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Desulfurization

by

Eugene Pretorius and Helmut Oltmann Process Technology Group – LWB Refractories

1 Introduction In most steel grades sulfur has to be minimized below a certain level. Sulfur control is important throughout the iron and steelmaking process, from careful choice of raw materials, hot metal desulfurization, limited removal in primary steelmaking and the primary desulfurization process in the ladle. The main sulfur removal mechanism is the transfer of sulfur from the metal to the slag phase. This section reviews the fundamental principles of sulfur removal, with specific elaboration on the kinetics of desulfurisation in the ladle and thoughts on sulfur behavior in the EAF. It is generally known that the following parameters are important for good sulfur removal from the metal to the slag. • Liquid and fluid slag with a high dissolved lime content – high sulfur/sulfide capacity (Cs). • Higher temperature, which improves both thermodynamics and kinetics. • Oxygen content of the steel and oxidation state of the slag – lower is better. • Decent slag volume – higher slag volume removes more S, for same distribution ratio. • Mixing (stirring) of the steel – kinetics of mass transfer to slag metal interface. • The extent of slag and metal inter-mixing at the interface (stirring intensity and slag fluidity)

2 Thermodynamic principles of sulfur removal

2.1 Sulfur removal reaction and sulfide capacity The following reaction is generally accepted to describe sulfur transfer from the metal to the slag:

]O[)S()O(]S[ 22 +⇔+ −− (1)

Where [X] indicates X dissolved in the metal (infinite dilution standard state) and (X) indicates species dissolved in the slag. The equilibrium constant for this reaction cannot be evaluated directly, however, due to difficulty in describing the activity of the ionic species in the slag (O2-

and S2-) and has to be determined experimentally. Experimental determination is usually by gas-slag equilibration techniques based on the following reaction:

22122

221 OSOS +=+ −− (2)

Slag samples of known composition are equilibrated with a gas of controlled pS2 (partial pressure of sulfur) and pO2 and the equilibrium sulfur content of the slag determined by analysis. The sulfide capacity (CS) can then be calculated according to equation (3), which is based on the expression for the equilibrium constant of reaction (2).

−

−

γ==

2

2

21

2

21

2

S

O)2(

S

OS

aK)S(%

p

pC (3)

LWB Refractories 2 Process Technology Group

The term sulfide capacity (CS) is used to describe a completely liquid slag's potential ability to remove sulfur and is a function of slag composition and temperature. Sulfide capacities for various slags have been experimentally determined and examples are shown in Figures 1 and 2. Sulfide capacities are usually expressed in log units, for example: log CS = -3.2. The less negative the logarithmic unit (larger the actual number), the better the sulfur removing capacity of the slag, i.e., a slag with log CS = -1 has a better sulfur removing capacity than a slag with a log CS = -3. Generally the sulfide capacity increases as the CaO-content increases in these slags.

Figure 1. Iso-sulfide capacity curves for CaO-Al2O3-MgO slags at 1600°C. (Numerical values show - log CS.)

Figure 2. Iso-sulfide capacity curves for CaO-Al2O3-SiO2 slags at 1600°C. (Numerical values show - log CS.)

Industrial slags are invariably more complex than the slag systems of which the sulfide capacity has been determined experimentally and a general expression for Cs as a function of slag composition and temperature would be of great use. The correlation of sulfide capacity with


Optical Basicity (OB) has been shown to have a good correlation over a wide range of slag compositions and is therefore widely used.

2.2 Optical basicity

Table 1 shows optical basicity (Λ) values for the most common slag components.

Table 1. Values of optical basicity (Λ) of slag components.

