limitations of the model and some just

to simplify the problem to a manageable level. Figure 4.8 show~

schematically, a comparison of two lining configurations. The

only difference between the two configurations IS the

incorporation of an insulating tile layer between the s~(ety

1ining and the 1adie shel1. The thickness of the working 1ining

75

r

""-l cr>

CDI SAFETY WORKING L10UID

L1NING L1NING STEEL 3 .3 .3

p=2292 kg/m p= 2915 kg/m p=7830 kg/m

c= 1000 J/kg-OC c= 1100 J/kg~C c= 450 J/kg-OC

k= 1.4 W/m~C k= 2.2 W/m 2C 1 k= 50 W/m~C

1.27 cm ... , 1-- 7.6 cm ... 1-- 10.2 cm --, INSULATING .. 1

TILE (VI

3 p= 961 kg/m SAFETY WORKING UOUID

c =1000 J/kg..?C L/NING LINING STEEL

k= .17 W/m!C

Figure 4.8: Lining configuration and thermal properties used in the simulation. l-Regular Lining 2-Insulated Lining.

• .-..

1 was assumed cons tan t (i. e. wear was not accoun ted for). Th i s

assumption i5 valid because after a normal life the majority of

the lining maintains its original thickness, he~vy wear areas

(about 60% of original thickness remaining) are usually local

(Le. at the slag line or in the splash area). The therm,ll

resistance of the steel shell was assurned to be negliqible. The

effect of curvature of the ladle sidewall was alsn neglected.

Thus, the calculations were based on a l-dimensional heat

transfer system in rectangular coordinates.

An ideal ladle cycle based on a shop production

rate of 14.5 heats per day was simulated, as shown in

Figure 4.9. lwplied by this idealized cycle, are a number of

assumptions which are now detailed. with two ladIes in

operation each ladle produces 7.25 heats per day. Moreover,

tapping from the EAF into the ladle, was assurned instantaneous

since the time required to fill the lacile wa5 relatively short

compareci to the total residence time of the steel in the ladle.

The steel was assumed ta have a uniform initial temperature in

the lacile of 1650 0 C. During casting, the rate of descent of

the steel level in the ladle was relatively slow thus the ladle

was divided into 5 segments in the vertical direction. 8ach

segment of 16.0 tonnes required 20 minutes to empty, thus the

uppermost segment was assumed ta hdve a liguid steel/lining

contact time of 50 minutes (30 minute hald for stlrring dnd

ladle metallurgy plus 20 minutes emptying time) while the bottrnn

segment had a contact time of 130 mInutes (JO minutes plus 100

minutes before emptying) •

77

~ Cl)

-,

17 hours preheat

10 min ladle wait

1 [ ~fl35 min o preheat

u newly-Iined

ladle J 0 -

25 min

ladle preparation

Figure 4.9: Simulated ladle cyc~e.

tap (T=1650t: )

30 min hold

& stir

'\

~

100 min cast

.....

f

with FASTP, each step in the cycle was allowed to

run for the corresponding amOU'1t of time specified in

Figure 4.9. At the end of each time pe:iod, execution stopped

and a thermal profile was produced. prior to execution of the

next step, the appropriate!:>oundary conditions (or other

conditions) were changed and the last calculated thermal profile

of the lining became the initial thermal profile for the next

step in the cycle. This was continued until the thermal

pr 0 files from one cycle to another converged (i.e. essen t ially

became the rame).

By monitoring the change in energy content of the

sidewall lining during the time it was in contact with the

liquid steel, it was possible to quantify the liquid steel

temperature 10ss by the simple erlergy balance shown in

Figure 4.10. Equivalent liquid steel temperature referred to in

this section was calculated by adding the increase in heat

content of the sidewall to the heat lost by convection and

radiation to the environment.

4.1.2 Discussion and Results

1) preheating of Newly Lineà LadIes

The importance of adequately preheating h';'gh

quality ladle linings is discussed by Saunders. 36 The energy

content of the lining prior to tapping the first heat on a ladle

79

, j

(J) C)

.....

En ergy lost to

surroundings

Energy fost

by steel

Energy

absorbed

by

lining

Energy lost by steel

Energy absorbed Energy lost to +

by lining surroundings

Figure 4.10: Energy balance equating temperature lost by liquid steel to energy gained by lining and energy lost to the su[roundings.

......

1

"f

is vpry Important in that if the lining is too co1d, the risk of

skulling increases and 5:) does the probability of a'l aborted

cast. At McMaster Works, about 5% of the heats are tapped in

newly 1 ined ladIes or in ladIes taken out of service for a

pr.olonged period of time. The l ining energy content after

preheating is dependent on various factors, three of the more

important being: rated burner power, efficiency of heat transfer

between the flame and the 1 ining, and the preheat t ime.

For the pur pose of the simulation, it was assumed

that an effective heat flux of 6750 w/rn 2 was uniformly

supplied to the ladle for a duration of 17 hours. This value i s

in accordance with burner design and rated capacity.

Figures 4.11 and 4.12 show the predicted progression in rise of

the co1d-face (Le. ladle shell) and hot-fdce temperatures

respectively. As can be seen, the cold-face temperature of the

regular lining rose faster and to d higher final levei than that

of the insulated lining. This implied that heat losses ta the

surroundings were higher for the regular ladle. As for the

hot-face, higher temperatures were attained at the end of

preheating when the ladle was insulated. Figure 4.13 shows how,

cl t the end of prehea t i ng, the enerc;y input 0 f the b urner was

di stributed between energy absorbed by the l ining and enerqy

lost to the surroundings from the shell surface. For this

particular case, the ladie lining energy content was equi\1alent

ta 140°C of l iquid steel superheat for the regular 1adie and

170 0C for the insulated ladle.

81

1

1

C­• rot :::t a • CE

~ . . ~ . . G . , ·

o ., ...

· · · ta} . .

• f'4 V • ..J

v • 4.t • r1 :» • c: ...

· . ~ .

a

o "If ..

. . . . Ci) . . .

'Cil . .

o o ..

(~o) 3l:tn.1 Vl:t3dH31

. . . . a ..

. ~" . . ,

Ci)

o Il

... 'C1

ID .. ln ... .. ... (ri ... N ... ... ... 0 ...

CI

ID

...

ID

Figure 4.11: Predicted progression of cold-face temperature rise for regular and insulated ladles during prehea t i ng 0 f newl y li ned lad les.

82

...... U) a: J: -UJ X H t-

(!) Z H t-oC(

l1J J: lU a: a.

. \

" ~ • II) C:> .-4

" u ~

• ..J . <:) L ln • .... .... • :l

rot Q 01

~ U • ~ • a: ..J , . U

~ fi') • .... ., • rot '. ::J

Q CU • ~ C .... " . · ~

~

0 ....

en

CIl

" ID

ID

,.

~--~--+---~--4---~--~---+--~~-i----r---~--+---~~~ .... o o (JI

o o o o o o la

o o ~ i " i

o o O'l

o 2

--U) ct 1: -W L H t-

ID Z H t-oc( UJ 1: W ct Q.

Figure 4.12: Predicted progression of hot-face temperature rise for regular and insulated ladIes during preheating of newly lined ladles.

83

•

1

Regular Ladle

413 MJ/m2

262 MJ/m2

151 MJ/m2

supplied retained lost by by to

burner lining surroundings (100 %

) (63 0/0 ) (37%)

Insulated Ladle

413 MJ/m2

322 MJ/nf 91 MJ/rrf supplied retained lost

by by to burner Ilning surroundings

(100 % ) ( 78°/0) (22% )

T Figure 4.13: Energy distribution after preheating of newly 1 i n ed lad 1 es.

84

1

(

2) Attaining Working Temperature

A ladle is said to have attained working

temperature once the energy content of the lining becomes

repeatable from one cycle to the next. Figures 4.14 and 4.15

show how the lining energy content increased, for a regular and

insuldted ladle, as the number of heats on the lining was

increased. These figures correspond to a liquid steel/lining

contact time of 130 minutes. After about the fourth heat (i.e.

cycle 4), the energy content of the lining converged to a

maximum ann a dynamic steady-state temperature profile (i.e.

working temperature) was reached. In an operating environment,

the various steps in a cycle are not necessarily of fixed

duration. Thus, the energy content of the lining will vary

within an operating band. If the spread of the band is too

large, temperature control becomes difficult.

Table 4.3 shows the predicted liquid steel tap

temperature compensation for the first 4 heats on a new lining.

The predicted results show that the required temperature

compensation is about the same regardless of whether or not the

ladle is insulated. The large temperature compensation required

for the first heat can of course be reduced if the lining energy

content can be increased by longer or more efficient preheating.

85

1

1

l

1 i --.l • +---+ 0 0 'r- I . N \

~ \ .. i

\ .... N"'.1n . ~ t

~~~dd ~

l'

~< uuu or >->->->->- CI uuuuu z ... 0 .... rln UJ .... .. . , 00 ~ . . u -1

Z \ tt).~ H

:l: ..... r

LU ..J (J

d: 0 )-~ ~

\. ro (J '" ---" ... ~

~

~ t- e , ..... ot., <'- Z ~

l: H ,

t- '01>- ~

1< LU '. --- ~ , ,

:S~H'" 1 H 1- ~ < t' --- t- ..... l,

.... ,.0 .. i: 1 :

i ln

l- i: .~

, , .. ' , ° " ~ < ---/.

/:' ~Q. // 3Œ .• 1 .... .f 1 • 0 0 1 , , 0 0 0 0 0 . . . .

° . ID ln • tri N .... (satnor 80- 3 X .lN31NO:J lV3H

Figure 4.14: Increase in lining energy content during the first 5 heats on a regu1ar lad1e. (liquid steel/lining contact time = 130 minutes)

86

'f 4.

.... , i ..

1 1 1 L 1 0 0

~ • . . N

~\'. .i-

~r. ~N"'~" " f \ "

~~~:aJ~ " , i-~: , . uuuuu

~4h[.). )-)-)-)-)-

~ .. UUUUU al

: z " <J ~, Cl 0

H . 0 t- lin en ~

oC U -: Z

~ cfu; <: H :E -\

~ ~

UJ 1\ ~ ...J ., U

~

.~g >-~ c1!JD:- A u " --- ... ~ ,\ ,

9 1- e

' , 1-0, ,

èn ~o <: x. 0 Z ~ l: . H , ~ ~> <J -y . UJ --- ~ Il .

i/.':i/ 1 SoCHt- H

~ <J v .~ 1----il

1- .~ 0 ~ il oC ICI

W . il ~ ...

il Il. .Ia ~

#~ •. ::1 ---~ , . -i-

l,' ~~ ,/ 1

.. / l _L 1 jo.

B 1 1 0 1 0 0 0 0 0 0 • . . . . .

10 Il ~ " N ... (setnor 80-9 X ) l.N31NO:J AEl~3N3

Figure 4.15: Increase in lining energy content during the first 5 he a t son an i n sul a t ed lad 1 e • (1 i qui d steel/lining contact time = 130 minutes)

87

•

1

1

Table 4.3:

Heat t

EAF temperature cl)mpensation required for heats prior to attaining working temperature of the 1 in i ng •

1 (new lin i ng)

2

3

4

88

1) Lining Comparison at Worklng Temperature

FIgure 4.16 compares the en,=rgy content of the

two llrllnqs r)fter steady-state was atta~~jed. The energy content

1S shawn ~s a functlon of 1iqUld stee1/1ining con~act times

(I.e. for a minimum contact time of 50 minutes and a maximum

contdct time of 130 minutes). Figures 4.17 and 4.18 show how

t.he tpmper3ture between the hot face and cold face, for a

contélct time of 130 minutes, varied at the end of each step of

the cycle both for the regular linlng and the insulated 1ining,

rcspectlve1y. The followlng observations were made by tracking

the cyclt=> stc~p by step.

a) Ladle preparation

During ladle preparation a significant amount of

cnergy is lost. by the lining via the lad1e shel1 to the

surroundlngs and from th.? working lining hot face to the cold

preheater wall. It must be noted that during the time of 1ad1e

pre par . .3t ion (i. e. 5 1 ide 9 a te ma k e - u p), the pre h e a t bu rI' e ris no t

operatlng. In this particular case, the energy lost by the

1ining during ladle preparation is predicted ta be about 42

"1J m2 • ThIS flgure appl ies for both the regular and the

Insulated 1adle. In equlvalent 1iquld steel superheat this

figure represents about 22oC~ In order to reduce this 1055,

the followlng actlon was contemplated:

89

..

o 0 ~ r 1 1 w • ~

m z M ~ ~ ~ U

9 c ~

~~H~

~ ~

! Π~

o o N

c m

o ~----~--~L-·~L4----~-----+-----+----~----~ ____ ~ ____ +c c o o o .

