Physico chemial Aspects of Steelmaking Processes By E TT URKDOGAN Synopsis.pdf

8/17/2019 Physico chemial Aspects of Steelmaking Processes By E TT URKDOGAN Synopsis.pdf

1/21

Physicochemical Aspects of

Steelmaking Processes*

Reactions in Ironmaking and

By E. T. T URKDOGAN**

Synopsis

This is a critical review of the physical chemistry of selected aspects of

ironmaking and steelmakingprocesses. The equilibrium data from gas-

slag-metal reactions are presented in a form suitable for easy application

in the study of the state of reactions in ironmaking and steelmaking. Plant

data indicate departures rom equilibrium or most reactions in the blast-

furnace bosh and hearth. In steelmaking, most reactions are close to equi-

librium at the time of furnace tapping. Comments are made on reaction

equilibria in relation to the ladle refining of hot metal. A summary is

made of new equilibrium data for desulfurization of liquid steel with cal-

cium aluminate slags. Various examples are given of the role of vapor

species n pyrometallurgicalprocesses, e.g., silicon- and sulfur-bearing spe-

cies and alkali cyanide vapors in the blast furnace, causes of fume emission

during furnace tapping, and consequences f fume emission in coal gasifica-

tion using liquid iron as the reaction medium.

I. Introduction

tudies of the physical chemistry of gas-slag-metal

reactions and measurements of the physicochemical

and thermochemical properties of materials pertinent

to ironmaking and steelmaking processes have gone

through many interesting stages of evolution during

the past six decades or more. The first textbook writ-

ten on the physical chemistry of steelmaking was the

classical work by Schenck, Einfuhrung in die Phys-

ikalische Chemie der Eisenhuttenprozesse, published

in the years 1932 and 1934. Belated English transla-

tion of this book in 19451)was of much value to steel-

makers in many countries and stimulated the birth of

a new era of research on the physical chemistry of

metallurgical reactions. The symposium on physical

chemistry organized by the Faraday Society in 19482)

set the main theme of research for the subsequent two

decades on the fundamentals of pyrometallurgical pro-

cesses.

Accumulated knowledge of the physical chemistry

of gas-slag-metal reactions and the physicochemical

properties of solid and liquid metallic solutions, mol-

ten slaps, and refractory materials have been critically

assessed and compiled by Kubaschewski and Alcock,3~

Richardson,4) and Turkdogan.5's) These reference

books are used for the background scientific and tech-

nical information in the preparation of the present

paper.

Upon kind invitation by the Iron and Steel Institute

of Japan, a critical assessment is presented in this pa-

per of some aspects of the present state of knowledge

of the physicochemical aspects of ironmaking and

steelmaking processes. For the paper to be technical-

ly comprehensive, the subject of discussion is confined

to a few selected topics of process metallurgy in iron-

making and steelmaking that are of current interest:

(1) the state of reactions in the blast furnace and the

converter in relation to the gas-slag-metal reaction

equilibria; (2) ladle refining; and (3) the role of vapor

species in metallurgical reactions, including coal

gasification in liquid iron bath.

II. Gas-Slag-Metal Reaction Equilibria

1. GeneralConsiderations

as-slag and metal-slag reactions are redox reac-

tions which are accompanied by exchange of electrons

between the reacting species as they are transferred to

and from nonpolar gaseous or metallic phase and ionic

molten slag. For example, the transfer of sulfur from

metal to slag is a reduction reaction involving ex-

change of electrons between sulfur and oxygen atoms:

-E-(O2-)= 0+(S2) ..................(1)

where underscore indicates solute in solution in the

metal and ( ) species dissolved in the slag. In most

cases, the transfer of a reactant from metal to slag is

an oxidation process, such as

+ 5 0 + 3 (02-) = (POr) ...............(2)2

Although the reactions as formulated above are

conceptually correct, their equilibrium constants can-

not be evaluated because the thermodynamic activity

or activity coefficient of an individual ion cannot

be determined experimentally. However, numerous

attempts have been made in the past, following the

work of Herasymenko,7'8) to represent the equilib-

rium constants of slag-metal reactions in terms of

assumed ionic concentrations.9'1°) Although the deri-

vation of reaction equilibrium relations from the ionic

structural models are of theoretical interest, their ap-

plication to gas-slag-metal reaction equilibria in mul=

ticomponent systems has not been particularly reward-

ing when compared with the conventional methods of

representing the reaction equilibria in terms' of the

measured thermodynamic quantities.

he solution models of slaps, either constitutional

or structural, have been much exploited in pursuit of

a rationale for the composition dependence of the ac-

tivity of constituent oxides in slaps. No matter what

major differences there might be between assumed

solution models, they all yield activity vs, composition

relations in reasonable agreement with the same set

of data for binary and ternary silicate melts, because

of the empirically adjustable parameters involved in

*

**

Received November 14, 1983.

Technical Center, United States Steel Corporation, Monroeville, PA 15146, U.S.A.

Review

(591)


2/21

(592) Transactions ISIJ, Vol. 24, 1984

the formulations. However, although the solution

models may be fictitious, the formalisms derived there-

from are helpful in the extrapolation of data to other

composition ranges. For example, Ban-ya et a1.11'12)

have shown that the composition dependence of the

activity coefficient of ferrous oxide in complex silicate

and phosphate melts can be represented by an em-

pirical formalism based on the regular solution model

for polymeric melts proposed by Lumsden.l3~

ecause of lack of activity data for complex slaps

and limited applicability of empirical solution models,

the slag-metal equilibrium relations are often repre-

sented in terms of the concentrations of reactants and

products by adopting the mass-action law. The iso-

thermal equilibrium constant K thus expressed varies

with the slag composition, on account of omission of

the activity coefficients of oxide reactants and prod-

ucts in the expression of the equilibrium constant K.

he mass-action law has been applied in the past

in a variety of ways to represent the state of slag-

metal reaction equilibria. The apparent disparities

sometimes seen in the experimental data of different

investigators on slag-metal equilibria are often at-

tributable to the format of representing the state of

equilibrium for slags of varying compositions. In an

attempt to develop a relatively simple rationale, Turk-

dogans~ proposed a method of application of phase

rule to describe the slag-metal reaction equilibria as

discussed below.

et us consider in a general way the equilibrium

distribution of an element X between slag and metal.

X (in metal) _ (X) (in slag)

(3)

The independent state variables which determine the

solubility of X in the slag (in a simple or complex

ionic form) coexistent with the metallic phase may be

deduced by invoking the Gibbs phase rule:

p+f =c+2

(4)

where, p : the number of phases

: the number of degrees of freedom

: the number of components of the system.

For the phosphorus reaction between liquid iron and

slag, the system may be described by the following six

components: Fe, P, 0, (CaO), (SiO2) and (MO),

where MO represents MgO plus other minor slag

components, e.g., MnO, A12O3,CaF2, ... For isobaric

and isothermal conditions, we now have from Eq. (4)

p+ f = 6. With the assumption that phosphorus has

virtually no effect on its activity coefficient in the

metal and slag phase, the equilibrium distribution

ratio (%P)/[%P] between slag and metal may be de-

scribed by the following functional relations, depend-

ing on the number of phases present* :

2 phases : slag and metal,

L %P]p,T

(5)

where, V

3 phases:

: the basicity.

gO-saturated slag and metal,

(%P)

4 phases

[%P]

P,T

{[%O] , (%SiO2), V} (6)

: MgO- and CaO-saturated slag and metal,

[(%P) [

[%P]

P,T

~b{[%o] , (%5102)} (7)

Although qualitative, the slag basicity defined by

the concentration ratio (oxides of network modifiers)/

(oxides of network formers) gives a general indication

of the extent of depolymerization of the melt, hence

is a simple indicator of the trend in changes of oxide

activities with composition. Consequently, the equi-

librium constants for slag-metal reactions expressed

in terms of concentrations of slag components are often

simple functions of slag basicity, at least for composi-

tion ranges of practical interest.

he concept of slag basicity appears to have been

first introduced by Herty for steelmaking slags in the

mid 1920's; he proposed a simple concentration ratio,

so-called the V ratio:

%CaO

The definition of slag basicity is somewhat arbitrary.

For example, if we were to assume that in steelmaking-

type slags the concentrations of CaO and MgO are

equivalent on a molar basis, and similarly on a molar

basis 1 2 P2O5 is equivalent to SiO2j in terms of mass

concentrations of oxides, the slag basicity B may be

defined by the ratio:

B = %CaO+1.4(%MgO) (9)

...............

%SiO2+0.84(%P2O5)

The basic steelmaking slags are saturated with mag-

nesia and dicalcium (magnesium) silicate. For such

slaps containing less than about 3 % P2O5, the basicity

B given by the above expression is essentially a linear

function of the V ratio with a proportionality factor

of about 1.17.

B =1.17 %CaO .....................(10)5

102

or blast furnace slags, which contain 10 to 20 %

A12O3, he basicity is often represented by the mass

concentration ratio:

%CaO + %MgO 11

2. Ironmaking Reactions

he state of equilibrium in metallurgical reactions

involving three or more phases cannot be evaluated

accurately without giving due consideration to inter-

relations among different reactions occurring simul-

taneously. In conventional laboratory experiments

with blast-furnace-type systems at unit activities of

carbon and carbon monoxide, many hours of reaction

* It should be noted that the argument presented here is the corrected version of that given by the author in Ref

. 6).