Oxide Optical Basicity (Λ) Na2O 1.15 CaO 1.0 MgO 0.78 CaF2 0.67 TiO2 0.61 Al2O3 0.61 MnO 0.59 Cr2O3 0.55 FeO 0.51

Fe2O3 0.48 SiO2 0.48

The average optical basicity (Λ) for a slag of any composition can be calculated by means of the expression:

Λ = XAoxΛAOx + XBOyΛBOy + ... (4) where,

componentsallofmoleculeoxideinoxygensofnumber*componentoffractionmole

moleculeoxideinatomsoxygenofnumber*componentoffractionmoleXΣ

=

(5)

Thus, in a CaO-Al2O3-SiO2 slag for example:

232 SiOOAlCaO

CaOCaO N2N3N

NX

++=

232

3232

SiOOAlCaO

OAlOAl N2N3N

N3X

++=

232

22

SiOOAlCaO

SiOSiO N2N3N

N2X

++=

While the optical basicity is currently the best compositional parameter available that can be used to describe the composition (basicity) of a slag, it does not say anything about the physical properties of a slag. For example, consider the following "slag" composition:


% CaO 62 % MgO 8 % SiO2 30

The optical basicity of this slag can be calculated as Λ = 0.756, which could be considered sufficiently basic to ensure good sulfur removal and minimum refractory wear. However, the calculation of this number is meaningless because the slag listed above is completely solid at steelmaking temperatures, and it will only start to melt at 1790°C and will be fully molten at 1950°C.

The application of the optical basicity concept to steelmaking slags is only useful if the slags are completely melted. From the above illustration it is clear that significantly more information, such as solidus and liquidus phase relations and viscosity data, are required in order to completely evaluate slags that would be suitable in, for example, the ladle furnace.

2.3 An optical basicity calculation example

Calculate the optical basicity of the following slag:

% CaO - 45, % SiO2 - 35, % MgO - 10, % Al2O3 - 10 Solution: Optical basicity (Λ) Molecular weight

CaO 1 56.08 SiO2 0.48 60.09 MgO 0.78 40.31 Al2O3 0.61 101.96

Moles CaO: 45/56.08 = 0.802 Moles SiO2: 35/60.09 = 0.582 Moles MgO: 10/40.31 = 0.248 Moles Al2O3: 10/101.96 = 0.098 Total moles = 1.730 Mole fraction (N) CaO: 0.802/1.73 = 0.464 Mole fraction (N) SiO2: 0.582/1.73 = 0.336 Mole fraction (N) MgO: 0.248/1.73 = 0.143 Mole fraction (N) Al2O3: 0.098/1.73 = 0.057 Total Base value = NCaO + 2NSiO2 + NMgO + 3NAl2O3 = 0.464 + 2(0.336) + 0.143 + 3(0.057) = 1.45


XCaO = NCaO/Total Base 0.464/1.45 = 0.320 XSiO2 = 2NSiO2/Total Base 2(0.336)/1.45 = 0.463 XMgO = NMgO/Total Base 0.143/1.45 = 0.099 XAl2O3 = 3NAl2O3/Total Base 3(0.057)/1.45 = 0.118 Optical Basicity = (XCaO*ΛCaO) + (XSiO2*ΛSiO2) + (XMgO*ΛMgO) + (XAl2O3*ΛAl2O3) = (0.320 * 1) + (0.463 * 0.48) + (0.099 * 0.78) + (0.118 * 0.61) = 0.691

2.4 Optical Basicity correlation with sulfide capacity

The relationship between sulfide capacity and optical basicity (Λ) is given by:

]2.25)*6.43[(T

)]*54640(22690[Clog S −Λ+Λ−

= (6)

Figure 3 shows the correlation between sulfide capacity and calculated optical basicity of slags at 1500°C (2732°F).

Figure 3. Correlation between sulfide capacity and calculated optical basicity of slags at 1500°C.


In an effort to improve on the accuracy of the correlations with optical basicity in specific slag composition regimes, various adjustments or corrections have been proposed. For example to improve accuracy with CaF2-containing slags a modified optical basicity has also been proposed, which ignores the effect of CaF2 on the optical basicity, but includes a correction term in the calculation of log Cs. Alternative correlations not based on optical basicity are also available to relate CS to the slag composition, but the applicability of these relations is probably limited to slag compositions close to the original compositions used to derive these correlations. In summary, the optical basicity correlation is easy to use and a good estimate over a wide range of slag compositions, but alternative sulfide capacity data can also be used in their applicable range of compositions.