~ o . .

~ . ~

. . N

(setnor 80-9 X ) lN31NO~ A9~3N3

Figure 4.16: Comparison of energy content of regular and insu1ated li~ings after working temperature has been reached, as a function ot contact time t.

90

~

C ~

e ~

w ~ u r u a ~ z ~

w ~ H ~

t

o o 1) ...

1 1 Y 1

o o

" ~

<J

1 1

6 1

1

o o o ... o o IIJ

(:J 0) 3~nl V~3dW3.L

o o CD

o o

"

i;

Figure 4.17: Temperature profile through a regular lining at the end of each step in a cycle after the ladle has reached working temperature.

91

-E E -lU U ct 11.

.... a :t:

~ a CI u.. UJ u Z < .... U) H 0

1

:r

.t

0 0 N

.&1--.--&1

... 1-

~ ~iffi~~ ~5~ ~mib~ uIDŒ~Q.ID'"

; 0 1 ....

... <J 0 0 ln E , .. E 1 -',/

'::1 LU l'~}'

~ U , ,/ oet , ,1

J~' lL

, Y t-/~ " / .. 0

~f 0 J: 0

/ ,'. .. :J: -/

0

kJ' a: / : lL I<J ,:; LU . 1 u .,

J <J # z oC(

/ l, t-

) <J# (J)

0 H ln a 1 ~b / \

1 , \

cri <J ~ / , ' ,

/ \ , \ 1

1 1 ,

Cl <J ~ / 1 ,

1

1 1 . , ,

0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 CD .. N 0 CD ca ~ N .. ~ .. ..

(:1 0) 3~nl'l~3dW31

Figure 4.18: Temperature profile through an insulated lining at the end of each step in a cycle after the ladle has reached working temperature.

92

1) The horizontal preheat burners will be automated in

arder ta allow the preheaters ta work at aIL times

when the ladleman is not working on the ladle.

2) The preheater wall, norrnally made up of high

temperature cement will be converted ta ceramic fiber

material in arder ta minimize radiative lasses from

the working lining.

b) Ladle Preheating Between Heats in a Cycle

During the 35 minute preheating between heats,

the mode1 predicts that the total lining energy content

decreases slightly for the regular ladle and increases slightly

for the insulated ladle however neither change is significant.

Thus the preheat burners are predicted ta supp1y just enough

energy 50 as ta compensate for lasses ta the surroundings. As

can be seen from the temperature profiles in Figures 4.17 and

4.18, even though the 1ining hot-face surface temperature

increases significantly during preheating, the temperature

within the brick (40 ta 140 mm deep) decreases. The model

predicts a working lining hot-face temperature at the end of

preheat of l225 0 C for the regular lad1e and l3lS oC for the

insulated ladle in the bottom 1/5 section of the sidewa11.

Typical measured temperatures f~r regular ladles after

preheating are between ll80 0 C and l2l0 0 C.

93

1

1

c) Ladle Wait for Tap at EAF

prior to tap the ladie is picked up and taken to

the furnace. Time studies done on the McMaster Works ladle

cycles show that on average the ladle waits for about 8 winutes

before the furnace tap begins while for 25% of the heats the

ladle waiting time is between 10 and 20 minutes. Since the

ladle waits with no lid, the hot-face of the 1ining radiates

freely ta the surroundings. The amount of energy lost by the

lining during this waitjng period is predicted to be equivalent

to lloe liquid steel superheat for the regular ladle and

13 0 C for the insulated ladle. Operating personnel have been

made aware of the effect of excessive waiting time of the ladle

and a special effort is being made to insure that ladIes are not

picked up too early prior to tap.

d) Liquid Steel/Lining Contact

In the ladie cycle simulated, the energy content

of the lining increases significantly during the time of liquid

steel/lining contact. As can be seen from Figure 4.16, this

increase is not Iinear and even at the end of cast (i.e. after

130 minutes of contact) the lining energy content is still

predicted to be rising. This clearly shows that thermal events

occurring during ladle cycling are dynamic in nature (i.e. not

steady state), thus the employment of steady state assumptions

in typicai ladie cycles, such as the one simulated, can lead to

erroneous results.

94

1

The highest rate of energy pickup by the lining

1 from the liguid steel occurs during the initial contact time.

The model predicts that for the bottom 1/5 section of the

sidewall, where the total contact time is the longest (130

minutes), about 65% of the total energy picked up by the lining

occurs during the 30 minute ladle hold (i.e. 65% of the energy

is picked up during the initial 23% of the contact time). For

the upper 1/5 section of the sidewall, where the contact time is

taken as 50 minutes, about 85% of the total energy is picked up

during the 30 minute hold (i.e. 60% of the contact time). These

figures apply to bath regular and insulated ladIes. This shows

that if a lan1e, full of steel, i5 held for a significant amount

of time prior ta casting, temperature stratification can be

considerable if the ladle is not adeguately stirred. The use of

high quality linings thus require a reliab1e rnethod of stirring

in arder ta reduce temperat'\re stratification which if severe

enough can lead to failure of slide-gate openings or poor

initial steel fluidity during teeming resulting in aborted

casts.

e) Maximum Shell Hot-Spot Temperature

Table 4.4 shows a comparison of the maximum shell

hot-spot ternperature predicted by the model and the average of

some actual measurements of ladle shell temperatures using

contact therrnometers. It must be mentioned that the plant data

1 was compiled during a period when one regular ladle and one

95

1

Table 4.4: Comparison of maximum shell temperature (oC) for regular and insulated ladIes; predicted results and actual plant data.

Regular Ladle Insulated Ladle f~::ence 1 Simulation 372 293 (14.5 hpd)

Measured 351 + 28 264 + 22

1

87

1 (11 hpd) - -

96

• Insulated laàle were ln service, both having started on new

1in1ngs. The data was gathered after bath ladIes had reached

worklng temperatures (ie: after 5 heats on the linings). As can

be seen, the results predlcted in the simulation show a hotter

hot-spot temperature for both regular and insulated ladIes.

This can be attributed ta the fact that the simulations were

based on a shop production rate of 14.5 heats per day (hpd)

whereas during the test period the actual production rate was on

average Il hpà. The lower actual production rate than the one

simulated had the effect of i~creasing the time during which the

1adle was empty, thus the average working temperature of the

11ning should be lower than predicted. If however, the model is

looked at strictly from the point of view of comparing the

effect of incorporating insulating tiles into a regular ladle,

it predicts that a decrease in the maximum shell hot-spot

temperature of 79°C can be achieved. This compares quite

favorably ta the value of 87°C obtained from actual plant

data. The model also predicts that the insulating tile hot-face

temperature in the si~ulated 11ning configuration exceeds the

maximum manufacturer recommended temperature. The effect of

thlS on the physlcal integrity of the 1nsulating tile must be

determined.

f) Analysis of Liquid Steel Temperature Lasses

As was previously stated, the change in the

thermal state of the 1ining during contact with liquid steel can

be equated to an equivalent 1iquid steel temperature change.

For the simulations, it was assumed that energy lasses

97

• experienced by the liquid steel were evenly distributed

1 throughout the bath.

Figure 4.19 shows a breakdown of where llqUld

steel temperature is lost for both the regular ladle and the

insulated 1adJe and for the top 1/5 and bottom 1/5 of the

ladle. Th~ energy terms have been converted to equivalent

liquid steel temperature losses ta the 1ining and surroundlngs.

The model predicts that, for the cycle simulated, the equivalent

liquid steel temperature lost to the lining is about the same

for bath the regular and insulated lining5, and for bath top and

bottom 1/5 sections. From Figure 4.16, it is evident that

during the time of liquid steel/lining contact, the lining

energy content profiles are essentially parallel. Thus the

model results imply that the incorporation of insulating tiles

does not cause the 11ning to pick up additional energy however

the 1ining with the insulating tiles operates at a higher energy

level (i.e. hotter working temperature) One benefit from this

can be a reduction in the incidence of skulling, if sku11ing is

a prob1em wlth the regu1ar 1adle 1ining.

Having shown that both regular and insulated

1inings absorb a similar amount of energy from the llquid steel,

it is necessary to examIne the other heat loss component (i.e.

losses to the surroundings by the ladie shell) to expldin the

difference between the two configurations. Since the effect of

98

(

(

Regular Ladre

to IInln9

COMPUTER SIMULATION RESULTS

o Equivalent liquid Steel Average Temperature Loss ( C)

Insulated Ladle

t = 50 min to hnmg

to su rrounding5' '5 Oc

53 Oc

to 5urrounding5: 10 Oc

45 Oc

to IInln9

to surroundln9S

Average

Figure 4.19:

40 Oc t = 130 min to linlng : 39 Oc

42 Oc to surroundings: 27 Oc 82 Oc 66 Oc

. 68 Oc Average 56 Oc

Difference: 12 Oc

Simulation results of equivalent liquid steel temperature 10ss (oC) for a regular ladle versus an insulated lad1e.

99

1

1

l insulating tiles is ta reduce the overall ladle shell

tp.mperature, energy losses to the surroundings are consequently

reduced. Thus the predominant effect of the insulating ti12s on

the system is ta reduce liguid steel temperature 1055 by

primarily reducing energy lasses from the 1adle shell ta the

surroundings. The model predicts that the incorporation of

insu1ating tiles ta the regular 1in1ng can result in an average

decrease in 1iguid steel temperature 1055 of 12 oC. This

implies that the steel tap temperature at the EAF can be reduced

by an eguivalent amount.

Since energy lasses through the sidewall are only

a portion of the overall lasses in the system and since actual

ladle cycles can deviate significantly from the idea1 cycle

simulated, it is difficult to obtain plant data ta verify the

accuracy of the predicted 12 0 C temperature saving. In light

of the fact however that a sufficiently accurate drop in shell

temperature was predicted by the model, it is deemed that 12 0 C

can be used as a conservative figure for the economic analysis

of incorporating insulating tiles. As it turns out, for the

results derived from the simulation, reducing the heating time

in the EAF by 12°C san result in about a 3% gain in shop

productivity for the McMaster Warks operation. Thus, the

incorporation of an insulating tile between the l~dle shell and

the safety lining can result in a significant improvement in

productivity provided of course the physical integrity of the

tile is proven to be adeguate.

100

1

Incorporation of a Captive, Lightweight Ladle Lid

The present practice at McMaster Works is to use

a ladle lid only during ceeming. This was considered in the

simulation performed. At Stelco, Hilton Works, a considerable

amount of effort has been expended in the deve10pment of a

captive lid system. The results of this effort have been

reported by Minion and Leckie. 37 In arder to evaluate the

potential benefits of incorporating a captive lid system to the

McMaster Works ladIes, a simulation, again using FASTP, was

performed. The ladle cycle used in the previous mode! was

modified by removing the preheating step prior to each tap (but

not the preheating of newly lined ladIes) and allowing the ladle

to have an efficient lid made of ceramic fiber in place at aIl

times during the cycle. The effect of this is to significantly

reduce the energy losses from the hot-face of the working

1 in i n9 •

Figure 4.20 summarizes the results of the

simulation. As can be seen, the effect of incorporating a

captive 1id is ta reduce the amount of energy the 1ining absorbs

from the liquid steel. The energy 10ss ta the surroundings

increases marginally reflecting the small increase in ladle

shell temperature. Since the lining loses less energy from the

working hot face while the ladlp is empty than when the

conventional cycle is used, the energy content of the lining

prior ta tap is higher and cl oser ta its "steady state" energy

101

•

1

Normal Cycle

to IInmg

-----------------------------

COMPUTER SIMULATION RESUlTS

Modifym9 the ladle Cyde by Usmg a Captive Ceramlc Fiber lid

For Insulated ladle; Equivalent Steel Temperature

Captive Cover

t = 50 min to lIom9

° to surroundmgs . ...1!l t d "

QC o surroun '"95

45 Oc

to IInlng 39 Oc t = 130 min to ""109 14 Oc to surroundings. 27 Oc to 5urroundIOgs 29 Oc

Average

Figure 4.20:

66 Oc

56 Oc

Average Difference.