Review


3/21

time are needed to reach the equilibrium distribution

of silicon, manganese, and sulfur between slag and

metal to satisfy the three-phase equilibria in the over-

all reactions:

Si02)+2C = Si+2C0 .....................(12)

CaO)+S+C = (CaS)+CO ............(14)

The studies made by Turkdogan et a1.14) ave re-

vealed that a partial equilibrium in such three-phase

systems could be realized within about 1-hr reaction

time using small amounts of slag and metal samples.

They obtained the equilibrium relations for the fol-

lowing three coupled reactions in graphite-saturated

melts and blast-furnace-type slags containing 10 to

20 % A1203.

(MnO)+Si = (Si02)+2Mn ...............(15)

CaO)+ 1 Si+S = (CaS)+ 1(SiO2) ......(16)

2 2

CaO)+Mn+S = (CaS)+(MnO) .........(17)

From their own equilibrium measurements and

those of Filer and Darken,15) Turkdogan et a1.14)de-

rived the following equilibrium relation for reaction

(15)

%Mn] 2 (%Si02) K

M11s1= %Mn0) [%Si]

log KMl,Si 2.8B'-l.16

where, B':

400 Withinhe

effect of to

ogethe

Darken,15)

equilibrium Stobo,17i

was represented by

%S) (%SiO2) lie 1

b S- -[%SJ - [%Si} ) (%C

og K515 6327 -4.43+1.4B'

the slag basicity as given by Eq. (11)

o 1 600 °C investigated, the range

mperature on KMnslcould not be detected.

with the experimental data of Filer and r

Hatch and Chipman,ls) and Taylor and

elation for reaction (16)

a0)

(19)14)

Equations (18) and (19) give the following equi-

librium relation for the manganese-sulfurcoupled re-

action (17).

%S) (%MnO) 1

Mils -[%S] [Mn] (%Ca0)

20

327

og KMfS T _3

In most blast furnace operations, the sum of the

concentrations of MgO, MnO, and A1203 (=MO)

does not change much. Also, in melts saturated with

graphite in equilibrium with a fixed partial pressure

of carbon monoxide, the activity of oxygen remains

constant for a given temperature. For these limiting

conditions, the functional relation for the slag-metal

equilibrium is simplified to

Transactions ISIJ, Vol.

24, 1984

(593)

(%X)

_

[%X]

0 {B'}

(21)

The equilibrium distributions of manganese and

silicon between slag and metal shown by dotted curves

in Fig. 1 are consistent with the functional relation

(21) deduced from the general thermodynamic con-siderations. Since CaO is one of the reactants as

given by Eq. (14), the sulfur distribution ratio has to

be a function of both the slag basicity B' and the con-

centration of CaO in the slag; the dotted curves in

Fig. 1 for (%S)/[%S] are for 40 % CaO in the slag.

The dotted equilibrium curves in Fig. 1 for graphite-

saturated melts at 1 bar CO pressure were derivedl4~

from the appropriate free energy data, oxide activities,

and experimentally determined equilibrium relations

given by Eqs. (18) and (19).

n his 1978 Howe Memorial Lecture, Turkdoganls)

discussed the state of slag-metal reactions in the blast

furnace hearth, based on the analysis of the daily aver-

age plant data for various blast furnaces of U.S. Steel

Corporation and those reported by Okabe et al.l 9> or

Chiba Works of Kawasaki Steel Corporation. The

data analyzed are typical for most blast furnace opera-

tions. The melt temperatures at tap are in the range

1 450 to 1 550 °C, mostly in the range 1 475 to 1525

°C

. The average slag compositions in mass percent

are in the range : 38 to 44 % CaO, 8 to 10 % MgO,

34 to 38 % Si02, 10 to 12 % A1203, 0.5 to 1.0 %

MnO, 1 to 2 % S, 0.1 to 0.6 % K20,


4/21

(594)

Transactions ISIJ, Vol. 24, 1984

area shown in Fig. 2 (top diagram) ; the equilibrium

lines (dotted) are from Fq. (20)) for 40 % CaO. How-

ever, for a given blast furnace there is much less scatter

in the data which show the expected linear relation

between (%S)/[%S] and [%Mn]/(%MnO).

alues of KMnsi rom plant data scattered within

the shaded area in Fig. 2 (bottom diagram) are be-

low the equilibrium line (dotted). This observation

from the plant data may indicate that as the metal

droplets pass through the slag layer in the blast fur-

nace, the direction of the reaction might be

(MnO)+Si -> 2Mn+(Si02) ............(22)

Such an intuitive deduction, however, may not be cor-

rect. Observed departure from equilibrium for re-

action (15) may be attributed to competition between

FeO and MnO in the slag for oxidation of silicon. It

is a generally accepted view that with a blast-furnace

burden of low basicity, the residual iron oxide in the

melting zone of the bosh is higher than with a burden

of high basicity. In slags of low basicity, the higher

concentration of iron oxide in the upper part of the

slag layer would be in competition with MnO for re-

action with both silicon and carbon. Consequently,

departure from equilibrium for reactions (13) and (14)

is expected to be greater in slags of low basicity. The

trend seeing from the plant data in Figs. 1 and 2 sub-

stantiate the deductions made from the foregoing argu-

ment. A study of plant data from a ferromanganese

blast furnace20~also show departures from equilibrium

for reaction (15) in a manner similar to that for the

iron blast furnace.

elve et a1.21~ ave studied the titanium-silicon re-

action in graphite-saturated melts using blast-furnace-

type slags

TiO2)+Si = (SiO2)+Ti ...............(23)

Their experimental results, for temperatures of 1 500

and 1 600 °C and basicities of 1 to 2, may be sum-

marized by the following equation.

KTiSi [%Ti] _ (%S102)%Ti0

2) [%SiJ 24

Ca0

ogKTisi 0.46 - -Si0

2 +0.39

From the blast-furnace-plant data compiled by Delve

et al.21j and by Hess et al.,22~ t is noted that the values

of KTisi are two to three times lower than those given

by Fq. (24), which presumably represents the equi-

librium relation. Observed departure from equilib-

rium for reaction (23) in the blast furnace hearth is

similar to the state of silicon-manganese reaction (15)

in iron- and ferromanganese-blast furnaces.

3. SteelmakingReactions

any equilibrium measurements have been made

of reactions between liquid iron alloys and slaps rele-

vant to steelmaking processes. The equilibrium rela-

Fig. 1.

Daily average blast furnace data are compared

with equilibrium relations for slag/metal

distribution of silicon, manganese and sulfur

in melts saturated with graphite at 1 bar pressure

of CO.

Review


5/21

Transactions Isu, Vol.

24, 1984

(595)

tions for the key reactions given below are taken from

the author's critical assessment6r of the experimental

data of various origin.

he recommended value of the equilibrium con-

stant for the carbon-oxygen reaction in liquid iron is

that obtained by Fuwa and Chipman23} :

0=C+0

K[ac] [ao]

co ............(25)

168

og K = - T --2.076

where Pco is in bar and the activities ac and ao are

taken equivalent to [%C] and [%O], respectively, at

low solute concentrations. It should be noted that

the ratio C02/CO in the gas increases with decreasing

concentration of carbon. For a total pressure of

Pco2+Pco=1 bar, the CO2+C=2C0 equilibrium

gives for 1 600 °C : 0.05 bar CO2 for 0.03 % C and

0.14 bar CO2 for 0.01 % C in iron.

rom the known oxygen solubility in liquid iron

and the iron oxide activities in the quaternary melts

CaO-MgO-FeO~-SiO2, the following relation is ob-

tained for CaO (MgO)-saturated simple stags in equi-

librium with iron-carbon melts (< 1 % C) at 1 600

°C and CO

2+CO=1 bar.

[%C](%FeOt)-0.33 (26)

where the subscript t indicates the total iron in the

slag as oxides represented as FeO, i.e., %FeOt=

%FeO+0.9 X %Fe2O3.

or the slag basicity V> 1, the equilibrium relation

for the manganese reaction is reduced to a simple form

for temperatures of 1 550 to 1 700 °C.

Mn+(FeO) = (MnO)+Fe

Mn~e= (%MnO)%M

n](%FeOt)

CaO

- -

%SiO2KMnFe6

(27)

or s gs CaO and Ca4P2O9, the

following equilibrium relation was obtained by Ban-

ya and Matoba24~ for the phosphorus oxidation at low

concentrations in iron.

4CaO(s)+2P+50 = Ca4P2O9(s)

1 K

_

[%P]2[~/

pO]5 ......(28)

log Ky = 75-660660-31.77

The temperature dependence of KP determined in an

earlier study by Fischer and vom Fnde25~ differs con-

siderably from the above equation. However, for the

temperature range 1 530 to 1 585 °C investigated by

Ban-ya and Matoba,24~ the values of KP in these in-

Fig. 2.

Daily average blast furnace data are compared with

equilibrium relations for manganese-sulfur and

manganese-silicon coupled reactions in melts satu-

rated with graphite.

Review


6/21

[596)

Transactions

ISIJ, Vol. 24, 1984

dependent determinations agree within a factor of

about 1.5. It should be noted that the enthalpy

change accompanying reaction (28) estimated from

the thermochemical data is closer to that given by the

above equation.

arious formalisms have been used in the past to

represent the slag-metal equilibrium with respect to

the phosphorus reaction. 26-29)Using the slag-metal

equilibrium data and thermochemical data available

in the early 1950's,Pearson and Turkdogan30) erived

an empirical equation for the effect of temperature

and slag composition on the activity coefficient of

P205 in the slag, relative to the hypothetical pure

liquid P205. A similar formalism was adopted by

Suito et a1.31)o interpret their recent results for phos-

phorus distribution between liquid iron and slagssatu-

rated with magnesia. These different formalisms of

varying complexity do not fit all the experimental data

on the phosphorus reaction.

new formalismproposed recently by the author°

is based on the functional relations given by Eqs. (5)

to (7). Using all the available experimental data,

including the recent work of Suito et al.,31) t was

found that for slaps not saturated with lime or dical-

cium silicate, the following functional relation de-

scribed the data reasonably well*.