2.5 Importance of liquid slag

One of the most important aspects that is frequently overlooked when evaluating and discussing slags, is that most slags consist of two fractions, i.e., a liquid fraction and a solid fraction. Slag fluidity in basic slags is controlled by the liquid and solid fractions of the slag. The higher the solid fraction of the slag - the lower the fluidity of the slag (higher viscosity). A basic slag that is completely liquid has maximum fluidity. Only the liquid portion of the slag is active in desulfurization, a high solid content effectively reducing the slag volume.

The following example illustrates the importance of fluidity and effective slag volume.

Consider the following two slags at 1600°C (2912°F):

Slag O (Figure 4) Slag K (Figure 4)

% SiO2 38 35 % CaO 45 50 % MgO 17 15

Phases present 100% liquid (Optimum slag)

56% liquid, 39% Ca2SiO4, 5% MgO

In many calculations of sulfur removal, the total composition of the slag is sometimes mistakenly considered. It is the lime dissolved in the liquid fraction of the slag that is removing sulfur from the steel and not lime as determined by the chemical analysis of the "total" slag. The undissolved lime in the slag does not remove any sulfur. The liquid fraction of slag K has the identical composition as slag O. The addition of more lime than required (slag O to slag K), only resulted in a decrease in the amount of liquid available for desulfurization (liquid composition stayed constant) and an increase in the viscosity of the slag.


9080705040302010 60

10

20

30

40

90

80

70

60

50

70

50

60

40

30

20

80

90

10

S + L

M + L

Sea area (all liquid)

Land area (all solid)

Swamp area (liquid + solid)

C2S + L + MM2S + L + M

M2S + L

C2S + LCa2SiO4

Mg2SiO4

Ca3SiO5

SiO2

MgOCaO

K

O

1600°C (2912°F) Isothermal Section

Figure 4. Isothermal section of the CaO-MgO-SiO2 system at 1600°C

The errors that can be made by not considering only the liquid portion of a slag is illustrated by the following example:

Table 2. Example of possible errors in the calculation of final metal sulfur content Considering the

total slag chemistry (Slag O)

Considering the total slag

chemistry (Slag K)

Considering only the liquid fraction

of slag K Temperature (°C) 1600 1600 1600 Si (wt%) 0.035 0.035 0.035 WM (tons) 100 100 100 WS (lb.) 2000 2000 (2000*0.56) = 1120Λ 0.698 0.717 0.698 [% O] (ppm) 15 15 15 log CS -3.01 -2.74 -3.01 LS 17.83 32.84 17.83 SF (wt%) 0.0297 0.0263 0.0318 Calculated result Correct Wrong!! Correct


To summarize: the sulfide capacity can be defined as the ability of a liquid slag to absorb sulfur, determined experimentally as the equilibrium sulfur content for certain conditions (temperature, pS2 and pO2). The sulfide capacity is a function of temperature and slag composition and can be calculated according to correlations based on the concepts of optical basicity.

3 Sulfide capacity in slag-metal systems.

3.1 Log Ls The concept of sulfide capacity was explained and experimentally determined according to the gas-slag reaction (2), but we would now like to apply it to the slag-metal reaction (1). While in the gas-slag system the partial pressure of O2 and S2 are important, in the slag-metal system it is the O and S in solution in the metal that determines the equilibrium sulfur content of the slag.

]O[)S()O(]S[ 22 +⇔+ −− (1)

22122

221 OSOS +=+ −− (2)

By subtracting reaction (2) from reaction (1), we obtain reaction (7) for which the equilibrium constant is calculated as in (8).

]O[SO]S[ 22122

1 +=+ (7)

log K(5) = -935 / T + 1.375 (8)

The derivation is not shown in detail here, but involves a substitution of the expression for sulfide capacity (3) into the equilibrium constant for (7) and converting to the logarithmic form, to arrive at equation (9). This equation relates the sulfide capacity and other known thermodynamic properties of the metal, to the one parameter of prime importance for further evaluation – the sulfur distribution (Ls).