Simulation results of temperature 1055 (OC) having a captive lido

43 Oc

Average 34 Oc

22 Oc

equivalent liquid steel for an insulated ladle

102

{f

.,l ..

contp.nt. Minion and Leckie showed that the measured working

lining hot-face temperature could increase from 795+184 oC

w i th t he normal lad l e cycle (us i ng prehea ti ng between hea ts), ta

~n pstimated 11000C using a captive caver practice. For the

McMaster Works insulated ladle described, the predicted decrease

in 11quid steel temperature loss is 22 0 C. This, combined with

12°C saved with the incorporation of an insulating tile, can

result ln a predicted EAF tap temperature reduction of 34 0 C

over the regular ladle lining and cycling practice used. This

translates into an overall productivity gain of about 8%.

Moreover, savings in natural gas consumption due to the

Elimination of preheating the ladle between heats also result.

An increase in lin i ng li fe is also expected because of the

reduct ion in spalling normally caused by large therma 1

fluctuations of the working l ini ng hot- face. Another positive

effect will be ta renuce within-heat and between-heat

temperature variability in the tundish which is an important

objective in order to achieve stability in operations and

improvements in billet quality.

103

~I "

1

1 4.2 Ladle Refinlng Furnaces

In traditional EAF steelmaking, the raIe of the

furnace was to melt scrap, dephosphorize, desulfurize and

decarburize. AIso, because the furnace provided the only energy

input to the molten steel, it was necessary to heat the steel to

high enough temperatures in order to account for aIl heat losses

in the system, as was described previously. In the past (ew

years there has been considerable interest in ladle refining

furnaces. Their cost effectiveness within the overall

steelmaking operation has resulted in a tremendous growth in the

number of ladle furnace (LF) installations around the world.

The incorporation of LP's, and their related

technology, has allowed the EAF ta be used strictly for melting

and for other refining operations which are best performed under

oxidizing conditions, such as dephosphorization. Thus a LF, aS

described by Fruehan,38,39 allows steel to be tapped at lower

temperatures from the EAF (between 2S-7S oC lower) resulting in

an increase in EAF productivity and overall lower operating

costs. Sorne of the main advantages offered by LF's are as

fol1ows:

1) productivity is increased by decreasing furnace time because of lower tap temperature and desulfurizing in the ladle.

104

(

f l

2) Steelmaking furnace refractory life and furnace availability are improved.

3) Power and electrode consumption is decreased in the EAF.

4) Temperature and composition control is improved with corresponding savings in alloys.

5) Ladle reflning practices can be carried out without the need for an increased tap temperature.

6) The steelmaker is given the ability to sequence cast several heats jn cases where the steelmaking facilities would not allow hlm ta do 50 previously.

7) A buffer is provided between steelmaking and casting which gives the steelmaker more flexibility.

8) Desulfurizing efficiency of the synthetic slag cover is increased because the ladle furnace heats the slag, and thus increases its fluidlty and causes better slag metal mixlng.

The penalty which must be paid with the LF is usually an

increase in ladle refractory costs. With this considered, there

is still a $2S/tonne claimed net saving when a LF is used.

The heating equipment used by conventional LF's i5

similar ta that used by EAF's. In general, the LF has 3

graphite electrodes, the diameter of which range from 25 to 46cm

and with a distance between electrodes usually less than that in

EAF's. LF's are powered by transforrners ranging from 6 to 45MVA

depending on the heat size being heated. The units can also be

equipped with other features depending on subsequent steel

treatment required. ThIS can range from simple stirring for

bath hornogenization to vacuum treatment for the purpose of

degassing. The cited references provide a thorough description

of equipment and suppliers.

105

. '

\

~

•

1 As mentioned previously, a ladle refining facility

can improve steel quality by allowing accurate control of steel

composition and temperature. This can be seen from data

presented by U.S. Steel, Fairfie1d Works in Figures 4.21 to

4.23. 39 •1 Temperature control performance has becn reported to

be as good as ~2oC from aim at the end of lad1e refining (i.e.

delivered to the caster). It is thus evident that with the

exceptional temperature control which a LF cao provide,

variability at the caster can be greatly decreased consequently

allowing casting with overa11 10wer superheat.

106

1 '"

Humber of Ot..nlrtloM

12

l1l 10 L

9

8

7

6

6

4 -0 ..... 3

2

1

0

,..- --

K·55 Gr8d.

X - 1.29 P.rœnt

a - 0.020 Percent

134 1.36 1.38 1.40

" Mn

Figure 4.21: Final manganese distribution for LF practice at u.s. Steel, Fairfield Works.

r

111.""'''

e"ll

Nu .... t. ... f Ob_ ...........

30

28

26

24

22

20

18

16

14

12

10 -C> (l) 8

6

4

2

0 0.42 0.44 0.45

" Carbon

0.46

K·55G"*'t

x - 0.44 Percent

CI - 0.006 Pl1ctnl

0.47 0.48

Figure 4.22: Final Carbon distribution for LF practice at u.s. Steel, Fairfield WOrks.

• ....

o CI)

..--",

HIIIII .... of OtH.nJatID ...

;.-' <1\

15r'-----------------------------------------------------------------------------,

14~

131-

121-

111-

10~

91-

8~

71-

61-

51-

41-

31-

,. ~

I/.~~

• ~

, Oc = 5/9 (oF)

-'In

~ W~

:LIIJI-Lf. RJ~'I/IIIJIIIJJ.L 1-4 58 912 13-16 17·20 21·24

1 III • Delta T. TtJnd .. h (OF)

Figure 4.23: Within-heat temperature 105s distribution in the tundish with LF practice at U.S. Steel, Fairfield WOrks.

;> -<,

• CHAPTER 5

Controlling Steel Temperature in the Tund ish

Tempera ture Vd r iati on wi th in the tund i sh can occur

in two forms:

1) variation with time (i.e. from the beginning to

the end of a ca~t)

2) variation with position within the tundish (Le.

between inner and outer strands of multi-strand

casting machines)

The first form of variation is primarily a

function of the thermal cond i tian of the steel suppl ied by the

ladle. It is typically characterized by a relatively cold steel

temperature in the tundish at the beginning of a cast due ta the

initial stratification effects in the ladle and enhanced by the

chilling effect of the cold tundish if a cold board tundish

lining practice is used (Le. no preheating of the tundish

refractory) .40 With time, under normal conditions, the profile

stab lli zes and for most of the cas t the temper a tures rema i ns

relatively constant. Towards the end of the cast the

temperature again drops. Figure 5.1 shows a typical tundish

temperature profile for a heat cast at McMaster Works.

Temperature measurements were taken at various times in the

tund ish pourbox wi th di sposable, immers i on thermocouples.

110

.,. \ ,

... i

~----~------~----~------~-----+------+------+---r--~~ .. --------- ------- ------- ----------------,r---------4---; 0

~

'0 UJ c: ID UJU

UJ :J lJ -M

:::'1 0" .-t

...J

~

0 ....

0 ID

o .. ~-----+------+------+------~----~------~------~~--~o ~ ~ ~ ~ 2 ~ i , i .. .. .. ~ ! .. .... ... ~

Figure 5.1:

(::1) 9'-'n~ g'-'9dW91.

Example of a typical temperature profile in the McMaster Works Tundish 8

1 1 t

-c .... E -Q)

E .... 1-

D C .... ""' CD lU U

1

----------------------------,.

The second form of temperature variation is

related ta flow conditions within the tundish. In this chapter,

these two forms of variation, and their sources, will be

elaborated on and sorne œethods available for reducing this

variation will be discussed. Specifically, the use of flow

control devices in the tundish, such as dams and weirs, the

application of units ta heat the steel while in the tundish and

the addition of cooling scrap into the tundish during a cast in

arder to reduce steel superheat will be focused on. The merits

of these and their applicability ta the operation at McMaster

works will be discussed.

5.1 Flow Control Deviees

At McMaster Works, an Il tonne capacity, T-type

tundish is used. The ladle stream enters the tundish pourbox

via a submerged alumina-graphite, argon purged tube. The steel

th~n flows into a distributing trough ta the nozzles which

subsequently feed the copper moulds. Figure 5.2 shows a

photograph of the McMaster Works tundish, ready for use.

In order ta develop an optimal flo~ control

configuration for the McMaster Works tundish, a full scale water

model as is seen in Figure 5.3, was constructed at Stelco R&D,

Burlington Ontario. The water modeling work described here was

112

I---~-

l

i I~

Figure 5.2: McMaster Works tundish •

.r 1 13

1

Figure 5.3: Full-scale water model of the McMaster works tundish.

1 14

1 performed by Research personnel and completed in November 1984.

Thi s work was the basis for the design of the present tund ish

used at McMaster Works (the previous tundish had a capacity of

only 5 tonnes). Water modeling is used extensively at Stelco

R&D for defining problems, recommenCling changes, development and

optimizing equipment prior to commissioning. 41

The practical and theoretical aspects of water

modeling, including the methodology of calculating residence

9 times and volume fractions, are weIl described by Kemeny ,

Oksana wilshynskylO and Lawrence. ll Figure 5.4 shows how

flow patterns can develop in the McMaster Works tundish water

model, with no flow control devices in place. with no such

devices, the fluid is short circuited to the two inner strands.

The minimum retention time for the inner strands was determined

to be 12 seconds and for the outer strands, 27 seconds. Also,

excessive surface turbulence is imparted by the ladle stream.

This implies that in a real system, it may be difficult to

maintain a uniform synthetic slag layer over the entire trough

thus resulting in excessive heat lasses, reoxidation of the

exposed steel and slag entrainment which can ultimately result

in dirty steel.

115

1

Figure 5.4:

FRONT VlEW

SIDE VIEW

Flow patterns in a tundish with no flow control devices incorporated.

1 16

J-

characterized

The total volume within the tundish can be

into three fractions:

Vt= Vp+Vm+Vo

where: Vt= Total Volume Fract ion

vp= Plug Flow Vol ume Fraction

Vm= Mi xed Flow Va lume Fract i on

Vd= Dead Flow Volume Fraction

The plug flow fraction represents a region of

larnellar motion with no backmiKing. The mixed flow fraction

represents a region where no gradients exists (i.e.

instantaneous dispersal of concentrations). The dead volume

fraction represents a reg ion where the fluid is moving so slowly

that it is assumed to be stagnant.

An optimal flow control configuration should

provide the following:

1) Containment of surface turbulence within the pourboK.

2) Acbieve equal flow distribution between the inner and

outer strands for the purpose of product uni formi ty.

3) Maximize minimum retention time to allow floating out

of non-metallic inclusions.

4} EUm i nate dead volume in order to decrease temperature

stratification.

1 17

•

1 Table 5.1 presents water model data ot how the

'optimal' flow control configuration compares to the no-flow

control configuration. Figure 5.5 shows the flow dcveloped with

the optimum dam/weir configuration. It can be seen that minimum

residence time has been ~:.ignificantly increased wi th the opt IJnum

flow configuration. More uniformity between inner and outer

strands has aiso been achieved. The optimum flow control

configuration also shows a higher plug flow volume fraction and

the dead flow volume fraction has been virtually eliminated.

From this work, it can thus be concluded that the

use of flow control devices can help in reducing between-strdnd

variations and stratification within the tundish. Efforts are

presently concentrated on implementing the optimum flow control

configuration on a permanent basis at McMaster Works. problems

encountered during plant trIals have 50 far been related to

holding the monolithic dams and weirs in place during a cast.

Solutions ta this problem are now being investigated.

5.2 Heating Steel in the Tundish During Casting

Steel temperature drops in the tundish at the

beginnning and at the end of a cast are common. The magnitude

of these temperature drops is not very predictable and depends

on a number of factors such as the history of the steel while in

the ladle (Le. stirring time and intensity, ladle lining heat

1 1 e

Table 5.1: Results of water model analysis showing retention time and volume fraction comparisons for tundish without flow control devices versus tundish with optimum flow control configuration.

Residence rime (sec) Vol fraction (~) .

Con figuration Strand Hin Peak Mean Plug Mixed Oead

No Flow Control Inner 9.6 147.9 471.9 1.8 88.9 9.2 Outer 23.0 149.7 480.3 4.4 87.9 7.6

Optimum Flow Control Inner 71.0 179.9 407.3 3).4 74.7 0.0 Outer 81.8 272.4 433.4 32.3 77.7 0.0

119

, 1

•

i

Figure 5.5:

MAIN FLOW

~ ~--

....... _-.....,-~~" FRONT V\EW

SIDE VIEW

SECONOARY FLOW THRU SLOTS

Flow patterns in a tundish with optimum flow control configuration.