%P205)

PS - -~%P~ (1-%Si02) = {%FeOt, V}

.......................(29)

For double saturation of the slag with lime and mag-

nesia, when the basicity V>2.5, the isothermal equi-

librium relation KPS s a function % FeOt and essen-

tially independent of basicity, in accord with the

derivation of the functional relation (7).

or slags of the system CaO-FeO-Si02 P205 satu-

rated with either CaO or Ca4P209, the phosphorus

distribution ratio would be a function of both % FeO~

and % Si02. However, the data of Fischer and vom

Ende25) and of Knuppel and Oeters28) in Fig. 3 for

these singly saturated phosphate melts may be repre-

sented by a single curve. The scatter in the data

masks the apparently small effect of % Si02 on the

relation:

%P205)

-' °'F eO

~} ...............(30)- [%P}

for singly saturated slags containing less than 8 %

S102.

From the analysis of available experimental data,

the relations in Fig. 4 were obtained for steelmaking-

type slags saturated with both dicalcium silicate and

magnesio-wustite, for general guidance in evaluating

the state of the phosphorus reaction in steelmaking.

Fig

. 3. Equilibrium distribution of

xide or calcium phosphate.

phosphorus

between

slag and iron

for

saturation of slag with calcium

Fig. 4. Variation of equilibrium relation KPS with tempera-

ture and iron oxide content of saturated steelmak-

ing-type slags. After Turkdogan.s~

* All the experimental data considered in this evaluation of the equilibrium relations are for slags saturated with magnesia; therefore

, the

ndependent variables used in Eq. (2) are those for the functional relation in Eq. (6) with % Fe01 as a measure of [% 0].

Review


7/21

The representation of the experimental data in the

manner described did not show any effect of CaF2 on

the phosphorus distribution ratio, indicating that CaF2

in moderate proportions is equivalent CaO. This

subject is discussed further later in the paper in rela-

tion to the ladle refining of hot metal.

or the low oxygen potentials sustained in the pres-

ence of metallic iron, the sulfur dissolves in the slag

as a sulfide ion. For the slag-metal reaction (1), the

equilibrium relation may be represented by

KS- (~°-)(%FeOt) ..................(31)

where, K5 a function of temperature and slag com-

osition.

Various schemes were proposed in the past to repre-

sent the variation of Ks with slag composition. As

suggested by the author32~ in an earlier study of the

available data, the simplest form of the composition

dependence of KS is that shown in Fig. 5 for wide

ranges of slag compositions and temperatures of 1 530

to 1 730 °C; the relation is well represented by a single

curve for the concentration of acidic oxides above

5 %. Loscher33> studied the sulfur reaction using

slaps saturated with tricalcium phosphate. For melts

containing about 42 % P2O5 and 3 to 6 % FeOt, re-

mainder being CaO, the KS is in the range 2 to 4

which is in general accord with other data in Fig. 5.

The greater scatter of the data points at low concen-

trations of acidic oxides is attributed to the greater

Transactions ISIJ, Vol. 24, 1984 ( 597)

sensitivity of KS at very high basicities to the relative

concentrations of basic oxides. For CaO-FeOt melts,

the curve intercepts the ordinate at Ks 540.

lthough the steelmaking methods and practices

differ to varying degrees, the same fundamental prin-

ciples of gas-slag-metal reactions apply to them all.

In keeping with the theme of the present discourse,

the deliberation here is confined to the discussion of

the state of reactions during steel refining as affected

by slag composition. For low phosphorus in the fur-

nace charge, i.e.,


8/21

(598)

Transactions ISII,

Vol.

24,

1984

blowing (Q -BOP) steelmaking practices. Variations

in composition within the shaded areas suggest that

some basic steelmaking slags may not be saturated

with dicalcium (magnesium) silicate, despite the fact

that BOF and Q; BOP slags always contain entrained

particles of undissolved lime. The Q- BOP slags usu-

ally contain more GaO than BOF slaps. For the

range 1 550 to 1 650 °G, the tap temperature has no

detectable effect on the distribution of data points in

this composition diagram.

he dotted line in each diagram is for dicalcium

(magnesium) silicate saturation in the simple GaO-

6%MgO-Fe0~-SiO2 melts in equilibrium with liquid

iron at 1 600 °G. There is a pronounced shift in the

position of the univariant equilibrium in the composi-

tion diagram for the saturated-steelmaking slags to

higher concentrations of GaO and to the correspond-

ing lower concentrations of Si02 as compared to the

quaternary system GaO-6%MgO-FeOt-SiO2 in equi-

librium with liquid iron. This change in the melt

composition to higher concentrations of lime is due to

the presence of P205, MnO, Fe203, GaF2, and small

amounts of other oxides which shift the saturation iso-

therm to higher concentrations of GaO. Another

rationale is that the slag basicity changes from low to

high levels as the iron oxide content of the saturated

slag increases, e.g., the V ratio is 2.0 to 2.5 for slags

containing 4 to 6 % FeOt and increases to 3.4 to 3.6

for slaps containing 28 to 30 % FeOt. That is, in

saturated slags the basicity is fixed by the concentra-

tion of iron oxide, hence by the concentration of car-

bon in the steel.

he oxygen activity in liquid steel measured with

the EMF-oxygen sensor gives the concentration of oxy-

gen in solution in the steel (for low-alloy plain carbon

steels). The BOF and Q; BOP data for concent-

rations of oxygen in steel and total iron oxide in slag

at tap, for temperatures of 1 550 and 1 650 °G and

for carbon contents in the range 0.01 to 0.8 %, are

distributed within the shaded area in Fig. 7(A).

The dotted curve is calculated from the data on the

solubility of oxygen in liquid iron and the activity

of iron oxide36> along the lime-saturation isotherm

for relatively simple slaps GaO(MgO)-SiO2-FeOt.

The position and shape of the shaded area (where the

plant data are distributed), relative to the dotted

line for the simpler saturated slaps, indicate that the

activity coefficient of iron oxide in steelmaking slags

decreases with increasing concentration of iron oxide

and with the corresponding increase in slag basicity.

rom the compositions of slag and metal samples

taken at tap in BOF and Q -BOP steelmaking at tem-

peratures of 1 590 to 1 610 °G, a plot is made in Fig.

7(B) . of the product {(%MnO)/[%Mn]} (%GaO/

%Si02) against the concentration of iron oxide. The

broken line drawn with the slope of six represents the

slag-metal equilibrium for the manganese reaction as

given in Eq. (27). It is self evident that there is close

approach to the slag-metal equilibrium with respect

to the manganese reaction in oxygen-steelmaking pro-

cesses, over wide ranges of carbon content of steel at

tap.

In oxygen+lime bottom blowing as in Q-BOP, the

slag formation begins in the melt near the tuyere zone

where the oxides of iron, manganese, silicon, and phos-

phorus react rapidly with the lime powder. As the

Fig. 6. Composition of steelmaking slags containing


9/21


(599)

slag particles traverse the bath with CO bubbles, they

react with carbon. Because of the secondary reaction

with carbon in the bath, the concentrations of iron

oxides and manganese oxide in Q- BOP slaps are

lower than those in BOF (or BOP) slaps. That is,

for a given concentration of carbon in the bath, the

metal/slag manganese-distribution ratio in Q- BOP is

higher than in BOF.

he plant data plotted in Fig. 8 show the relation

between the concentrations of oxygen and carbon in

steel at tap. With the exception of low-carbon melts

in the Q -BOP, the oxygen contents of steel at all car-

bon levels for BOF and Q-BOP steelmaking are close

to the equilibrium values for a pressure of 1.5 bar CO

which is an average total gas pressure in the bath.

The lower oxygen contents of steel in Q-BOP at low

carbon levels suggests that oxidation of CO to CO2

in the bath is well above the equilibrium values for

the reaction;

CO2+C = 2CO .....................(32)

That is, when the rate of oxygen-bottom blowing ex-

ceeds the rate of consumption of oxygen by carbon,

manganese, and iron, the excess oxygen will oxidize

the CO generated in the bath near the tuyere zone to

C02, thus diluting CO and allowing the decarburiza-

tion to proceed at lower levels of residual oxygen in

the steel.

n the basis of the foregoing process analysis, which

was discussed elsewhere5,6~ n more detail, certain ad-

vantages are anticipated in the use of argon+oxygen

mixtures in the later stages of refining of steel in both

the top and bottom blowing practices, particularly for

steels containing less than 0.02 % C. In fact, these

anticipated advantages have been realized in recent

developments of combined top and bottom blowing

practices in steelmaking, e.g., LBE and K-BOP.

any previous studies of plant data have shown

that at the time of furnace tapping in all types of steel-

making processes, the sulfur-distribution ratio between

slag and metal is close to the equilibrium relation

shown in Fig. 5.

he BOF and Q-BOP plant data for temperatures

of about 1 600 °C are plotted in Fig. 9(A) showing

the variation of log K 5 with the total iron oxide con-

tent of the slag. For low P205 contents in slags satu-

rated with both dicalcium silicate and magnesia, the

concentrations of iron oxide and silica or lime in the

slag are interrelated. Therefore, for doubly saturated

slaps, the phosphorus-distribution ratio alone will be

a single function of the concentration of iron oxide.