OS alogflogCslog375.1T935

]S[%)S(%logLslog −+++

−== (9)

In this equation, fS is the activity coefficient for sulfur in the metal according to the equation for the activity of S in dilute solution: aS = fS * [%S] and similarly for the activity of oxygen: aO = fO * [%O]. The activity coefficient of S in metal fS can be calculated based on the metal composition and interaction parameters. For common carbon steel grades, the activity coefficient is close to unity (log fS ≈ 0) and this assumption thus has little effect on the calculations. The activity and activity coefficient of oxygen are more variable and need to be considered more carefully.


If the sulfide capacity (Cs) for a particular slag is known, then the sulfur distribution ratio (Ls) can be calculated for a particular slag-metal system in which the values of fS and aO are known. For best sulfur removal a large value of Ls is required, which can be obtained by high temperature, large Cs and low oxygen activity aO. The activity coefficient of sulfur fS is fixed by the metal composition and not really a controlled variable. Note that the Ls ratio is not dependant on the relative slag and metal mass or initial sulfur content.

3.2 Log aO The activity of oxygen can be measured or calculated and has an important impact on the magnitude of Ls. Generally for deoxidized metal in the ladle, the higher the concentration of deoxidizer in the metal, the lower the oxygen activity. The oxygen activity is also typically lower for Al-killed, than for Si-killed steel. The lower the activity of the corresponding oxide (in slag or in metal as inclusions), the lower the oxygen activity. According to (9), a lower oxygen activity in the metal will result in higher Ls and thus improved sulfur removal. The oxygen activity for the desulfurisation calculation, can be determined in one of two ways: • Actual measurement of oxygen activity and temperature in the ladle by an oxygen probe

produces an electro-motive-force or emf value (measured as a voltage), relative to a reference cell of known oxygen potential. The emf is usually re-calculated or converted by the measuring unit into an oxygen content in parts per million (ppm) and sometimes carbon content. (For the oxygen content: [%O] * 10000 = ppm[O]).

• Calculation based on slag-metal reaction equilibrium of the primary deoxidizer specie (usually Al or Si). The difficult part of this calculation is to estimate the activity of the oxide-specie in the slag (Al2O3 or SiO2), which can be done by correlation with experimentally determined activities in similar slag systems or slag models. With knowledge of the temperature, activity coefficient (and concentration) of the deoxidizer specie (Al or Si in metal) and activity of the oxide, the oxygen activity aO can be calculated for use in equation (9).

• Note that the oxygen activity of the bulk metal is measured by the probe, which can differ from the conditions at the slag-metal interface, which is calculated by considering slag-metal equilibrium. When the oxygen content is determined, it should be combined with the activity coefficient fO (which can be calculated based on the metal chemistry of a sample) to obtain the activity aO. In most low carbon steel compositions, the effect of Al, Si and C in the metal are most important to determine the value of the activity coefficient of oxygen (fO). The accuracy in measuring/calculating the oxygen activity that actually controls/determines the sulfur distribution ratio is critical – especially at very low oxygen activity or oxygen content in the steel. The calculation of Ls is very sensitive to oxygen activity at low oxygen activity, so that small deviations have a major impact on the resulting Ls as shown in Figure 5. A slight difference between the true oxygen activity determining Ls and the estimated or calculated value, can result in large deviations between the calculated and true S-behavior. For the curves in Figure 5, the activity coefficient of oxygen was assumed to be one fO = 1, so that aO = [%O] = ppm[O]/10000. At very low oxygen activity there is also a possibility that the conditions at the


slag-metal interface differ considerably from the conditions in the bulk of the metal – both due to concentration gradients and oxide activity. The activity of Al2O3 or SiO2 are lower at the interface due to solution into the slag, whereas in the bulk of the metal, the pure oxides with activity of unity could be present (as inclusions).

0

100

200

300

400

500

600

700

800

900

1000

0 5 10 15 20 25 30 35

Oxygen content - ppm O

Ls =

(%S)

/[%S]

logCS=-2 logCS=-1.8 logCS=-1.6

Figure 5: Sensitivity of Ls to oxygen content.