120

content, etc.) and tundish brick or lining system used (i.e.

cald tundish practice or preheated tundish). In arder ta reduce

these temperature fluctuations at the beginning of cast, end of

cast and during ladle changes in sequence casting, systems

capable of rapidly heating steel entering the tundish have been

developed and are employed in several plants around the world.

Heat input to the steel can be achieved by methods

such as oxy-fuel burners, direct arc or inductive heating.

Experimental trials using electromagnetic, channel type

induction heating, as described by Ohara et al. 42 and Yoshii

et al. 43 of Kawasakt Steel Corp., have shown that temperature

drops in the arder of lO-20 oC, observed under normal casting

conditions, were reduced to o-soc when a tundish heating

system was employed. Figure 5.6 shows a schematic of the

heating system. Figure 5.7 shows a typical temperature profile

for conventionally cast heats while in Figure 5.8 it can be seen

how temperature control is improved when the heating unit is

employed. The result of this work showed that since the normal

temperature drops obtained at the beginning and at the end of a

cast represent a 'non-steady state' operating condition,

decreasing these temperature fluctuations by employing a heating

system has a marked positive effect in decreasing the frequency

of material rejection. This can be seen in Figure 5.9.

121

1

"

Iron yoke

Inductor

'" Refractory

Channe' Coollng jacket

MoIte" atee'

,~

Figure 5.6: Schematic of heating system for molten steel in the tundish during continuQus casting.(42)

\22

II

""' 0 • ~ G)

!5 ... CIl '-G)

~ G) .. ca c -E 4D Q) ...

Conventlonal + 10 heata

+6

Chan~ of ladle Te.ming end from ladl. to Tundllh

" ~~-~-=~~~~~~/~~~~~

Standard te.mlng tempe rature

o 10 20 30 ~ 50 eo 70 80 QO 100110120130140150 180

Teeming tlme (min)

Figure 5.7: Temperature variations of steel in the tundish for conventional casting practice. (42)

123

•

1

"

Experimental"- Change of heats r:1Ladle

" ",.'~

20 40 60 80 100

End of Ladle

"-~

120 140 Teeming time (min)

Figure 5.8: Temperature variations of steel in the tundish for experimental heats using the heating system. (42)

124

1

"0 G) -Q G) () () CIl

1: o c

Figure 5.9:

o 0

SUS-30.4 SUS~30

c::::J Experimental

n=20G EZI ConventtonaJ

n=21e

def.ct

SB : Fim sla b, ET: Elld .lab

Effect of using the heating system on first and last slab reject frequency. (42)

125

1

1

Tundish heating systems like the one described can

provide fine-tuning in temperature control required for critical

quality type steels. The capital cost requirements for

equipping tundishes with such a system can be significant. The

applicability of such a system for McMaster works is not

presently justifiable sinc~ temperatures are generally high.

This system may be considered at some time in the future once

better temperature control is achieved and superheats are

lowered substantially.

126

• 5.3 Scrap Addition to the Tundish to Reduce Superheat

Although scrap addition to the tundish to reduce

superheat is not (at least from an operating point of view) the

ideal method to control temperature, it can aid in solving, to a

certain extent, sorne quality and productivity problems inherent

with operating with high steel superheat. When the caster

operating crew at McMaster Works receives a 'hot' heat, the y

have no choice but to cast the heat at low casting speed in

order to reduce the chance of a strand breakout (note that a

combination of high superheat and high casting speed results in

a thinner, weaker solidified shell at the mould exit). Low

casting speed is achieved by keeping a low tundish level, which

is undesirable since this could lead to slag entrainment into

the moulds and to non-optimal flow patterns within the tundish.

Hence sorne benefits could be achieved by scrap cooling,

especially for heats which are on the high side of the

distribution in superheat, as is indicated in Figure 4.1 and/or

heats of selected grades whose final application makes them

sensitive to axial segregation.

The advantages of using scrap as a coolant in the

tundish are the following:

1) Excess heat, which would have otherwise been dissipated

uselessly (in the mould, spray zone, etc.) can be used

to melt scrap thus increasing product yield at no extra

cost.

127

1 2) A1lows operation at higher steel level in the tundish

causing less problems when sequence casting since the

'buffer' in the tundish is larger.

3) 8crap is easily available from the operation.

The amount and composition of the scrap added must be such that

the overall compc=ition of the steel cast is not significantly

altered.

5.3.1 Thermal and Kjnetic Considerations

A heat balance was performed using the REACTION

subprogram of F*A*C*T (Facility for the Analysis of Chemical

Thermodynamics), avai1able at the McGil1 University Computing

center. 44 The balance was as follows:

18.02 Fe + <A> Fe = (T,I,L) (25,1,8)

<18.02 + A> Fe (157l,1,L)

•••• (5.1)

The molten steel and the scrap were considered

to be pure Fe. To an initial amount of 1000 Kg (Le. 18.02 Kg

moles) of molten steel (denoted by 'L'), at an unknown

temperature T, an arbitrary amount, A, of scrap was added at

25°C. The final condition desired was a resulting bath

128

(

temperature of 1571oC. In REACTION, by specifying an initial

molten steel temperature T and setting AH=O (for an ideal,

adiabatic condition), a value of A was calculated. By running

the program for several values of T, Figure 5.10 was obtained.

In the form presented, this graph can be used to determine the

amount of scrap required to drop the temperature of the molten

steel by the desired amount. Thus the relation established is

as follows:

where:

S = .62 (AT)

S = Kg of scrap per tonne of liquid steel AT = drop in superheat in Oc

•••• (5.2)

The kinetics of scrap melting must also be

considered. If conditions are such that the melting rate of the

scrap is too slow, then there is risk of having an accumulation

of unmelted scrap in the area of addition. For obvious reasons

this i5 undesirable. Ta simulate the kinetics of scrap melting,

FASTP was again used. 34 This was made possible by a feature

incorporated within FASTP whereby the liguid bulk can be

instantaneously stirred thus a 'moving boundary' can be

simulated. As each node in the salld scrap, immediately in

contact with the liguid, undergoes a phase change and reaches

its liguidus temperature, its energy is instantaneausly mixed

with that of the bulk. The data generated by FASTP was treated

in order to yield the graphical reprcsentation as seen in

Figure 5.11. The model is based on l-dimensional heat transfer

129

1 \

• •

\ \

~ \

\ \

1

lb N

, \

~

1

o N

Q

\

1

11 ....

1 1

...--X ru ID

Cl 1

X

IL

\.(

lSl \

q

\

1 1

-0 ID .. (aUU01/6~) UOl~~PPV de~~s

~ \

\ \

1- 0 " "If

. r-II

"g

"1- 11 N

0 N

Il ..

!-o ....

"1- ID

--1 0 o

-U 0 -r-<l .

c. 0 t. Cl

CU t. :J ~ ID t. CD a. E CD r-'C CD t. ..... ::l 0' Q)

a:

Figure 5.10: Enthalpy balance for determining amount of scrap addition required to drop liquid steel superheat by the desired amount.

130

r

-(

",

~ J

"

~-------4--------+-------~------~~------~-------+~ .. Il ·

~ · .. N · ...

0 · ...

CD

0

CD · 0

0 . ... L ~ • · Q 0 • i ., 0 N .. · ID 0

0 . 0 0 0 0 0 0 "0 0 0 0 0 0 ... 0 0 0 0 0 ... 0 0 0 0 0 ... 0 0 0 . . . . . . . 0 0 ... 0 0 0 0 0 ... ~

..JsqwnN ~sl..Jno.:J

..-. r-4 CU

...-4 +J -ri C -ri -a: "-a:

Figure 5.11: Dimensionless relationship for melting kinetics of steel cylinders immersed into liquid steel at l570oC. Both liquid steel and cylinders are assumed to have a liquidus temperature of 153SoC and a so1idus temperature of 1490 oC.

1 31

•

1 in the radial direction of a cylinder and ln this partlcular

° case applies to liquid steel (at a bulk temperature of 1570 C)

and scrap (at an initial temperature of 2SoC) both having

simi1ar solidus temperatures of about 14900 C and similar

1iquidus temperatures of about 15350 C.

As can be seen in Figure 5.11, the relations are

plotted as functions of the weIl known dimensionless numbers

Fo (Fourier Nurnber), Bi (Biot Number) and R/R(, 't' 1) wherc; ln1 la

Fo = rA t

R(initial)

Bi = h R ... 1 k-( InItIa )

.•• (5.3)

••• (5.4)

where: = thermal diffusivity of the cy1inder (rn 2 /sec) t= tirne of immersion (sec)

R - initial radius of cylinder (m) (initial) -

R= radius at any tirne, t (m) h= conv2ctive heat transfer coefficient in the bath

(W/m -K) k= thermal conductivity of the cylinder (W/rn-K)

The versatility offered by presenting the relationships in the

form of dimensionless quantities is apparent. The chart can be

used for any combination of thermal propertles or convective

conditions in the bath.

\32

l In examining Figure 5.11, it is seen that for

lowcr values of Bi R/R goes ta values higher than l , (initial)

and then reverts ta values lower than l, ultimately going to O.

This implies that if conditions are such that Bi is sufficiently

small, upon immersion of a cylinder, a shell will form (i.e. the

cylinder diameter increases). After the shell has reached a

maximum size it will begin to melt back. This effect is very

important in determining the total time required for the

cylinder to completely melt (Le. R/R(initial) = 0).

For the case described, an expression relating the

Biot Number to the Fourier number was also derived for

determining the total required melting time (i.e.

R/R(initial)= 0). This is seen in Figure 5.12.

Fa = 14.88(Bi)-0.87 ••• (5.5)

5.3.2 Plant Trials Conducted at McMaster Works

Several st~~l cylinders of nominal diameter 23.9mm

were eut to a length of about 600mm and immersed individually

into the tundish pourbox for periods of time ranging from 5 to

25 seconds. Table 5.2 shows details regarding the composition

of the cylinders and the bath. The objective of this exercise

was ta determine an average heat transfer coefficient (denoted

by h), within the pourbox area, where the liquid metal

133

1

1

/ ~/

1

/

1

1 1

/ /

/ ,~'

o r--ru + x r-­CD

o 1 U

,1

~

x IJ..

1 /

d

/

1

1 ,.

ID

t-------~--~N~-----------r------------~----------~O o '}' '1

(0.:1) 601

Figure 5.12: Regression analysis for determining the relationship between Bi and Fo for to tal t irne requ i red to completely me 1 t the cyl i nders.

134

-·ri Dl -m o -'

{

Table 5.2:

Liquid steel Chem i s try

Cylinder Chemi s try

Compos i tion of test cyl i nders, 1 iqu id steel and bath cond i t ions.

C Mn P S Si Cr

CI.39 1.15 0.011 0.026 0.30 0.12

(1.38 1.60 0.028 0.023 0.18 0.11

Bath Tempera ture = 1560 0 C

Open stream casting ladle to tundish

135

i

-----------..

turbulence is highest, in order to see what the scrap melting

kinetics would be like. Figure 5.13 shows a cylinder after a 5

second immersion.

Due ta the fact tha t open st ream ca st i ng was used

during the test, a high degree of turbulence on the sur face 0 f

the bath was present. This l imited the immersion Ume of the

cylinders to a max imum 0 f 25 seconds since for h igher immers ion

times, the strong surface turbulence created a neck on the

sample which subsequently caused it ta break as it was being

pulled from the ba th. The immersion depth into the ba th was, on

average, about 300mm whereas the total bath depth WilS 600mm.

After the samples were allowed ta cool down,

measurements of the diameter of the cylinders were per formed.

Each cylinder was divided into IOOmm segments starting at the

immersion line. This was necessary because the diameter of the

cylinders varied from the immersion line to the base. Also, the

sur face of the samples was non-uni fo rm and wavy in na t ure

therefore severai measurements had ta be made and an average WilS

recorded. For the purpase of this a nalys i s, an overa 11 cyl i nder

diameter for each immersion time was calculated and reported in

Table 5.3. For each immersion time, a Four ier Number was

calculated. From the measured average diameter of each cylinder

a value for R/R(initial) was aiso calculated and reported in

136

(

Figure 5.13: Test cylinder after a 5 second immecsion.