The equilibrium relation for doubly saturated, low-

phosphorus slags is shown by the dotted curve in Fig.

9(B). The points for Q; BOP are scattered about the

Fig. 7(A).

Oxygen contents of steel

content of slag at tap.

is related to iron oxide

Fig. 7(B).

Equilibrium relation for

(- - - -) is compared with

Turkdogan.s~

manganese

plant data.

reaction

After

Review


10/21

(600)

Transactions ISIJ,

Vol. 24,

1964

dotted curve representing the equilibrium relation.

The data for BOF, for slags containing more than

about 15 % Fe01, show departures from equilibrium

in favor of high phosphorus in the steel. The same

plant data are plotted in Fig. 10 showing the variation

of the phosphorus-distribution ratio with the carbon

content of the steel at tap. For steels containing 0.03

to 0.05 % C at turndown, the slag/metal phosphorus

distribution ratio for Q -BOP is about twice that for

BOF. Furthermore, lower levels of carbon attainable

by oxygen-bottom blowing brings about more exten-

sive dephosphorization of the metal.

Fig. 8.

Relation between

carbon in liquid

blowing practices.

the concentrations of oxygen and

steel for oxygen-top and bottom

After Turkdogan.6~

Fig. 9.

State of phosphorus reaction at time of furnace tap-

ping in BOF and Q -BOP for temperatures of about

1 600 °C is compared to the equilibrium relation

(- - - -) for saturated slags. After Turkdogan.6

Review


11/21


(601)

As was noted in previous studies,5,6~ he analysis of

the plant data substantiates the view that at the time

of furnace tapping there is close approach to slag-

metal equilibrium for most reactions. Therefore, the

equilibrium data for slag-metal reactions in conjunc-

tion with plant data will provide a reliable basis for

static and dynamic process control in steelmaking,

with certain adjustments in the charge control model

to accommodate small differences in melt-shop prac-

tices.

III. Ladle Refining

rior to the 1970's, the ladle treatment of hot metal

was confined almost entirely to desulfurization when

it was considered necessary. In the case of liquid

steel, the ladle refining has been practiced for a much

longer time. However, prior to the 1950's, the ladle

refining of steel meant deoxidation, recarburization,

resulfurization for free machinability of steel and mod-

erate additions of alloying elements. The advent of

continuous casting and an increase in customer de-

mand for high-strength, low-alloy steels with improved

transverse mechanical properties made it necessary to

develop increasingly more sophisticated methods of

treating steel in the ladle for deoxidation, desulfuriza-

tion, degassing, and shape control of inclusions. In

recent years, there has been a growing demand for

cleaner steels, i.e., steels containing very low concen-

trations of nonmetallic inclusions, low sulfur steels

(< 10 ppm S), alloy steels low in interstitial elements

H, N, C, 0, and S, and for other quality steels for

special applications. Such high levels of steel refine-

ment can be achieved only by a suitable method of

ladle treatment.

ince the late 1970's, there has been a rapid growth

of technology for ladle refining of both hot metal and

steel, the so-called secondary steelmaking or ladle

metallurgy. There is an overwhelming volume of

publications of research and development on ladle

metallurgy. Regretably, many of the publications,

both from the academia and the industry, are on re-

petitive work. A critical assessment of the accumu-

lated knowledge and technical information on ladle

metallurgy in a comprehensive, yet in a condensed

form, would be a worthwhile undertaking. In the

present discussion of a few selected areas of ladle re-

fining, with particular emphasis on slag-metal reac-

tions, no attempt is made to cite in a chronological

order the large volume of repetitive publications on

the subject. References to almost all the publica-

tions related to ladle metallurgy can be found in the

proceedings of three series of symposia, Scaninject ,

held in Lulea, Sweden, in 1977, 1980, and 1983,3'-39)

and in other publications cited in this paper.

1. Rate of Refining n Stirred Melts

o enhance the rate of refining of steel with the

overlaying slag, with the injected powder reactants

and the flotation of reaction products out of the steel

bath, various methods of stirring have been developed

in the steel industry; electromagnetic, mechanical,

and argon injection. References may be made to re-

Fig. 10.

Phosphorus distribution

bon content of steel at

and Q-BOP steelmaking.

ratio

tap

is related to the car-

(4600°C) in BOF

Review


12/21

(602)


view papers40> on the practical merits of these pro-

cesses. Reference may be made also to Szekely41 for

studies of better understanding of the fluidynamics of

mixing in stirred melts through mathematical model-

ing and computer-aided calculations. For the pres-

ent, it will suffice to state that the experimental and

plant studies have shown that the rate of slag-metal

reactions or flotation of reaction products in stirred

melts can be represented by an equation of the form42~ :

Ct = C0 exp (-kt)

(33)

where, C0: the initial concentration of reactant in

he steel

t : the concentration at time t

: the apparent volumetric rate constant for

given type and intensity of stirring.

he mixing time, z, to achieve about 95 % homo-

genization in stirred melts is a good measure of the

rate of ladle refining. From measurements of the rate

of solute homogenization in 50 t argon stirred melts in

the ladle, 50 t ASEA-SKF, 200 t RH and 65 kg water

model experiments, Nakanishi et a1.43~erived the fol-

lowing empirical relation for mixing time z as a func-

tion of the rate of dissipation of energy, ~, in the

stirred melt.

r = (6O0±100)°°

(34)

where z is in second and in W/t liquid steel. For

gas stirred melts, is that due essentially to the buoy-

ancy energy of the injected gas, the contributions of

the kinetic energy of the gas, and the thermal expan-

sion of gas to the circulating flow are small enough to

be neglected.

= 6.18 X10-3VT In 1+10pg5~ .........(35}

W P

where, V: the gas flow rate (nm3/min)

: the melt temperature (K)

: the mass of the melt (t)

: the density of the melt (kg/m3)

: the gravity acceleration (m/s2)

: the depth of the gas inlet (m)

: the gas pressure at inlet (bar).

For p=7 000 kg/m3 and g=9.81 m/s2, is given by

E=

0.014VT

log 11+

H

W 1.46P

(36)

Mori and Sano44~ proposed the following equation

for time of mixing for the purpose of scale-up

z=FL

W/p)2/3

1/3

(37)

where, F (dimensionless) : a proportionality factor

hat depends on the geometry of the melt

ontainer.

In model experiments of solute homogenization in

aqueous solutions, Haida and Brimacombe45> used

torpedo- and ladle-shaped vessels, and found that F

increases with an increase in the diameter/height ratio,

D/H, for the cylindrical ladle and length/height ratio,

L/H, for the torpedo ladle. The following relation

Review

may be obtained from their results

D 0.85 L 0.85

F=150(H)150(H)

(37-a)

Haida and Brimacombe45~ also found that their re-

sults on mixing time in aqueous solutions as repre-

sented by Eq. (37), agree well with those of Nakanishi

et a1.43~or stirred liquid steel mentioned earlier. An-

other formulation of the mixing time given below is

that derived by Sano and Mori46~ rom a mathematical

model that describes circulating flow in a molten steel

bath with inert gas injection.

z=100(D2/H)20.337

- - -

(38)

This derived formula is also found to agree well with

the measured values of mixing time in stirred steel

baths and in water-model experiments.

ikuchi et a1.47~ measured the rate of dephosphori-

zation of low-carbon steels in stirred melts. They

found that the volumetric rate constant, k', for de-

phosphorization increased with an increase in the in-

tensity of stirring. Their results may be approxi-

mated by the relation:

k' (min-1) ^' 0.00 70.4

(39)

which, in combination with Eq. (34), gives k'z ~ 0.07.

That is, as one would expect from the rate of mass

transfer considerations, the volumetric rate constant is

inversely proportional to the mixing time.

2. Hot-metal Refining

ince the late 1970's there has been an intensified

R & D effort, particularly in Japan, to explore the

uses of fluxes in the refining of molten pig iron to low

levels of silicon, sulfur, and phosphorus. The objec-

tives-with an economic and environmental bias-are

two-fold :

i) to increase the recycling of steelmaking slags

(low in phosphorus) to the blast furnace for the re-

covery of iron and manganese; and

ii) to develop slagless steelmaking from hot-metal

low in sulfur, phosphorus, and silicon, thus minimizing

slag disposal.

In his 1981 Howe Memorial Lecture, Fuwa48~present-

ed a detailed account of research and development

in Japan on the ladle treatment of hot metal. This

subject will not be pursued further here with the ex-

ception of citing a few selected interesting examples of

slag-metal reactions pertaining to the desulfurization

and dephosphorization of hot metal by ladle treat-

ment. In practice, the hot metal is first desiliconized

to about 0.1 % Si in the tap-torpedo ladle by the

addition of iron ore or manganese ore followed by

nitrogen stir of the melt. After deslagging, the hot

metal is desulfurized and dephosphorized by injecting

with nitrogen and/or oxygen a mixture of sinter fines,

burnt lime, calcium fluoride, and some calcium chlo-

ride. In some practices, sodium carbonate alone is

injected. At low temperatures and with fluxes having

high capacities for phosphorus and sulfur, the hot


13/21

metal can be refined to the residual levels of 0.005

% P and 0.002 % S. It is well to remember that the

hot metal thus refined contains hardly any silicon,

therefore there is severe limitation to the amount of

scrap steel that can be melted during decarburization

in the converter.

t sufficiently low oxygen potentials, the phospho-

rus dissolves in molten slags and fluxes as a phosphide

ion (P3-). Tabuchi et a1.49) studied the following

phosphide and phosphate reactions in CaO-A1203

and CaO-CaF2 melts.