3.3 Effect of oxidized slag.

The effect of the amount of reducible oxides (FeO, MnO and Cr2O3) in ladle slags on the desulfurizing efficiency is shown in Figure 6 and low reducible oxide content is thus preferable for good desulfurization. For this reason the slag is deoxidized by addition of aluminum shred, FeSi or CaC2 when desulfurization is critical. The definition of the %desulfurization in this widely quoted Figure is not quite clear, but it is reasonable to assume that the reducible oxides act as a source of oxygen at the slag metal interface, which will decrease the effective Ls ratio to a lower value than calculated by considering the bulk oxygen content. A simple correction of the Ls ratio for oxidized species in the slag can thus be made in the calculation.


1 4

100

80

60

40

20

2 3(FeO + MnO), %

% D

esu

lfu

r iza

tio

n

Figure 6: Effect of (FeO & MnO) in slag on desulfurization.

3.4 Equilibrium Sulfur content. In addition to the Ls-ratio, the slag and metal weight (Wm and Wslag) and initial sulfur concentration in the metal (%S°) do affect the final equilibrium sulfur content (%Seq) according to:

slagm

meqW*LsW

W*]S[%]S[%

+°

= (10)

Only initial sulfur content in the metal is considered here and if the slag contains significant sulfur at the start of treatment it should be included, but in most cases it has very little influence on the final sulfur content. (The top half of the equation then becomes: {[%S]°*Wm + (%S)°*Wslag}) According to equation (10), to achieve lower final sulfur content in the metal, the slag volume (Wslag or slag to metal ratio) or Ls can be increased, or alternatively the initial sulfur content in the metal decreased. As shown earlier, the value of Ls is affected mainly by the slag composition (sulfide capacity), temperature and oxygen content of the metal.

3.5 Slag design Published phase diagram and liquidus slag models can be used to design slags for specific metallurgical requirements. The choice of fluidizers (SiO2, Al2O3, and/or CaF2) in the design of these slags is very important because the solubility of CaO and MgO is strongly dependent on the type of fluidizer used. As seen in the optical basicity calculation, CaO has the highest optical basicity of the common ladle slag components and will have the most impact on the optical basicity and thus sulfide capacity. In order of decreasing basicity the order is CaO, MgO, Al2O3 and SiO2. It follows that replacing Al2O3 with SiO2 as fluidizer will lower the sulfide capacity,


both because the optical basicity of SiO2 is lower than Al2O3 and because lime solubility decreases. In practice we cannot use any slag composition (or temperature), but are restricted by slag fluidity and compatibility with the ladle refractories. The impact of the choice of slag composition or main fluidizers is clearly illustrated by the slags in Table 3. All of the slags have a high fluidity and are just saturated with respect to CaO and MgO ("creamy") and are therefore compatible with magnesia and dolomite refractories. Different slag composition choices, result in different desulfurisation potential and equilibrium sulfur content for the same conditions of oxygen content of the metal and slag mass. Using available phase diagrams or liquidus models, various combinations of fluidizing agents can be used so that an infinite number of slags with varying sulfide capacities can be designed to attain specific metallurgical goals.

Table 3. Compositions of slags that are just CaO or CaO and MgO saturated at 1600°C (Λ=optical basicity, Cs = sulfide capacity, and Sf = final sulfur)

Slag 1 Slag 2 Slag 3 Slag 4 % CaO 45 52 53 57 % MgO 17 10 13 8 % Al2O3 17 23 % SiO2 38 21 23 9 % CaF2 11 3 B5 Ratio 1.63 1.63 1.94 1.86 Λ 0.698 0.742 0.758 0.785 -Log Cs 3.01 2.38 2.14 1.76 Sf

* 0.0424 0.0285 0.0217 0.0121 *Using 100 tonnes of steel, 1000 kg of slag, Si = 0.05%, O = 15 ppm

Slags can also be “designed” for re-sulfurized steel grades, that is not to remove much sulfur – e.g. Slag 1 in Table 3. This usually requires substitution of MgO for CaO (and SiO2 for Al2O3) in order to generate slags that are still fluid but with a decreased thermodynamic sulfide capacity, while maintaining basicity for basic refractory compatibility. These slags will have lower CaO/SiO2 ratios that will increase the solubility of MgO in the slag. All these slags are 100% fluid at steelmaking temperatures and compatible with magnesia slag line refractories. We have so far only considered the thermodynamic factors that affect the potential for sulfur removal – i.e. the best case scenario of what can be achieved for the conditions after sufficient or infinite time. The kinetics or rate of sulfur removal are, however, also important and will be considered next.