( 137

1

Table !).3:

t ( sec)

5

10

15

20

22

.~2

~~-- ---~-------------------------

Results of cy1inder immersion tests used for determination of h. (d(initial)=23.9mm)

--d Fo R/R (' . t' 1) Symbo1 used

(mm) ln lIa Figure 5.13 on

29.1 o .378 1.22 Â

31.0 0.756 1.30 • 29.1 1.134 1.22 • 28.0 1.513 1.17 • 27.2 1.664 1.14 4-

26.~ 1,.891 1.10 _X

138

1

Table 5.3. By plotting values of R/R(initial) vs. Fourier

Number cnto Figure 5.11, a value of the average heat transfer

coefficient in the tundish pourbox was determined.

Figure 5.14 shows the experimental points as they

1 ie on the curves established in Figure 5.11. It is seen that

as time progresses (Le. Fo increases), the points converge to

the curve pertaining to a value of Bi of about 5. It is also

observed tha t for the lower immersion times, the points obtained

experimenta11y lie significantly away from the predicted curve.

A possible explanation for this discrepancy, as is also

described by Mucciardi,45 is that upon immersion of the cold

cylinder into the l iquid steel, an interboundary resistance in

the form of a gdp forms between the cylinder and the solidified

shell due to the chilling effect of the cylinder. For the

shorter immer sion t imes the res i stance causes a th inner than

predicted she1l to solidify but with time, and as the cy1inder

expands as it heals up, the effect of the resistance diminishes

and the experimenta1 curve is brought in line with the

predicted. It must be kept in mind that in the development of

the pred icted curves, i t was assumed that there was per fect

contact between the cyli nder and the solid if ied shell. This of

course is not the case in reality. Thus for a value of Bi=5,

-using Equation 5.4, h was calculated to be about 16000

W/rn 2_oC.

139

+-----~~----_+------~------4_------~----~~ .. ID ·

~ · .. N · ...

0 · ....

ID · 0

ID · 0

0 . .... L ~ • · i 0

i 4J 0 N 4'4 · al 0

0 . 0 0 0 0 0 0 "0 0 0 0 0 0 .. 0 0 0 0 0 .. 0 0 0 0 0 .. 0 0 0 . . . . . . · 0 0 .. 0 0 0 0 0 .. ....

""sqwnN ""sl..Jno,:j

Fig ure 5.14: Exper imental da ta po i nts as they appea r on the dimensionless relationship estab1ished in Figure 5.11.

140

-~ CO .... +J .... C -ri -CI: ....... a::

1

. l

The present tundish at McMaster Works, with the

optimal flow control configuration, has a pourbox of about 5

tonne capacity. Assuming that the entire volume of the pourbox

is uniformly mixed, for a normal casting rate of 0.9 t/min, the

residence time of the liquid steel in the pourbox is

approximately 335 seconds. Thus the melting time of the scrap

as it is added, must be under 335 seconds in order ta insure

that a build-up of unmelted scrap won't take place. By using

Equations 5.3, 5.4 and 5.5, substituting h=16000 w/m 2_oC and

t=335 seconds, the derived 'critical' diameter for cylindrical

shaped scrap (such as rad or bars) i5 133mm. This model

therefore predicts that crops from the McMaster Works Bar Mill,

which are not aIl necessarily round but do have 'wetted

perimeters' weIl below this calculated size, can effectively

meet the aforementioned kinetic criteria. The use of 102Xl02mm

billet crops from the continuous caster can also be considered

but the smaller sections available from the bar mill are more

suitable for this application.

Scrap cooling in the tundish was performed on 3

trial heats at McMaster Works. Table 5.4 shows the condition of

the trial heats to which scrap was added. The results obtained

are shawn in Figures 5.15 to 5.17 in the farm of tundish

temperature profiles during the cast. The scrap additions

consisted of 102X102mm billet crops weighing approximately 23 kg

each. The addition was made into the pourbox immediately

beneath the impacting stream where the turbulence was highest •

t 4 t

'.

Chemical Analysis

Heat No. %c %Mn %P %S %Si

23875 .48 1.53 .031 .105 .24

23982 .28 .87 .023 .019 .18

24329 .54 .86 .016 .027 .33

Table 5.4:

Liquidus Total Scrap Scrap Addition

Temp(oC) Addi tior(Ka) Rate (kg/t)

1485 320 12

1505 680 9

1485 455 13.5

Condition of trial heats for scrap additions in the tundish.

........

Predicted Drop

in Superheat (oC)

19

15

22

1 "J 'li.

o o ID ....

m c 4'4 r4 0 0 u Cl • L u en ~ ~ 4'4 a

lit c .... >r4 LO aB li • :J~ ." ~~ 1

., • • u .. 0

" .fi

1 1

1 1 1 1 , ,

1 1

1 1

1 1

1 1

1

1

o .. ., ..

-C or4 e -ID E or4 .... ca C or4 ~ ID CD U

Figure 5.15: Tundish temperature profile for trial heat 123875.

143

Il C .. ,.. a 8 a. • c. ~

Il C .... >,..

c.o aa V! 1 ::J-C .41 ~~ 1 1 1 1

1 1

1 1 1 1 1 1 1 ,

1 , 1

1

1 1

1 1

1

1 1

/ /

/ 1

1

1 1

1 -C .,.. e -CI) e .,.. t-

a c .,.. .6J 10 ., U

Figure 5.16: Tundish temperature prof i le for trial heat f 23982.

144

1

1 1

1

o Il ...

u • • u

1 1

1

1 1

1

/

" / , 1

1

/ /

/

, '" " "

a c 4'4 ... 0

:1 0. • L

t\l I! ~ 4'4 lE

1

ca C ...

> ... La a8 va I C :II! .u ... c.

o ... ln ...

-c: of'4 E -Il e of'4 .... D c: of'4 ..., ID ID U

Figure 5.17: Tundish ternperature profile for trial heat *24329.

145

1 The billet crops used were of the same grade as the liquid steol

being cast in order to avoid any possible chemistry variations

in the final product. At the scrap addition rates used, no

prob1ems were encountered except that the manua1 addition method

used was awkward. Also, when open-stream casting ls used, even

without scrap additions, a skull tends to form in the pourbox

towards the end of a heat. Tt was felt during the trials th~t

scrap additions beyond the rate of about 30 kg/tonne of liquid

steel could 1ead to skul1ing problems thus major temperature

adjustments should be made in the ladle prior to teeming while

for "fine-tuning", additions in the order of 0 to 20 kg/tonne

can be performed in the tundish.

Figure 5.15 shows the temperature profile in the

tundish when 320 kg of scrap was added at an average rate of

12 kg/tonne. The 1iquid steel temperature became lower as the

scrap addition was begun at about 12 minutes into the cast.

After about 43 minutes into the cast, scrap additions were

stopped and the temperature profile reestabl1shed itself as it

would have been without scrap additions. Figure 5.16 shows the

temperature profile when 680 kg of scrap were added at an

average rate of 9 kg/tonne and Figure 5.17 shows the effect of a

455 kg addition at the rate of 13.5 kg/tonne.

The scrap additions substantially decreased the

liquid steel superheat while at the same time allowing the heat

to be cast faster and increasing the yield by the amount of the

146

scrap added. It is believed that tundish temperature can be

successfully maintained to within + sOc by adopting a more

controlled scrap feeding rate. Smaller and more frequent scrap

additions can be used to maintain the tundish temperature

profile ln a ~ SoC control band. This however can only be

achieved by an automated control loop requiring the continuous

monitoring of liquid steel temperature in the tundish. This

technology is today still in its infancy since the traditional

method of using expendable probes has been an accepted standard

throughout the industry for yeats. This need for a method of

continuously monitoring steel temperature in the tundish, at a

cost competitive to that of expendable probes, has resulted in

the initiation of a research project in this area at McGill

University. Present studies are focusing on the development of

a direct temperature measurement method referred ta as the "self

cooling thermocouple" and on an indirect temperature m~asurement

method based on the use of the implicit solution for the finite

difference formulation of the general heat transfer governing

equations.

141

1

- -----------------------------------------------------~

CHA PTE R 6

Controlling Steel Temperature in the Mould

In the continuous casting process, the last liquid

steel transfer occurs between the tundish and the mould. In

order to insure good steel flowability between the tundish and

the mould, thus avoiding nozzle clogging, the liquid steel

temperature needs to be sufficiently higher than its liquidus

temperature. This therefore means that liquid steel entering

the mould will always carry with it sorne superheat. At McMaster

Works, 13.5mm nozzles are used to control steel flow from the

tundish to the mould. Although there can be flowability

problems when operating with very low superheat, other factors

such as steel chemistry and dissolved oxygen content in the

steel also affect flowability. The use of alternate flow

control systems such as slide gates or stopper rods can permit

casting with lower superheats relative to metering nozzlps. Tt

is generally accepted that even with these flow cO!1trol devices,

superheats in the order of lSoC are required.

It has already been shown in the previous chapter5

that lower casting superheat promotes lmprovements ln the

internal structure, as characterized by an increase in the size

of the equiaxed zone and a corresponding decrease in the size of

148

the columnar zone. With operating constraints on the superheat

1 reguirements for flowability, an external method of enhancing

heat removal within the mould would be most desirable. The use

of EMS to achieve similar results has already been discussed.

The method of interest, and the topic of investigation in this

chapter, is the use of high purity iron powder, added to the

mould during continuous casting, as an agent to provide

supplemental cooling to the liguid pool. This chapter reviews

previous work performed in this field anl presents results of

trials conducted at McMaster works.

6.1 previous Work

~ The use of iron powder as a supplemental cooling

agent during steel solidificatIon has been ir.vestigated by

severa1 workers both on continuously cast sections and on

conventionally cast ingots. Bohm46 discusses sorne of the

qua1ity improvements reported when high purity iron powder is

added to the mould during contirJous casting. Figure 6.1 shows

140Xl40mm billet samples cast at the Terni (Italy) Steelworks.

Samples 23 and 24, taken from conventionally cast billets, show

a significant amount of axial segregation and center1ine

porosity. An addition of iron powder, in the order of 1.0 to

1.4% resulted in a much more homogenous structure, relatively

free of segregation and center porosity, as can be seen from

samples 25 and 26. Trials with iron powder, produced by the

149

1

Figure 6.1: Sample Sample Sample Sample

23-Transverse section, no iron powder (46) 24-Longitudinal section, no lron powder 25-Transverse section, with iron powder 26-Longitudlnal section, with Iron powder

t carbonyl process, ta the tundish and mould have also been

conducted in West Germany. Figure 6.2 shows a diagram of the

fe~ding technique. Unfortunately no published results of these

trials could be found other than a report stating that results

were encouraging.

Published results describing experience with Iron

powder additions during solidification of molten steel, are from

two sources:

1) USSR, Institute for Casting Problems of the Academy of Sciences of the Ukranian SSr and the Krasnoe Sormovo Works.

2) Italy, Centra Sperimentale Metallurgico, Rome and Terni Steelworks.

This work is further elaborated on in the pages following.

6.1.1 Work From USSR

Zatulovskii et al. 47 investigated the effects of

adding up to 4.5% iron powder to 8 tonne forging ingots and 1-2%

iron powder ta several sizes of continously cast slabs, blooms

and billets. Both carbon and alloy grades were investigated.

Ingots cast with iron powder additions (trial ingots) and

conventionally cast ingots (reference ingots), were examined by

machining longitudinal sections in the axial regiOtl. The

sections were subsequently etched in arder ta reveal the

t 5 t

1

1

-~----------------------------.

Position A Carbonyllron powder + carrier gas

Pou ring nozzle --

Sirei rnpll

- -- DIC;IJlbuhonlaundpr (Iundlsh)

Posilion B CartJonylllon powder 1 carrlcr gelS

POslllon C Carhonyl Iron powder camer gas

Waler -c:::~~~~

ConlllllfOll') caslll1q molli

=: .. .....,,........ 4--- Willer

Water - '---...--r==-="=

Figure 6.2:

Slab

Iron powder addition rnethod using a carrier gas. (46)

152

mIcrostructure. It was found that the powder addltlon had a

favorable effect. The columnar dendritic zone was substantially

reduced whlle the equiaxed zone was larger.

Segregation studies were performed by two methods;

by measuring hardness and by chemical analysis of drillings from

the axial sections. The trial ingots exhibited more uniform

hardness readlngs throughout the entire section and did not

exhlbit the troughs and peaks characteristic of the reference

ingots. The trial ingots also featured a more uniform

dIstribution of phosphorus, sulphur, manganese and silicon.