-P2 + 3 (02-) = (P3~)+ 3..

...............(40)2 4

±p24(Q2-)Q2E- = (POr).

...........(41)

2 2 4

They found that the critical partial pressure of oxygen

for transition from the phosphide to phosphate reac-

tion at 1 550 °C is about 10-18 bar 02 for the calcium

aluminate melt and about 10-16 bar 02 for the lime-

saturated CaO-CaF2 melt. Their experimental data

give for the phosphate capacity, in units of wt%/

bar5/8, at 1 550 °C:

_ (%P205) (2)/ ..................4P(P

P2)12(Po2)54

GP=4.1 x 1022 for the aluminate melt (41 %CaO+

59%A1203) and C =1.4 x 1021for the lime-saturated

CaO-CaF2 melt. To facilitate comparison with other

slag-metal equilibria for the phosphate reaction, Cp

may be converted to another formulation of the phos-

phate capacity

(%P205)

p [%P][%0]2.5

~ )1/2

P2 02

= CP {%P} [%]2

.5 ...............(43)

With the known free energy data for solution of gase-

ous oxygen in liquid iron50)and of gaseous phosphorus

(P2) in liquid iron,51) the CP values of Tabuchi et al.49)

for 1 550 °C give KP=4.1 x 109 for the aluminate melt

(N 40 % CaO) and KP =1.4 x 108for the fluoride melt

(N20 % CaO).

n the system CaO-Fe0r-P205 saturated with CaO

and Ca4P209, the melt contains about 28 % P205 and

15 % FeOt. For this saturated slag, the equilibrium

constant KP in Fq. (28) gives K=2.1 x 106 for 1 550

°C and KP

=0.6 x 106 for 1 600 °C. The values of KP

derived from the results of Tabuchi et al.49) are two

to three orders of magnitude higher than that for the

lime- and phosphate-saturated melt.

n laboratory experiments, Kawai et al.52) mea-

sured the concentrations of oxygen and phosphorus in

iron samples equilibrated with lime/fluoride- and lime/

alumina-based slags at 1 600 °C. Their data give

values of KP increasing from about 105 at 50 % (CaO

+CaF2) to about 106 at 85 % (CaO+CaF2); for the

aluminate melts (with no CaF2) containing 35 to

40 % CaO, the value of KP is in the range 4 x 104 to

1.3 x 105. High values of KP derived from the data

of Tabuchi et a1.49)are inconsistent with other slag-

metal equilibrium data. This disparity may be due

Transactions ISIJ, Vol. 24, 1984 ( 603 )

to the indirect method of estimating the oxygen poten-

tial in their experimental system.

ecause of current interest on fluoride-based slaps

or fluxes for hot metal dephosphorization, numerous

speculations have been made of the possible effect of

CaF2 on the phosphate capacity of slags. An empiri-

cal equation proposed by Healy29) to describe the

phosphorus distribution ratio as a function of slag

composition when used for fluoride slaps gives some

unusual results. As shown by Kawai et a1.,52) or ex-

ample, such an exercise would predict a distribution

ratio (%P)/[%P] in an 80 % CaF2-based slag about

four orders of magnitude greater than that in the

CaF2-free slag, an impossible situation.

s discussed elsewhere,6) and briefly mentioned

earlier, for lime- and magnesia-saturated slags the

KPS in Eq. (29) for a given temperature is a single

function of the concentration of iron oxide in the slag.

On the basis of this reasoning, the data of Kawai

et a1.52) nd Kikuchi et a1.47) or 1 600 °C are plotted

in Fig. 11 where the dotted curve represents the equi-

librium relation for complex steelmaking-type slags

(containing 0 to 5 % CaF2) assessed in a previous

study.6) It should be noted that Kikuchi et a1.47) ave

their results in a graphical form as isoplethes of (%P)/

[%P] in a pseudo-ternary system CaO-CaF2 FeOt

(total being 88 %) of MgO-saturated slags containing

about 4 % Si02 and 1 to 2 % P205. Considering the

diversity in slag compositions, there is a general agree-

ment between the results of vastly different types of

experimental work.

he alumi.nate anions copolymerize with silicate

and phosphate anions to form complex alumino-silico-

phosphate anions. Therefore, in the functional rela-

tion KPSwe may include alumina. Assuming that 0.5

mol A1203 is equivalent to 1 mol Si02, the concen-

tration of A1203 as molar silica equivalence would be

1.18 x %A1203. In the lower diagram in Fig. 11, log

(1+ %Si02+ 1.18 x %A1203) (%P205)/[%P] is plot-

ted against the total iron oxide content of the slag. It

follows from the foregoing analysis that CaF2 has no

particular effect on the phosphate capacity of the slag

where CaF2 behaves much like CaO. The major role

of CaF2 in slaps for hot metal dephosphorization is to

lower the liquidus temperature and lower the viscos-

ity of the slag for use at hot metal temperatures. The

highest distribution ratio (%P)/[%P] attainable is

with lime-saturated slags containing 15 to 25 % FeO,

and low concentrations of Si02 and A1203.

s would be anticipated from the thermodynamic

properties of phosphate melts, the ratio (%P)/[%P]

obtained with soda-based slaps is higher than that with

lime-based slaps. For example, Marukawa et a1.53)

have found that in the dephosphorization of liquid

iron, essentially saturated with graphite, the ratio

(%P205)/[%P] is about 1 000 at 1 300 °C for slag

basicity of Na2O/Si02=3; a similar distribution ratio

is obtained at a lower basicity of 1.5 when the tem-

perature is decreased to 1 200 °C.54) It should be

noted, however, that carbon increases the activity

coefficient of phosphorus dissolved in iron; for hot met-

al compositions the activity coefficient of phosphorus

Review


14/21

(604)

Transactions

ISIJ,

Vol. 24, 1984

is about 5,55) which in part accounts for the attain-

ment of high values of the ratio (%P)/[%P] in the

dephosphorization of hot metal.

n a recent publication, Takeuchi et a1.56jpresented

an interesting account of transitory reactions that oc-

cur in hot metal refining by injection of flux with oxy-

gen, based on their experiments with 100 t hot metal.

A mixture of lime, iron ore, and fluorspar was injected

with a 4: 1 02: N2 mixture at a depth of 1.5 m below

the melt surface in a transfer ladle. The oxygen ac-

tivity in various locations in the metal bath was mea-

sured with an EMF-oxygen sensor. The oxygen ac-

tivity in the bath at locations away from the gas

stream and well below the slag layer were close to the

value expected for the carbon-oxygen equilibrium

with C0; the oxygen activity was about 100 times

greater near the slag layer and 1 000 times greater

near the lance tip. The variation in oxygen activity

with location in the melt was in general accord with

the pattern of liquid flow. The metal was dephos-

phorized at 1 300 °C from about 0.1 to 0.01 % P with

carbon remaining at about 4.4 %. From the com-

positions of metal and slag samples and measured oxy-

gen activities near the slag layer, they estimated the

phosphate capacity of the slag. For CaF2-fluxed

slaps with basicity VN 3, they obtained values in the

range 1025 o 1026 or Cp at 1 300 °C; this corresponds

to Kp 1010which is close to the value estimated by

extrapolating the data for steelmaking slags from 1 600

to 1 300 °C, with due adjustment of Kpo for the ac-

tivity coefficient fp N 5 in graphite saturated iron.

ompared to dephosphorization, it is much easier

to desulfurize hot metal by various means, e.g., injec-

tion of magnesium, calcium carbide, burnt lime, and

fluoride- or soda-based slags used for dephosphori-

zation. In a recent publication, Turkdogan and

Martonik57~ presented some equilibrium data for the

gas-slag-metal reaction:

aO(s)+S+C = CaS(s)+CO............(44)

It will suffice for the present to state that at 1 bar

pressure of CO and for graphite-saturated melts coex-

istent with solid CaO and CaS, the equilibrium con-

centration of sulfur in liquid iron is about 4 ppm at

1 600 °C and increases to about 8 ppm at 1 400 °C.

With injection of burnt lime, the hot metal can, in

practice, be desulfurized to about 20 ppm S. Emi

and Iida,58~however, were able to desulfurize hot met-

al to 12 ppm S at about 1 400 °C by injection of pow-

dered limestone into hot metal in the transfer ladle.

Improved efficiency of desulfurization close to the

equilibrium levels of sulfur is attributed to the higher

reactivity of finely dispersed CaO that is formed upon

in situ dissociation of injected limestone powder. An-

other advantage is that the CO2 generated reacts

exothermically with silicon in the hot metal, thus

compensates for the heat losses.

3. Liquid SteelRefining

n the ladle refining of steel we are concerned

primarily with deoxidation, desulfurization and de-

gassing of steel. Recently, the author59~ presented a

critical evaluation of current knowledge on ladle de-

oxidation, desulfurization, and inclusions in steel, both

fundamental and practical aspects. The present dis-

cussion of this subject, which does not include degas-

sing, is an addendum to the author's previous paper.

alcium aluminate-based slags play a key role in

the ladle deoxidation and desulfurization of steel. To

ensure the attainment of isotropic mechanical prop-

erties in the steel, the sulfide inclusions that form

during solidification of the steel should be in the form

of evenly dispersed globules of sizes in the micron

range. The presence of alumina inclusions in the

steel is undesirable in continuous casting and in many

applications of the steel; the calcium aluminate, as

finely dispersed particles, is the preferred form of oxide

inclusions in killed steels. These objectives can be

fulfilled for killed steels containing less than a total of

25-ppm sulfur+oxygen by achieving a fine dispersion

of minute particles of molten calcium aluminate in

Fig. 11, Variation of the equilibrium relation Kp5 with the

iron oxide content of calcium fluoride and calcium

aluminate slags is compared with saturated-steel-

making type slags at 1 600 °C,

Review


15/21

liquid steel during the ladle treatment.

n the ladle desulfurization of steel with aluminum

and prefused calcium aluminate slag, the following re-

action prevails.