Furthermore, trial ingots also were found to be superior in

mechanical properties. A significant improvement in ductility

was observed.

( The iron particles, all other factors being equal,

cause changes in the solidifying liquid pool within the ingot:

1) the superheat ln the liquid pool is dissipated

relatively quickly

2) temperature gradients in front of the solidIfication

front decrease

3) there lS an Increase ln the number of heterogenous

nucleation sites

4) the liquid pool viscosity increases thus reducing

convection currents near the solidification front

(

153

1 These effects aIl contribute to minimize columnar grain growth

and severity of axial segregation.

Similar studies using iron powder were also

conducted on continuously cast 420X175mm slabs. Powder addition

ranged from 1 to 2%. Among the beneficial effects observed

werei

1} the wi thd rawa 1 speed was i ncreased by a factor 0 f

1.3-1.5

2) the equiaxed zone was increased by a factor of 1.4

3) greater structural and physiochemical homogeneity

4) reduced central looseness and segregation

Simi1ar resu1ts were obtained when adding 0.5-0.8% iron powder

of partic1e size 0.5-1.0mm to 280X290mm and 175XI020mm sections.

The studies were taken one step further when a

1.0-1.5% addition of iron and ferrotitanium powder was made ta

100XIOOmm billets. This was found to be very promising because

the addition of iron powder, together with powders of

ferroal10ys can give the same effects as described but with a

smaller total quantity of powder.

The authors of this work thus concluded that the

use of additions of iron powder may result in marked technical

improvements and savings. with this technique it may be

154

(

possIble to increase withdrawal rate by a factor of 1.5 to 2.0.

The authors aiso expect a reduction in material rejected, an

improvement in rhomboidity in billets and an increase in the

0rades of billets that can be continuously cast.

48 Another study, conducted by Yur'ev et al.,

describes how powdered pig iron was added to the liquid pool of

solidifying billets, slabs and ingots. The study stresses the

importance of selecting fine powders in order to insure that the

particles completely melt. It was found that for coarse

particle sizes (for instance 3-5mm) the powder did not

completely melt resulting in sorne particles appearing as

exogenous inclusions. The same effect was observed when too

much iron powder was added causing excessive cooling of the

liquid pool resulting in incomplete rnelting of sorne particles.

The conclusions from this study suggest that powder additions

should be based on liquid steel superheat. Additions should be

enough to lower the superheat to levels close to the liquidus

temperature but not to supercool the liquld pool. Improvements

in the internaI structure were aiso observed, thus in agreement

with resu1ts pub1ished by other researchers.

6.1.2 Work From Ita1y

.. 1 49,50 Rarnacclottl et a • have conducted detailed

studies on the effect of iron powder addition during the casting

of 140Xl40mm billets for reinforcing bar applications. An

155

1

! , j

apparatus for feeding the powder to the mould WdS developed and

1s shown in Figure 6.3. The system consists of a stordge hopper

and a feeding hopper. A screw feeder transports the powder to a

discharge point. The quantity of powder supplied can be

regulated by accordingly adjusting the rotational speed of the

screw. The powder is subsequently delivered to the mould by an

inert carrier gas. This system provides good consistency of

material flowrate and accuracy of control of the powder jet

trajectory. This is essential for the regularity of the

process.

For the trials, iron powder was added to a single

strand while the other seven strands were cast in the

conventional way. By simultaneously sampling the trial billets

and the reference billets, a good comparison of the cast

structures was obtained. A total of 13 experimental casts were

performed; 9 at standard speed (2m/min) and 4 at increased speed

(2.5-3.0m/min). Two powders, differing only in average particle

s1ze were studied; O.6mm and O.4mm diameter. The chemical

composition of the powder was the Eollowing (in percent) :

c si Mn S P Al 02

1.0 .68 .69 .015 .013 .04 0.1

The powder was fed at a rate between 1.5-3.0% for the standard

casts and 1.4-1.6% for the high speed casts.

156

1

AIICON

Figure 6.3: Schernatic drawing of rnetal powder feeding apparatus.(49)

157

•

1

From the study, the experimenters wer~ able to

draw the following conclusions:

1) the feeding of iron powder into the mould during

continuous casting did not create negative

interferences with the casting process and can be

completely automated.

2) an increase ln casting speed in the order of 40 to 50%

was achieved.

3) the finer powder (O.4mm diameter) was thought to bo

easier to melt, possibly due to it greater surface areQ

per volume added.

4) a uniform distribution of solute elements was achipvpù.

5) an improvement in the size of the equiaxed area was

obtained.

6) a more homogenous inner structure also led to improved

mechanical properties of the product.

7) in the case of a curved mould casting machine, the

addition of iron powder can contribute towards limiting

the problems of breakouts due ta erosion caused by the

casting stream on the outer radius shell. A reduction

of the assymetry in the solidification structure

between the inner radius and outer radius is also

possible.

8) there is a possibility of using thi.c; technique for: the

purpose of producing steels which are otherwise

difficult to cast.

158

1

(

6.2 Trial Work Performed at McMaster Works

The trial work conducted at McMaster Works was

cdrried out in two stages. In the first stage of the work, the

prepara tory stage, aspects such as the economics of adding

powder, the source and availability of iron powders for both a

trial period and for eventual permanent or semi-permanent

utilization, and the characteristics of the powder, such as

granulometry and chemical composition, were investigated. Aiso

in this stage, the development and construction of an adequate

powder feeding system for the trial period was undertaken. The

second stage of the work, the experirnental stage, consisted of

making iron powder additions into the mould during casting and

subsequently examining the effects of the addition on the

internal structure of the as cast billets when compared to

billets with no powder addition.

6.2.1 Preparatory Stage

A local producer of metal powder products, Quebec

Me ta 1 powder s Ltd. (QMP), was con tacted in order to obta i n

information on the different iron powders produced and to

determine the most suitable for the application of feeding into

the continuous casting mould. QMP operates a large, modern and

highly dutomated plant near Sorel, Quebec. The raw material for

the production of their ATOMET Iron powders is a high purity pig

159

1

,

iron supplied by the smelter of QMP's parent company, Quebec

Iron and Titaniun Corp. ATOMET powders are available in a

number of different grades of varying screen ana1ysis and

chemica1 composition. Iron powders have a multitude of

applications ranging from metallurgical uses such as powder

metallurgy to use a5 additives ta food substances to provide a

supplementary source of dietary irone

Implementing a practise of adding iron powder ta

the mould will result in an increase in billet production cost.

This increase is of course a function of the grade and purity of

the powder, thus affecting its cast, the relative amount of the

addition and the se11ing price of the billets produced, since

the powder addition also increases the billet production by the

amount of powder added. Table 6.1 shows how the billet

production priee increases as a function of the priee of powder

and the amount of the addition. As can be seen, it can be

economically feasible ta use iron powder as a supplemental

cooling agent if the amount of the addition can be kept

relatively small and lower priced powders are used. Typical

powder priees range in the order of $O.SO/Kg.

As ~as seen in the previous section, the two

benefits available from using iron powder as a supplemental

cooling agent are:

1) Quality improvements in the cast product.

2) Casting speed, thus cas ter productivity, can be safely increased.

160

Table 6.1: Increase in billet production cost ($ per tonne) as a function of amount of powder added (shown in weight %) and price of iron powder.

Po wd e r P rie e Amount of powder Addition ($/Kg)

0.5% 1.0% 1.5% 2.0%

0.44 0.83 1.65 2.48 3.30

0.66 1.93 3.8 r:: 5.78 7.70

0.88 3.25 6.05 9.30 12.10

1.10 4.13 8.25 12.38 16.50

1. 32 5.23 10.45 15.68 20.90

1.54 6.33 12.65 18.98 25.30

1. 76 7.43 14.85 22.28 29.70

161

•

~t present, the McMaster Works caster Cdn meet the overall shop

requirements in terms of productivity. If future shop operation

levels make it such that the cas ter becomes a bottleneck in th~

operation, the use of iron powder additions can be considered as

a viable method for increasing caster productivlty. Thus if

iron powder additions are to be considered, it will be strictly

in the context of improvlng billet quality. This can permlt

McMaster Works to:

1) Improve quality on presently produced grades such as cold heading quality billets in order to insure market retention.

2) Achieve quality improvements, namely reduced axial segregation, on grades which cannot presently be consistently produced, thereby introducing areas of potentlal new business. These grades include hlgh carbon wire and high carbon forgings.

Achieving the two goals stated above can more than make the

increase in production cost due to the iron powder additIon

worthwhile.

Any addition made to the mould during continuous

casting must meet the following criteria:

1) must not significantly affect the chemical analysis of the steel grade being cast.

2) must readily go into solution and uniformly distributc within the cast section.

3) must not contaminate the steel being cast with non-metallics or other unwanted elements.

162

Thesp cri ter i a can be adequ a tel y :net by choos i ng a powder of

1 dppropriate purity and granulometry. Discussions with a QMP

technical sales re[?resentative confirmed that highly pure

powders, suitable for this application are readily available.

Initially, 2-20 Kg sam[?les, 10 Kg of ATOMET 28 and

10 Kg of A'roMET 602 were r ecei ved for evalua t ion. Table 6.2

shows sorne character istlCS of the two powders. The main

di fference between the two powders is in the particle size

dIstribution. ATOMET 602 is of larger average [?article size,

beIng [?redominantly a -150/+'75 micron powder (Le. no fines).

1\5 a consequence, ATOMET 602 also contains significanlly less

oxygen and is thus of higher purity. Excessive oxygen levels in

the powder are undesirable since non-metallic inclusions in the

final product can be generated through reoxldation reactions.

l\na 1 yses [?erfo rrned a t McMas ter Wo rks showed t ha t the oxyg en

con tent () f the +45m i c ron fr act ion of ATOMET 28 was on aver age

about 1350ppm whereas the -45rr i '-'ron fraction contained in the

ord er 0 f 12000 ppm ox ygen. As can be expected the penal ty for

this is in the price of the powdersj AT OMET 602 is substantially

more expenslve than i ts less pure counterpart.

Any powder add 1 t ion made to the mould, must be

made ln such a way that it is unifor1nly distributed within the

l iquid pool. Also, the powëler must be fed at a steady rate.

Fa il urE' ta do 50 can resul t in a non-un i forro cast billet, which,

needless to say i5 highly unàesirable.

163

Table 6.2: Properties of ATOMET 28 and ATOMET 602 i ron powders (as per QMP catalogue) •

ATOMET 28 ATOMET 602

Density ( g/cc) 2.85 3.00

Granul)metry

+ 212m i cron Trace Trace -212 +150 5% 0.3% -150 +106 28% 52% -106 +75 23% 45% -75 +45 24% 2.5%

-45 20% 0.21,

Chemica1 Analysis

C 0.07% o .10% 0 0.18% 0.10% Fe 99+% 99+%

164

-

1 Thus an adequate powder feed i ng system, able to

hdndle thp trial powders chosen had to be developed. The first

systems tr ied were based on the use of a carrier gas. In

9 en e raI the y con sis t ed 0 f a ho l d i n 9 b i n w i t h a met e r i n 9 no Z z le

élischarCjlng into a powderjgas mixing chamber. The blend of

powder dnd gas travelled through a tube and was discharged from

~he end of the tube. Several versions of this system were tried

including opened systems and pressurized systems. AIl however

gave the same problems in that clogging or bridging at the

metering nozzle was occurring. Th j s caused lhe di scharged

material ta come out in erratic pulses. Another problem also

associated with the lise of a carrier gas, was that at the exit

of the tube there was particle stratification, with the larger

particles having a relatively small trajectory whereas the finer

par tic 1 es we r e b l 0 w n 0 f f cou r se. Th i 5 pro b l ern wa 5 rn u c h m 0 r e

ev iden t w i th ATOMET 28 si nce the large -4 Sm icron fract i on ca used

signi f icant fuming at the exit point. Based on these

deficiencies, it was decided that the powders bejng tried could

no t b e t:? f f e c t ive l y f ed w i th a car rie r 9 a s s ys t em the r e for e

further t:?ftorts were concentrated on developing a mechanical

feeding system.