CaO)+ 2 AI+S = (CaS)+ 1(A1203) ......(45)

3

The equilibrium constant for this reaction determined

experimentally by Schurmann et al.60~ or 1 600 °C is

about three-fold of the value computed59~ from ther-

mochemical data and reliable equilibrium data for

various gas-slag-metal reactions. To resolve this dis-

crepancy, Ozturk and Turkdogan61 have recently

determined, under closely controlled experimental

conditions, the equilibrium distribution of sulfur be-

tween calcium aluminate melts and liquid iron con-

taining aluminum. Also, the measurements were

made of the solubility of CaS at 1 550 and 1 600 °C

over the composition range of the liquid phase from

the aluminate to the lime saturation. The results of

these equilibrium measurements were found to be in

close agreement with those computed previously59~

from other equilibrium data.

s is seen from the equilibrium relation in Fig.

12(A) for steel containing 0.04 % Al in solution, the

sulfur-distribution ratio increases appreciably with

increasing ratio CaO/ A1203. Obviously, the lime-

saturated aluminate slag should be used for effective

desulfurization of steel. The effect of aluminum con-

centration in steel on the sulfur-distribution ratio is

shown in Fig. 12(B) for the lime-saturated aluminate

melts in which the solubility of CaS increases from


(605)

1.8 % S at 1 500 °C to 2.2 % S at 1 650 °C. These

equilibrium data indicate that the steel containing

0.03 to 0.05 % Al can be desulfurized to the levels of

a few ppm S by interacting the argon-stirred liquid

steel with the lime-saturated aluminate slag. In fact,

in a well-controlled ladle refining practice,* it has

been possible to desulfurize steel on industrial scale to

levels below 10 ppm S.

f the ladle treatment of steel with calcium alumi-

nate is intended to achieve cleaner steel, i.e., low in

total oxygen as aluminate inclusions, without desul-

furization, the ratio CaO/A1203 in the synthetic slag

and the residual aluminum in the metal should be

reduced as indicated by the equilibrium data. For ex-

ample, for the steel to contain 0.02 % S and 0.02 % Al

after the ladle treatment, the prefused calcium alumi-

nate to be used should be saturated with sulfur, e.g.,

-'2 % S, and have a mass ratio CaO/A1203 of about

1: 1. This is in the middle of the liquid composition

range; therefore, the slag has capacity to absorb alu-

mina inclusions from the steel when mixed with the

metal with argon stirring. For the particular case

considered as an example, the equilibrium concentra-

tion of oxygen in solution in the steel would be about

2 to 3 ppm.

he aluminum-killed steel may be desulfurized also

by injecting a mixture of burnt lime and fluorspar.

The extent of desulfurization that can be achieved

depends on the efficiency of fluxing of the deoxidation

product alumina with the injected lime. Kor and

Richardson62> determined the sulfide capacity, GS, of

CaO-CaF2 A1203 melts at 1 500 °C. Their data are

Fig

(a)

(b)

. 12. Equilibrium sulfur

distribution between liquid calcium aluminate and low

alloy steel.

No slag carry-over during furnace tapp

Basic refractory lining for the ladle

ing

(c)

(d)

High argon flow using porous plug

Ladle cover to avoid air oxidation of th

After Ozturk and Turkdogan.61

e melt

Review


16/21

( 606)


presented in a different form in Fig. 13 as a plot of

log KS against the concentration of CaF2 for indicated

mass ratios of %A1203/%CaO. It is seen that CaF2

has only a small effect on the sulfide capacity of the

aluminate slag. However, CaF2 fluxes both CaO and

CaS; for the invariant equilibrium at 1 500 °C, the

liquid phase contains about 60 % CaF2, 20 % CaS,

and 20 % CaO. Therefore, an addition of 5 to 10 %

CaF2 to the injected burnt lime would curtail the coat-

ing of lime particles with a layer of CaS, hence would

improve the efficiency of lime usage in ladle desul-

furization.

In the silicon-deoxidized steel, the desulfurization

is governed by the reaction :

CaO)+ Si+S = (CaS)+ 1(Si02)

2

...........................(46)

The sulfide capacity, and the deoxidation capability

with silicon, of calcium silicate melts is low even at

lime saturation; therefore, desulfurization by reaction

(46) is possible only when the activity of silica is low-

ered significantly by using a suitable flux such as pre-

fused calcium aluminate. The relation shown in Fig.

14 for sulfur and silicon concentrations is for steels

equilibrated with calcium aluminate+5%Si02 melts

saturated with CaO and CaS. In this slag, the ac-

tivity of silica is in the range 10_5 to 10_4 and the

activity of alumina about 5 X 1013, both with respect

to the pure solid oxides, for which the following are

the equilibrium concentrations of silicon and alumi-

num in the steel at 1 600 °C: 0.01 % Al for 0.07 % Si

and 0.03 % Al for 0.3 % Si. However, at these low

oxide activities in the slag considered, the kinetics of

Si-Al redox reaction would be sluggish, consequently

at low solute concentrations, e.g.,


17/21

Transactions ISII, Vol. 24, 1984

(607)

blast furnaces have revealed that while the sulfur con-

tent of the slag increases with the descent of the burden

in the bosh, the sulfur and silicon contents of the metal

droplets reach maxima and then decrease as they pass

through the slag layer.19'63,64) This phenomenon of

vapor-phase mass transfer has been substantiated by

the experimental work of Tsuchiya et a1.65) nd Tukr-

dogan et a1.14) nder conditions partly simulating the

bosh region of the blast furnace.

n the presence of carbon and depending on tem-

perature and activity of silica in the slag or coke, the

SiO vapor is generated by one of the following reac-

tions.

Si02 (in coke ash or slag)+C --~SiO+CO......(47)

iC+CO -> SiO+2C .................................(48)

At 1 bar pressure of CO, as1o2=1and above about

1 500 °C, Si02 is converted to SiC and SiO is gen-

erated by reaction (48); for a5102=0.1, reaction (48)

applies only at temperatures above about 1 610 °C.

In the experiments cited,14) the SiO generated by pas-

sage of CO through a coke bed at temperatures of

1 600 to 1900 °C, were found to be close to the equi-

librium values for reaction (48). Volatile SiS is gen-

erated by reactions of the type with the coke ash :

aS (in coke ash)+SiO -->SiS+CaO ............(49)

eS (in coke ash)+SiO+C -p SiS+CO+Fe

.......................(50)

The reaction of these vapor species with metal and

slag may be represented by the equations:

iO+C = Si+CO

etal SiSSi

+S ..............................(51)

SiO+0.5 (02-)+0.5C = ( )Si-O-)+0.5C0

lag SiS+ 1.5 (02-)+0.5C = ( Si-O-)+(S2-)

0.500

.......................(52)

Experiments have been made to simulate the events

in the upper part of the blast-furnace hearth where

metal droplets pass through the slag layer.14) Within

a short residence time (1 to 2 s) of metal droplets in

the slag, appreciable changes occurred in the composi-

tion of the metal droplets : manganese oxide in the

slag oxidized silicon in metal droplets from 1.0 % Si

to about 0.2 % Si and the metal was desulfurized

from 0.1 % S to about 0.02 % S.

oke samples taken from the tuyere zone of the

blast furnace usually contain 12 to 16 % ash which

has a basicity (CaO+MgO)/Si02 of about 0.6; for a

typical blast furnace slag, the basicity is about 1.5.

Therefore, the partial pressures of SiO and sulfur-

bearing species in the gas near the coke surface are

expected to differ from those interacting with the slag.

For the purpose of comparison, calculations are made

for assumed local equilibrium at 1 500 °C for the sys-

tems gas-slag and gas-coke ash; the results are given

in Table 1. The amount of ash in the tuyere coke

and its composition suggests that, although most of

the silica and sulfur in the coke ash are removed from

the furnace by the slag and liquid iron, there is sig-

nificant recycle of silicon and sulfur in the bosh region

by the vaporization and condensation of these species

on the coke particles.

he alkalies constitute another dominant vapor spe-

cies in the bosh and stack regions of the blast furnace.

Many studies have been made of the alkali recycle in

the blast furnace via the process of vaporization and

condensation; the subject is well covered in the ref-

erences cited.66,67) General concensus evolved from

elementary thermodynamic calculations is that the

alkali silicates in the ore and coke ash decompose at

elevated temperatures in the lower part of the bosh

and the combustion zone. In part the alkali vapors

carried away with the ascending gas react with the

slag and can therefore be removed from the furnace;

part are converted to alkali cyanides and carbonates

and deposited on the burden and the refractory lining

of the furnace stack. Some new thoughts are pre-

sented in the following discussion on the sequence of

reactions and vapor species that may govern the alkali

recycle in the blast furnace.

s an example, let us calculate the equilibrium va-

por pressure of potassium for the reduction reaction :

K20)+C = 2K (g)+CO (g) ............(53)

The computed equilibrium values in Table 2 are for

a blast furnace slag containing 0.5 % K20 and having

a basicity of B'= 1.5 : K is the equilibrium constant

for reaction (53) from the thermochemical data and

aK2os the activity of K20 in the slag, relative to the

hypothetical solid K20, from the experimental data of

Steiler.6S)

If the potassium input with the blast furnace burden

is 4 kg. Kf t-HM and 80 % of it is removed by the

slag, about 0.8 kg. Kit-HM will be in the vapor phase

Table 1. Calculated equilibrium vapor pressures in

ar for the systems gas-slag and gas-coke

sh at 1 500 °C and 1 bar pressure of CO.

Table 2. Equilibrium

eaction (53)

nd B' -1.5.

vapor

with

pressure of potassium for

slag having 0.5 % K20

Review


18/21


carried away by the furnace gas. For a typical total

gas volume of 2 300 nm3/t-HM, the partial pressure of

the total potassium vapor species would be 7.2 X 10.4

bar (for a total gas pressure of 3.6 bar in the bosh

region). Comparison of this total potassium pressure

in the gas with those in Table 2 for the gasslag equi-

librium indicates that the slag particles will pick up

potassium from the gas only at temperatures below

1 200 °C. The ash in the coke samples taken from

the tuyere zone contains 2 to 4 % (Na20+K20)

which is several fold of that in the original coke ash.

Because of the low basicity of coke ash, the amount

of alkali therein is expected to be 5 to 10 times greater

than that found in the slag. In the upper part of the

bosh, the slag and coke will pick up alkalies from the

gas. As the temperature of the slag and coke increases

during descent in the bosh, they will emit alkali vapors

to the gas phase. However, the partial pressure of

potassium vapor in the gas would probably be below

that given in Table 2 for the gas-slag equilibrium.

Nevertheless, the numerical examples cited suggest

that much of the alkalies accumulated in the bosh and

hearth zone of the blast furnace, via recycling, are in

the gas phase.

ccording to a mass spectrometric study of vapori-

zation of potassium cyanide by Simmons et al.,69) the

equilibrium pressures of vapor species are in the order

p2 KCN) PKCN P3KCN) For atmospheric pressure of

nitrogen and at graphite saturation, the following are

the equilibrium ratios of the vapor species : at the

melting point of KCN (635 °C), (KCN)2: KCN: K

=2: 1: 0.76 and at the boiling point (1 132 °C),

(KCN)2: KCN: K=4.6: 1: 0.19.* It is all too clear

that (KCN)2 is the dominant alkali-bearing species in

the blast furnace. The sodium in the burden under-

goes cyclic reactions similarly to potassium; however,

the concentration of sodium-bearing species are about

one-tenth of the potassium-bearing species.

n addition to the well-known problems associated

with alkali accumulation in the blast furnace stack,

the alkali cyanides are also responsible for the genera-

tion of ammonia and hydrogen cyanide in the upper

part of the stack by the reactions :

2KCN+3H20 = K2C03+2NH3+C .....

NH3+C0 = HCN+H20 ....................

.(54)

(55)

These and other organic volatile matters thus gen-

erated in the stack end up in the wash water of the

blast furnace off-gas. Recent experimental work of

Turkdogan and Josephic,70) under conditions partly

simulating the blast furnace stack, substantiated the

validity of this mechanism of generation of ammonia

and hydrogen cyanide in the stack.

2. Iron Oxide Fumes

volution of iron oxide fumes is a common occur-

rence in all steelmaking operations, e.g., in melting, in

refining with oxygen blowing, and during furnace tap-

ping. The mechanism of this interesting phenom-

enon was resolved in an earlier study by Turkdogan

et al.,71) based on the concept of diffusion-limited en-

hanced vaporization of metals in an oxidizing envir-

onment. In this counterflux-transport process, at

some short distance from the surface of the metal, the

oxygen and metal vapor react to form a metal oxide

mist (fume). The formation of an oxide in the gas

phase provides a sink for the metal vapor and oxygen

resulting in the counterflux of these two gaseous spe-

cies. An increase in the oxygen partial pressure, and

or an increase in the gas-film mass-transfer coefficient

for oxygen, decreases the thickness of the sub-bound-

ary layer through which the metal vapor is diffusing,

hence increases the rate of fuming. Also, the higher

the vapor pressure, or higher the temperature, the

faster would be the rate of fuming. For given tem-

perature and conditions of gas flow over the metal sur-

face, the rate of vaporization is represented in a gen-

eral form by the relation :

Rate of fuming = P(1-0) exp (-EJRT)

0(1- D)p~,,

}

where,

(56)

When

metal

diffusi

becom

-f

the fur

fumin

zation

expect p

tally.

During furnace tapping, the metal stream is ex-

posed to air, consequently the iron oxide fumes are

emitted by the mechanism described. The fume emis-

sion is particularly intensive in the blast furnace metal

runner, because of the high vapor pressure of man-

ganese at the concentrations present in the hot metal.

Furthermore, the metal droplets ejected by gas evolu-

tion from the hot metal in the runner increase the

gas-metal interfacial area, consequently intensify the

fume emission. It follows from the basic concept of

the mechanism of fuming that the extent of fume emis-

sion can be suppressed by either eliminating gaseous

oxygen from the exposed surface of the metal or by

covering the metal surface with a layer of slag. De-

pletion of oxygen near the metal surface can be

achieved by shrouding the metal runner with a gas

F: is the energy of vaporization

, ~ : constants for given conditions of

teady-state vaporization

(1-0) : the fraction of metal surface exposed

o the oxidizing gas stream.

the flux of oxygen toward the surface of the

is greater than the equivalent countercurrent

ve flux of the metal vapor, the metal surface

es oxidized, i.e., (1-D) , and consequently

ing ceases. Just before this cutoff, the rate of

g will be close to the maximum rate of vapori-

that is attainable is vacuo. These theoretical

ations have been well substantiated ex erimen-

* It should be noted that in the free energy tables in p

. 12 of Ref. 5), the value given for the vaporization of liquid KCN is incorrect;

t should have been, according to the data of Simmons et al.ss~

KCN (1) _ (KCN)2 (g)

G° = 52180-36.64Y(cal)

Review


19/21

flame. Extensive plant trials done recently in U.S.

Steel have substantiated the viability of this method

of fume suppression during furnace tapping.

3. White Fumes rom Blast Furnace Slag Runner

hite fumes emitted from the blast furnace slag

runner is another example of vaporization-induced oc-

currence. Samples of white fumes collected over the

slag runner have been characterized recently by Simon

and Kelly,72~ n U.S. Steel Research Laboratory, using

the scanning electron microscope. The X-ray scan

spectra revealed that the potassium and sulfur were

the most dominant elements in the fume samples.

Such a finding is consistent with the expected reaction

of slag with air during tapping.

he blast furnace slag contains 1.5 to 2.0 % S, and

as discussed earlier, the potassium vapor pressure of

the slag is relatively high. Therefore, upon exposer

to air, the potassium and sulfur will be evolved from

the slag as potassium sulfate by the reaction:

(K+)+(S2-)+202 = K2SO4 (white fume)

..........................(57)

In addition, there is SO2 evolution, readily evidenced

by the characteristic odor in the casthouse,

S2-)+ 3 02 ` (02-)+S02 ............(58) 2

In the oxidizing atmosphere over the slag, the sulfur

dioxide is partly oxidized to sulfur trioxide. In the

cooler ambient atmosphere away from the slag sur-

face, 503 and K2SO4 will react with moisture in the

air and form a sulfate complex xK2SO4.yH2SO4 which

appears as white fume. The X-ray scan spectra indi-

cated an atom concentration ratio K/S of about one

in the collected fume samples, which apparently has

the composition of K2SO4. H2SO4. The higher the

slag temperature, the greater the extent of fuming

caused by the above reactions.

he emission of white fumes can be suppressed by

spreading a thin layer of sand over the slag surface in

the runner. This simple remedy works for two rea-

sons : added sand, or other inert powder, chills the salg

surface and also reduces the surface area of slag ex-

posed to air.

amamoto and Harashima73~ studied changes in

chemical composition of slag during the refining of

hot metal with sodium carbonate and oxygen-top

blowing. They deduced from the composition of slag

and fume collected that there was preferential vapori-

zation of sulfur and sodium from the slag. This ob-

servation can be explained also in terms of reaction

(57) written for Na+ in the slag and Na2SO4 in the

fume.

4. Vaporization in Coal Gas cation

ertain vapor species that are generated in some

coal gasification processes may cause side reactions

which could affect the industrial application of the

gasification process. Because such processes are still

in the early stages of development, technical informa-

Transactions ISIJ, Vol. 24, 1984 ( 609)

tion available is meagre, therefore we can at present

only speculate on the possibility of vaporization-in-

duced side reactions and consequences thereof.

n the coal gasification processes that are now being

developed, e.g., CGS, COIN, KR, the pulverized or

granular coal is partially combusted with oxygen and

steam in a bath of liquid iron to generate a reducing

gas for use in the ore reduction and smelting processes.

In his doctorate dissertation, Fukagawa74~presented a

comprehensive review of much of the work done to

date on coal gasification using liquid iron as the re-

action medium. Technical developments made to

date on direct usage of coal in the steel industry are

presented in a series of papers published recently in

Ironmakingand Steelmaking.75~

ench scale experiments and pilot plant tests have

shown that about 15 to 30 % of the sulfur in coal is

absorbed by the molten iron and slag in the gasifier

and the remainder transferred to the gas and iron dust

(fume) emitted. The gas--metal interaction in the

combustion zone is conducive to enhanced vaporiza-

tion of iron by the mechanism already described.