A f ter mu che f for t, a s y stem, se e n i n Fig ure 6. 4 ,

·,v.IS rut ln place WhlCh COl'~j adequately feed the iron powder

Into the mou1d. It conslsted of a holding bin of 50 Kg capacity

discharging into a 'T' pipe. One end of the 'TI pipe was

attached to a stainless steel tube about lm in length used to

165

8

l :D-=-=-

.... ,....... ----1

7 ( ! / ) \ J

l 'J / 2 3 ---r . ----9 ~ _ ' r, of • .'

l, " l , , 1 \[)iL_~ (~=jV ~--:-. ~_:::. :::~ :::::: :)

Q. -9

;: ~)~

~T l '" -. - 01 ~~~- /- 7--'-",,? 0 C'~, --

0' ~-~~-.-'-o' o 0\

.4

- :: Storage 91n Screw Housing Screw

'\ ----- li l __ L \

J

1-2-3-4-5-6-7-8-g-10-11-

Drive Shaft Bearings Drill Motor Rheostat 115V-AC

'\ ~-----~ ~.r --- -

'-~-

~- : ~_.2~ ~ o:::::~ ~ - - --, - - -::J - - 1 L

----~ ~ -------~

, ,

\

- '" 1 ~ ........

~i '~--~ç,:~ 1-- I?-

/ ,~-------"

12-13-14-

Screw Speed Control Discharge Spout Feed Angle Adjustment Height Adjustment Trolley Handle

Flgure 6.4: Schematlc of powder feedlng system developeà for trlals at McMaster Works.

-

housc the feecling screw. The screw was fabricated using a 3.2mm

thick stainless steel strip. The screw ran through the 'T'

pi pe, whe re the i ron powder was pi cked up and transported to the

dlscharge spout. The other end of the screw was pinned to a

drlve shaft which in turn was supported by a pair of bearings in

order to provide stability and to prevent the drive end from

seizing. A 19mm drill motor was used to drive the screw. The

powder flowrate was controlled and regul~ted by means of a

rheostat. with this system, both types of powder could be

effectively fed. At the discharge spout the powder had enough of

a traJeccory to permit the placement of the tip of the spout

away from the molten steel stream.

The whole feeding apparatus was mounted on a

trolley which provided a height adjustment and an angle

ad]ustment of the feeding tube 50 as to permit regulation of the

powder stream onto the molten steel stream. The trolley had

large wheels making it maneuverable enough 50 it could quickly

be removed in case of any emergency.

Preliminary tests were performed with the feeding

apparéltus by running iron powder through the system and into a

l'ontinner. A compact stream with just the right trajectory was

obt.llncd by adjusting t~e angle of the discharge spout. A steel

rad was used to represent the liquid steel steam and the proper

distLlnce, height. and angle of the powder discharge were

determined. The requlred powder flowrate was also adequately

167

..

1 controlled with the rheostat available. Calibration of the

system was performed simply by setting the rheoatat controiler

at various points and measuring the powder discharge rate with cl

calibrated cylinder and a stop watch. Once aIl this prelimindry

work was completed the apparatus was deemed ready for the

trials.

6.2.2 Experimental Stage

For the trial, 500 Kg of ATOMET 602 were provided

by QMP. In order to properly assess the effect that the powder

addition had on the solidifying billet, the sampling schem~

outlined below was adopted:

1) Strand #1 was chosen for the trial.

2) The strand was allowed to cast normally for the first 30 minutes. Steel temperature and casting speed during this period were monitored for stability.

3) Once a stable casting operation was achipved a I:i llet sample, about 30cm ln length was eut and allowed ta cool. This was the "control" si3mple. Operatlng clat" such as casting speed and steel superheat were recorded.

4) with the casting parameters remalning relatively constant, d powder addition in the order of 1+0.2% was begun and was a110wed to run for about 30 min~tes in order for the casting strand to reach a new steady-state. A billet sample representatlve of this period was cut. ThlS was the "trial" samp]c.

lee

Several trlal heats were produced with the powder addition. Due

to the llmited work area available around the mould the trials

had to be limited to merchant quality r.einforcing bar grades

because on these grades, no tundish to mould str.eam shrouding

with inert gas nor aluminum wire feeding is required. AIso, the

present inert gas shrouding system employed at McMaster Works

would cause the powder ta be blown off course.

Two persistent problems were encountered which

lirnited the success of the trial. The major problem was that

the feeding apparatus could not be easily placed 50 that the

trajectory of the iron powder would fall directly onto the

molten steel stream. Evell if good alignment was achieved, with

aIl the movement on the caster floor the apparatus could have

easily been displaced causing the liquid and powder steams to be

misaligned. The ultimate result of this was that the iron

powder wound up cnte the billet corner causing a severe broken

corner condition. This however did not cause any strand

breakouts. The billets produced with the broken corners had to

be scrapped.

Another problem encountered was that the rheostat

settlng knob WdS too sensItive thus a calibration was required

before (ldch rune 1\150, occasional fluctuations in the line

voltage cdused inconsistent motor velocities thus inconsistent

powder (eed rates.

169

•

Despite the problems two relatively problem-frep

heats were produced. Table 6.3 shows the chemistry of the trial

heats along with sorne casting parameters. Figures 6.5 to 6.8

show macro etches of the trial billet sampI es ,md the control

samples. The etches were clone on billet samples cut

longitudinally at the center and immersed into hot 30%

o hydrochloric acid (about 100 C) for 45 mlnutes. They were

then immersed i nto a di 1 ute nitr ic aCld sol ut ion for l mi nu te

and wiped clean. This last step was found to give excellent

exposure of the macro-structure also allowing samples ta be kept

for months in the i r as-etched cond i t ion wi thou t any r ust

forming.

In comparing the macro-etches of the trial bi llet

sample and the control sample for the two heats, the following

observations were made:

1) Tte control samples showed a larger proportion of col umnar dend rites than the tri al samples. The columnar grains were also larger in size in the control samples.

2) The equiaxed reg10n was less pronounced and of a larqer graIn size in the core of the control samples than that of the trial sample.

3) Core porosity was much more pronounced in the control samples, being almost continuous while in the trial sample, although not completely eliminated, it was much less severe.

4) Quarter pOInt inclusion fields were found to be simllar for both the trial samples and the control samples suggesting that the powder addition did not have any significant effect on the number of non-metallic inclusions present.

170

1

Table 6.3:

AnJ l :i ~'>l s

IIL';l t- # l' l'-ln

~ ,) ]32,11 o H l • U6 " ... )

-4.

,J) }2':J l U.24 l.n

.. ,

Chemistry and casting parameters for iron powder addition trial heats.

( c" ) Cùstlng Speed Castlllg Tern)?er.:l turE:

S Si rn/nnn. Oc

.017 .2G 2.8 1574

.032 • 1 lJ 2.b 1552

171

.J

l

Figure 6.5: Heat #33241 control billet sample macro-etch.

, Figure 6.6: Heat #33241 trial billet sample macro-etch.

Figure 6.7:

\~

Figure 6.8:

. ' . . .. t, ! l" ",:

" ~ \

Heat #33291 control billet sample macro-etch.

. ,'. ...- ......... ..:{.

Heat #33291 trial billet samp1e rnacro-etch.

CHA PTE R 7

1

SummQry and Conclusions

The superheat of the liquid steel dellvered into

the continuous casting mould, determlnes to a gle~t extent lhe

quality of the cast product and the productivity of the steel

shop as a whole. Wi th respect ta quaI i ty, 1 t WilS shown through

a literature review and by way of examples pertainlng to the

McMaster Works operation, that casting stéel wlth cl relatlvely

high superheat has a d~trimental effect on the internal

structure of the final cast product.

A cast product's internaI quality is chdrilcterized

by the extent of axial segregation, core poroslty, mini-Ingotism

and columnar ta equiaxed zone ratio. It was seen ln Chapter 3

of this thesis that as the equiaxed crystal ralio increases, the

severity of core segreg~tion of elements such as C and S, ~nd

the incidence of core porosity decreases. Segregatlon of C ln

steels destined for critical end uses such as hi9h cArbon sprlng

wire, can cause breakage durlng the Wlre drawlng operatIon or

premature failure of the final product. It w<:ts S0cn thal

electromagnetic stirrlng of the llr]Uld pool Cdn homo<]enlzP lh,.!

solidifying core th us mlnlmizing segregatlon and core poroslty

t

174

•

while increasing the size of the equiaxed region. InternaI

quality improvements, similar to those obtained by casting wilh

relatively high superheat and wi th EMS can be acllleved by

casting with low superheat.

From an operating point of view, high casting

superhea ts are requi r ed when liqu id steel tempera t ure los ses are

not controllable. FIgures 4.1 to 4.3 clearly illustrate this.

Statistically it can be seen that a high mean liqUld steel

tundish temperature is requlred due to the large variability, as

evidenced by the standard devlation, present in the

distributIon. with the ]ower tail of the distribution alreddy

being near the liquidus temperature, the problem cannot be

corrected just by asking EAF operators to lower tap

temperatures. Doing so would shift the whole distribution in

the direction of the liqUldus temperature inevltably resulting

in an increase in aborted casts due to freeze-offs or 105s of

fIu.l.dity. Thus before asking the operator to lower dverage ta[>

temperatures, the variability in the temperature distribution

must be decreased. Doing so can result in significant

improvements in the over a 11 oper a t ion namely ln dec rease 0 f

refractory consumption, less downtime for furnace and ladle

rel ines, reduced risk of furnace and ladle breakouts, improved

cas ter productivity due to the allowance of higher casting spced

without increasing the risk of breakouts, etc.

175

In order to achieve temperature control, it was

shawn in Chapter 4 of this thesis that it is important te

und e r stand the energy losses in the system and to min im i ze

them. One of the mi3jor energy los ses for the l iquid steel

occurs during its contact with the ladle lining. A computer

software package developE.d at McGIll University was modified and

fea tut:es were added tù i t in ordet: to enable the sImulation of

thermal cycl ing of ladIes. From the data generated by the

prog ram, supported by ac tual plant data from Steleo McMaster

Works, the following conclusions were drawn:

1) An insulating tile between the ladle shell and the safety I ining can be effective in reducing ladle shell tem}Jerature thus also reducing energy losses to the env ironment. The model pred icted a tempera t ure decrease ln the ladle shell hot spot of 79°C (versus an auerage of 87 0 e actually obtained). ThIS can translate lnto a decrease in EAF tap temperature of 12°C. ThlS ln turn would resul t in an approximate produetlvity gaIn of about 3% for McMaster Works Steelmak] ng.

2) The inC'orporation of a hlghly insulatlng ladle lId system attached to the ladle at a11 times when the ladle lS less thdn full WJ th llquid steel can be an effectIve barr ier to energy los ses from the ladie brick hot face. By retaH'llng more energy Inside the brIck, the IjqUld steel WIll lose less temperature to the linlng. The model predlctç;d that EAF tap temperature ca n be fur ther decreaseù "::>y 22 0 C • Comb i n2d w i th the previous savlogs of 12

oe, shop praductivlty can be Increased ln the arder of 8%. The fact that the ladle working llnlng is subJected to much reduceù thermal cycling (le. changes ln hot-face temperature) can aLo result ln an Increase in 110lng life. Anolher potentlal beneflt of using thlS type of lid 15 that if ladle ..:ycle times ùre short enough, gas preheating between heats can be elimlnated thus provldlng further significant cost reductions. These flndings are well

176

• in agreement wlth results documenteù at Stelco Hilton Works where aIl the ladles were modified to incorporate a captive Ild system.

3) The rnaturity of ladle furnace rechnology combinr~d wlth ladle metallurgy practices is provlding steelmaklng shops with greater flexibllity in the control of sb.'d superheat, chemistry and InclusIon morpholagy. Signiflcant quality and productlvlty Incredses hllve been reported and can be dchieved with this technology.

During its trans[er fr.om the ladle tü the lundlsh,

and while l'es id i ng i ns ide the tund i sh pr i or ta enter i n9 lhe

moulds, the liquid steel is again subjected ta thermal and flow

conditions which lead to teroperature variability. For eXilmple,

at the beginning of a casting sequence, the steel entering the

moulds is relatively cold due to the combined effects of

stratification in the lafUe and the chilling eEEect of the cold

tundish l'efractory. It was shown in this thesis that the

thermal profile of the liquid steel in the tundish is strongly

in fI uenced by the thermal hi story 0 f the lad le. Temperature

adjustments ln the tundlsh are possible bl' several means.

Between strand temperature variability can be reduced by

emplol'ing flow control devices in the tundish 50 as to

homogenize residence t imes between inner strands and outer

strands. Optimum flow control device confif)uration for the

McMa ster Works t und ish was determ i ned by Steico resea l'ch

personnel using water rnodell ing techniques.