Away from the combustion zone, the iron oxide fume

is reduced to metallic iron dust which interacts with

sulfur-bearing vapor species in the reducing gas. As

the temperature of the dust-laden gas decreases, more

sulfur is absorbed by the iron dust. Upon cooling

and dedusting, the reducing gas generated contains

less than 300 vol-ppm H2S. The iron oxide fume

emitted during gasification of coal is an efficient gas

desulfurizer contained in the system only when the

reducing gas is cooled and dedusted. The dust re-

moved from the cooled reducing gas contains up to

about 8 % S, a material not suitable for recycle.

he sulfur and dust-related problems with these

coal gasifiers have to be surmounted for direct usage

of hot reducing gas in the shaft or fluidized bed reduc-

tion of iron ore. Noting that the hot reducing gas

enters the ore reducer at about 900 °C, the gas has to

be dedusted at a relatively high temperature, for

which a new technology is needed. In a reducing

gas containing 25 vol% H2, the entrained iron dust

will not desulfurize the gas below 800 vol-ppm H2S at

900 °C. Such a high sulfur-bearing gas is not suitable

for the reduction of iron ore. Hot dedusting and de-

sulfurization may be achieved by passing the gas, for

example, through a bed of calcined dolomite at about

950 °C. In such a process, the sulfided doloma has

to be removed periodically from the reactor and be

treated for disposal. Of course, other methods of hot

dedusting and desulfurization can be visualized; the

subject will not be pursued further here.

nother problem to be resolved is that involving

alkalies in the coal. Since coal contains some nitro-

gen, most of the alkalies will be contained in the gas

generated as alkali cyanide vapors, which will even-

tually condense in the ore reducer. Obviously, mea-

ures have to be taken to prevent the accumulation of

alkali carbonates and cyanides in the upper part of

he shaft reducer.

e see from the examples cited, there are many

amifications to the role of vapor species in pyrometal-

Review


20/21


lurgical or other high temperature processes. The

vapor species generated may cause pollution prob-

lems, may impede the operation of the process or may

assist the process reactions.

V. Concluding Remarks

cope of the physical chemistry of ironmaking and

steelmaking is, of course, much greater than what has

been presented in this discourse. Therefore, the con-

cluding remarks are confined to the subjects discussed

in this paper. The state of reactions in pyrometal-

lurgical processes can be evaluated with confidence

from accumulated knowledge of the physicochemical

and thermodynamic properties of materials and of

the process engineering concepts pertinent to high

temperature technology. The analysis of plant data

indicates departures from equilibrium for various

gas-slag-metal reactions in the bosh and hearth

zones of the blast furnace. The least departure from

equilibrium is for the distribution of silicon between

metal and slag. Departure from equilibrium for the

manganese-silicon reaction is greater for slags of low

basicity. Despite departures from equilibrium, the

silicon and sulfur contents of the hot metal change

in a systematic manner for a given blast furnace, i.e.,

high silicon/low sulfur and low silicon/high sulfur.

n oxygen steelmaking processes, most reactions ap-

proach slag-metal equilibrium at the time of furnace

tapping. However, some differences are seen in the

behavior of phosphorus reaction in the oxygen-top and

-bottom blowing processes

. The salg/metal phospho-

rus distribution is closer to equilibrium in Q; BOP than

in BOF steelmaking, particularly for low carbon con-

tents at tap.

greater use is being made of the gas-slag-metal

equilibrium data in recent developments of ladle re-

fining of hot metal and steel. At temperatures of

about 1 300 °C and with fluxed-basic slags of low

melting temperatures, it is now possible to dephospho-

rize hot metal to the levels of 0.005 % P at carbon

contents in the range 4.0 to 4.5 %. The aluminum-

killed steel can be desulfurized to levels below 10 ppm

S by interacting liquid steel with a lime-saturated

aluminate slag using argon stirring for slag-metal

mixing.

Vapor species are now known to play various roles

in pyrometallurgical processes. The alkali cyanide

vapors are responsible for alkali recycle and accumula-

tion in the blast furnace and also for the generation of

ammonia and hydrogen cyanide in the upper part of

the stack. The silicon and sulfur-bearing vapor spe-

cies are responsible for the transfer of silicon and sul-

fur from coke to metal and slag in the bosh region of

the blast furnace. Fumes emitted upon exposure of

liquid steel, hot metal or blast furnace slag to air are

consequences of some vaporization phenomenon.

REFERENCES

1) H. Schenck: Physical Chemistry of Steelmaking, The Brit.

ron Steel Research Assoc., London (1945).

2) Physical Chemistry of Process Metallurgy, Disc. Farad.

oc., (1948), No. 4.

Review

3) 0. Kubaschewski and C. B. Alcock: Metallurgical Ther-

ochemistry, 5th ed., Pergamon Press, Oxford, (1979).

4) F. D. Richardson : Physical Chemistry of Melts in Metal-

urgy, I & II, Academic Press, London, (1974).

5) E. T. Turkdogan: Physical Chemistry of High Tempera-

ure Technology, Academic Press, New York, (1980).

6) E. T. Turkdogan : Physicochemical Properties of Molten

lags and Glasses, The Metals Soc., London, (1983).

7) P. Herasymenko: Trans. Faraday Soc., 34 (1938), 1245.

8) P. Herasymenko and G. E. Speight: JISI, 166 (1950), 169

289.

9) H. Flood and K. Grojtheim: JISI, 171 (1952), 64.

10) T. Forland and K. Grojtheim: Met. Trans., B8 (1977),

45.

11) J. D. Shim and S. Ban-ya : Tetsu-to-Hagane, 67 (1981),

745.

12) S. Ban-ya, T. Watanabe and R. Nagabayashi: Japan/USA

eminar on Advances in the Science of Iron- and Steelmak-

ng, Kyoto, May 1983, (1983), 27.

13) J. Lumsden: Physical Chemistry of Process Metallurgy, I,

. R. St. Pierre, ed., Interscience, New York, (1961), 165.

14) E. T. Turkdogan, G. J.W. Kor and R. J. Fruehan : Iron-

akingSteelmaking, 7 (1980), 268.

15) E. W. Filer and L. S. Darken: Trans. AIME, 194 (1952),

53.

16) G. G. Hatch and J. Chipman : Trans. AIME, 185 (1949),

74.

17) J. Taylor and J. J. Stobo: JISI, 178 (1954), 360.

18) E. T. Turkdogan: Met. Trans., B9 (1978), 163.

19) K. Okabe et al.: Kawasaki Steel Tech. Res. Lab. Rept., (May,

974), 1-43; (Nov., 1974), 1-11.

20) E. T. Turkdogan: JISI, 182 (1956), 74.

21) F. D. Delve, H. W. Meyer and H. N. Lander: Phys. Chem.

rocess Metallurgy, G. R. St. Pierre, ed., Interscience, New

ork, (1961), 1111.

22) W. Hess, S. Mayer and P. Schulz: Stahl u. Eisen, 93 (1973),

07.

23) T. Fuwa and J. Chipman: Trans. Metal. Soc. AIME, 218

1960), 887.

24) S. Ban-ya and S. Matoba : Studies in Metallurgy for

achio Matoba, Tohoku University, Sendai, (1969), 50.

25) W. A. Fischer and H. vom Ende: Stahl u. Eisen, 72 (1952),

398.

26) T. B. Winkler and J. Chipman: Trans. AIME, 167 (1946),

11.

27) K. Balajiva, A. G. Quarrell and P. Vajragupta: JISI, 153

1946), 115; 155 (1947), 563.

28) H. Knuppel and F. Oeters : Stahl u. Eisen, 81 (1961), 1437.

29) G. W. Healy: JISI, 208 (1970), 664.

30) J. Pearson and E. T. Turkdogan: JISI, 176 (1954), 19.

31) H. Suito, R. Inoue and M. Takada: Trans. ISIJ, 21

1981), 250.

32) E. T. Turkdogan: Iron Coal Trades Rev., (Dec. 16, 1955),

471.

33) W. Loscher: Arch. Eisenhuttenw, 36 (1965), 237.

34) K. L. Fetters and J. Chipman : Trans. AIME, 145 (1941),

5.

35) G.J.W. Kor: Metal. Trans. B8 (1977), 107.

36) C. R. Taylor and J. Chipman: Trans. AIME, 154 (1943),

28.

37) Proc. Scaninject I, June 1977, MEFOS, Lulea, (1977).

38) Proc. Scaninject II, June 1980, MEFOS, Lulea, (1980).

39) Proc. Scaninject III, June 1983, MEFOS, Lulea, (1983).

40) Secondary Steelmaking, MS Book No. 190, The Metals Soc.,

ondon, (1978).

41) J. Szekely: Fluid Flow Phenomena in Metals Processing,

cademic Press, New York, (1979).

42) H. Sandberg, T. Engh, J. Andersson and R. Olsson:


21/21

Transactions ISIJ, Vol. 24, 1984 (611)

Jernkont Ann., 155 (1971), 201.

43) K. Nakanishi, T. Fujii and J. Szekely: Ironmaking Steelmak-

ng, 2 (1975), 193.

44) K. Mori and M. Sano : Tetsu-to-Hagane, 67 (1981), 672.

45) 0. Haida and J. K. Brimacombe: Proc. Scaninject III,

EFOS, Lulea, (1983), 5: 1.

46) M. Sano and K. Mori: Trans.

Documents

Physico chemial Aspects of Steelmaking Processes By E TT URKDOGAN Synopsis.pdf