177

Witllln heat vanability sueh as the front end

e[iects j\1st described may be controlled by a heating system in

the tundlsh. Th i s techno l ogy i s not yet fu Il y developed and a t

present would not be appl icable to the McMaster Works operation

Slnce as is the case, present practice is ta cast with

relatively high superheats. A more likely possibility for

McMaster Works, which was investigated in detail in this thesis,

is the use of scrap additions during teeming into the tundish in

o rder to red uce superhea t. Fur ther mod if ica t ions to the FASTP

software were made in order ta allO\<l the simulation of a moving

boundary 50 that a theoretical analys is of scrap melting

kinetics could be performed. Results of the simulations were

presented ln te,=ms of the dimensionless quantities Bi and Fo.

The use of this powerful technique allowed a general model to be

developed whereby any combinat ion of "effective" scrap radius

and convective heat transfer in the bath could be chosen and a

total rnelting time could be estimated. A value for the overall

convective heat transfer coefficient in the tundish pourbox was

determ ined by immer sing steel cylinders into the l iquid steel

for different durations of time, measuring their diameter after

withdrawal and calculating the corresponding dimensionless

qUdntities. An average value for the heat transfet coefficient

of about 16000 W/M 2 _OC was obtained. From this i t was

det~rmined that, from a scrap melting kinetic point of view,

estimating an average liquid steel residence time in the tundish

178

•

T

pourbox of 335 seconds, the steel scrap additions must be less

than 133 mm in order to avoid skulling ln the pourbox. Trial

heats were then performed whereby pleces of l02xl02mm billets,

readily available, were added into the pourbox. The trials

showed that th i scan be an effec t l ve wa y to reduce the super hea t

in the tundish. The major drawback was found to be in the

materials handling aspects. rf the addltions are kept

relatively small however, and the process automated, this

pr act i ce can be more acceptable to oper d t i ng per so nnel and can

at least be considered when producing the most critical grudes.

Liquid steel w.i.ll always carry sorne superheat as

it enters the mould. This superheat can be lowered and even

el iminated by adding Iron powder into the mould during casting.

The use of Iron powder as a source of supplemental cool ing can

provide the benefits of improving internal quality and

i ncreas ing casting speed, as was shown in the li terature survey

presented in Chapter 6 of this thE'sis. Trials were attempted dt

McMaster Works but only llmlted conclusions could be drawn due

to problems encountered with the feeding equlpment developed.

The two heats which were produced relat 'vely problem free

however, showed internal structure improvernents which were in

full agreement wi th what was observed by other rcsearchers.

\79

A reductlon ln the columnar zone Size and ln core

porosity was observed thus it can be said that based on these

limited results Improvements to the internaI quality of cast

billets are possible and further work on a larger scale is

warranted. The main obstacle will likely be in terms of

developing an adequate materials handling and delivery system

compatible with the Pollard inert gas shrouding system used at

McMaster Works and eventually a submerged ceramic tube shroudi.lg

sys tem.

Jn conclusion, Mct1aster Works is being asked to

produce more steel of better quality at a lower cast. In arder

to meet these requirements, engineering work is proceeding in

incorporating a captive lid system for the ladIes. Furthermore,

plans for adding a ladle furnace facility are also at the

engineering stage. Work is proceeding in the development of dam

and weir materlals for the tundish. rrhere are no immedlate

plans for making scrap additions to the tundish nor iron powder

additions to the mould but they may be consldered at sorne time

in the future, especially on the more critical grades of steel.

180

•

1

1

-----------------------------.

APPENDIX 1

FASTP (Faclhty for the Analysls of Systems ln Transport PhenOIllf'Ila) IS a softw.Hl'

package which was developed for the purpose of solving heat and ma~s transfpr probh'llI!-.

of relatlvely simple grometry lt is ba:t!d on the explicit fimtc dlff<,rencf' fornllllat.101I of thf'

Fourier general conductIOn equatlOn (1).

( 1 )

The formulation reqUires the sectIOn being analyzed he divided lOto discretf' nod('s and an

energy balance be performed on each node.

t 1 , ,

energy ln 1 • • , • • , • .. l 2 3 l-l 1 1+1

thus:

(Rate of energy ) (Rate of energy ) ( Rate of encrgy )

into no de i - out of node i + generated in node 1

or

where: VI = Volume of node i

PI = Density of node 1

CI = Heat capaclty of no de 1

e, = Dimensionless temperature at time t

e: = Value of 0, at t + llt

.6.t = Iterative lime in element

181

• t l~nl'lllY •

n-1 Il

(

Rat.e of em'rgy ) accumulation Hl

nod<, 1

(2)

( >l! L

•

·1 ",

q -- Rate of heat generatlOn

now,

(3)

(4)

where ~x = node spacing.

Comblning equation 2,3, and 4;

0' _ 0, = k.- 1-<,A'-1(8'-1 - O.)~t • P,C,V,ÂX.-l

(5)

ln order for equatlon 5 ta be solved, a stabllity criterion must be satlsfied. By ma-

n1pulating the terms of equatlOn 5, it can be found that the following cnterion must be

satisfied,

l (6)

Equations 5 and 6 represent respectlvely the exphC1t fimte dlfference formulation and its

correspond ing stabihty criterlOn for any mternal node 1 A simIlar balance must be con-

ducted for the surface nodes 1 and n, keeping in mInd the convention chosen that energy

flows ln the direction from 1 to n. Therefore, for surface no de l, the energy balance equates

ta the followmg,

(7)

" whcre q = Rate of heat generatlOn per unit area

11" - Smface area of nodal pomt. 1

A· =-- Surface art>a of nodal pOInt n

8a = Amblent temperature

a --:c Stt'fan-Boltzman constant = 5.663 X 10-8 W 1m2 - K4

€ -- [~mlsslvlty of surface of nodal point 1 (0 ta 1)

182

•

1 F = Vlew factor (0 to 1)

hl = Hcat transfer coefficient

Again, for mathcmatlcal stabllity, the followmg criterion must. be satu,tif'd,

(8)

and for surface node n;

= p"Cn Y.n ~~ --~~ (9) ".. 6t

the stability criterion beingi

( l 0)

If surface node n is a composite mat.erial, i e.

A B

• t t • n-l n l 2

then;

kn-l-mAn-l(On-l - On)~t _ kl_2A~OI - 02)~~ -\- q~ A. = Pnq~_!,-~(q_~_-_~,:,) 6Xn-l ~XI L':J.t

PtC l VI (0'1 - 0d +- ~t - (Il )

by incorporating into equation Il, the following substitution,

~ = m (or 6'1 = m) On lJ~

18 <

~---------------------------------------------------------- ------------

1 whcrc m - partItion coefficient

0' n

for stablllty;

k1-+ 2A 1 (mOn - (2)6t

6Xl(PnCnVn + mp1C1Vt)

(12)

(13)

The equatlOns just presented are the basis for the FASTP software These equations

arC' fllrther extcndcd to mass transfer by usmg analogous molecular properties instead of

thNmal propertlCs. The formulation of thcsc eqllations have also bcen expanded to incor-

pNate other fcatures ln the program Sllch as stlrring of a series of consecutive nodes, phase

changes and mtC'rboundary reslstance to provlde for non-Ideal contact between surfaces.

The followlng 18 a simple ex ample of a session wlth FASTP The problern solved IS

Illu~trated below.

2 () })= ïw /m - C

BRICK

What IS the temperature profile after 2 mmutes?

What is the temperature profile after 5 minutes?

What 18 the temperature profile after 10 rnmutes?

184

= 500°C

~

C

K

(Fixed dt time

2300 Kg/m 3 =

1000 0

= J/Kg- C

= 1.5 W/m-oC

= 0)

•

1

1

LJO , (''-1 "'J.N 1 1 0 ,,(II Vt (\ 1 \1\ J.l Th "N',I ~ K OR (21 MASS lRANS~ER PROBLEM •

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1 SI ( ME'ler9 Kil (l{JramlJ Spc- den l' loi", t·, 2-HI'lflSH Fppl :..b Ih. dpo ~ Il lili HR l-NH.R " l BRlllbll

Il) CYL 1 NDE.R OR t 2) SLAB (lR (J) '>l'llfRf ' l

ENTER THE NUMBER OF DI MENS IONS 1 MAX - '1

NUMBER OF BECllUNS IN EACIl DIMENSION' IHAX 51 DIMENS (ON 1

IPIS 15 A SUMMARY OF 'fOUR ENTRIES SO FAR

VOU ARE SOLVING A HEAT lRANSHR f'ROfll!'M

IN 5 1 UNITS

FOR A SlAB

IN DIHFNSIONtSI

DI MENS ION 1 HAS 1 SECTION<;

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PaulJe

Plealle press (reluru> ta continue

DO YOU WANT Ta (Il CORRECT ERRORS OR 121 f'ROCF ED • 2

<1> ENTER lHICKNESS OF EACH SElTION - ANU -NUMBER OF NOUES IJ TO 201 IN E .... CH SECTION FOR FACH DIHfNSION

SECTION 1 DIHfNSIUN 1 0 01 7

YOU ARE DEAL! NG W 1 TH SECT No 1 OF DI H No

ENTER 1 -NORMAL 2-ARITHHETIC J-GEOMETRIC 4-COMBINATION ~-YOlJr 0""1

YOIJR PROBlEH SETS UP .... S !'OLLOWS

NODE NODE. RADIUS OF NUDI': APEA vrJl UME No POSITION BOUND .... RY SPAC 1 HG

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Paut;{' Pl~B"le OTPOl5Cj 'rplurn.1 ta continue

185

l , l 'fI Il' 'vi AU T,r, l ,,. '1 f flll' ~ tH ,\11 rH Ir 'J .... !. 1'f."JlffIJJIIHII~I~ ,H'" L

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h HfllI MMH I.f/"I'I' (of ['fN' Iii Il! AI '''r' no iOU WANI TO [NItR Atl[) 1 OR CHAN' E ?

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'7' ~,JARIING TfMI'fRA1URE PROFILfb

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186

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DO YOII IoIAtll TO MAY~ "UMf CIIANt,H, II' Illf INl'ttT [l ... I ... < n t DO YOII IoIMll 10 FYfCIIIF IHl Fktl',PAM '1) no 1'011 IoIMH 10 'dur rXF(ïlll{lN (.' UO YOU IoI ... NT ... SUMM"'R~ m ~ OliR 1 Nf'l (1 l'''' 1" 1 N MF M(lR, (II

••• ~ ••••••• *~ •••••••••••••••••••••••••

INI UT U"'T'" IN HEMORr

._ ....•.•.. ~ ......•...................

5i r liON THICKNESS and NUMBlH OF ~OUF5

SEC r No 1 DI H No Ulon

flOUND ... RY SURFALE CotWI flONf,

OIM No DIM No

1 EMf'n,,, 1 1I1'E <IF ~(lPI"', E. -rEHPEflHUhE l'~ 'iIPf-,t.'E N

FOR SUR~ACE 1 OF li 1 M tlo HFAT TRANSFER COlf-FICIENI AHAIENr TFMPFPATURF NO RADI,t.TION CON51m "ED

7 l'un 25 00

-IOOU ou ~lln 00

HE AT TPANS OF DIH No

C(lt'rF RADIATION AND AMBIENI TEMP FOR SIJH~A' E N 1 ARE NU T REl EV ANl

NO PHASE O,"NGES ARE 1 NVOL VEO

THERE ARE NO INTER80UNDARY RESIS1ANCE.S

THERMAL PROPERTIES

SECT OIM ST"TE DFNSITY Co y

4 23UO 0000 IOVO 000 1 <")0

bTARTING TEMPERATURES

SECT No DIH No ISTART lEHF' 25 00 SI ... rE 4 ,; r ... filS - "HALE. f- 1 XE [)

THF_RE 15 NO HEAT GENERATION

AN~LYSIS IS DONE BASED ON THE NOIlE 1 Nl ERACT 1 ON SOLUT 1 ON

THERE ARE NO STIRRED SECTIONS

••• DO YOll IoIANT 10 MAKE SOME CH!'N<;E.S If} rHE IN~UT DA!.\ Ir) Df} 1'01' IoIAtIT Ta EXECliTE litE f ROfjRAM 'l'

••• DO YOIJ IoIAtJT TO ~~IOP EXEClll 1011 (n ••• 00 YOU WANT A SUMMAR) OF YOUR 1 NPU r DA TA 1 N MlMOPY 1 1 )

187