4 Kinetics of sulfur removal in the ladle.

4.1 Rate of desulfurization While thermodynamic describe the driving force and extent or limits of sulfur transfer from metal to slag, the actual change achieved in the allotted treatment time depends on a number of kinetic factors. It has been established that mass transfer of sulfur in the metal to the slag-metal interface, limits the overall kinetics of sulfur transfer to the slag. This means that the rate at which sulfur is transferred from the bulk metal to the slag-metal interface and transfer across the boundary layer close to the interface, are the slowest and thus rate limiting step – which is mainly affected by stirring in the metal. This implies that the slag-metal reaction is relatively rapid and mass transfer in the slag is not limiting – i.e. a well-mixed liquid slag. The mixing of the metal is thus the main limiting parameter, as controlled by the inert gas stirring through porous plugs or electro magnetic stirring. The following equations describe the kinetic theory of sulfur removal and equation (11) describes the rate of change in sulfur content as a function of the sulfur content at time t (%S), the initial sulfur content (%S°) and the equilibrium sulfur content (%Seq). The driving force for sulfur transfer is the difference between the current sulfur content at time t and the equilibrium sulfur content or (%S-%Seq) and k is the overall kinetic constant.

)S%S(%kdt

S%d eq−−= (11)

Integration of equation (11) between time 0 and t results in the following:

kt)S%S(%)S%S(%ln

eq

eq−=

−°

− (12)

It should be clear that the kinetic constant determines how rapidly the sulfur content in the metal decreases or sulfur is transferred from the metal to the slag. The larger the kinetic constant, the quicker will the sulfur removal be. As the rate is controlled by mass transfer in the metal, the most important parameter determining the value of k, would be the stirring intensity in the metal. The magnitude of the kinetic constant can be determined from actual plant conditions by sampling and analysis of metal sulfur content over time intervals. It can also be estimated by published correlations, which relate the kinetic constant to the intensity of stirring as a result of a particular gas flow rate.

4.2 Correlations of kinetic constant (k) and gas flow rate. Two different correlations between gas flow rate and stirring intensity were found in literature. Both are based on the concept of stirring energy due to gas stirring, incorporating the displacement and drag effects of the rising bubbles. The stirring energy is then correlated by empirical measurements to first the mass transfer coefficient and then the kinetic constant or directly to the kinetic constant. While the correlations might be only strictly valid for the particular conditions under which the measurements were done, it is nevertheless a good estimate


of the expected rate of sulfur removal. Specific conditions which can impact the applicability are: • Type and position of porous plug in ladle. • Ladle dimensions and slag/metal weight. • Slag-metal emulsification behavior at the interface (slag fluidity). Extrapolation of these ‘empirical’ relations to other plants can only be viewed as a rough estimation and should be checked with detailed sampling and correlation. Extrapolation of these correlation for bottom stirring through purge plugs, to stirring by top-lance should be done with care. The stirring energy according to Sundberg can be calculated by:

+−=

top

bottomP

Pln

T2981*T*V*183.6]W[E (13)

and the mass transfer coefficient (ms) and overall kinetic constant (k) [Ishii et.al.]:

)Area/E(10*33.8]s/m[m 4S

−= (14)

+

ρ=−

Ls*WW1

WArea**m

]s[kslag

m

m

S1 (15)

Alternatively the relations by Plushkell as used by Turkdogan can be used:

=ε

top

bottom

m PP

lnW

T*V*183.6]ton/W[ (16)

ton/W60for)(*013.0][mink 25.01 <εε=− (17)

ton/W60for)(*10*8][mink 1.261 >εε= −− (18)

In the above, symbols have the following meaning:

V = flow rate of stirring gas [Nm3/min] T = temperature [K] ρ = density of liquid steel at T [kg/m3] Area = surface area of slag-metal interface [m2] Ptop = pressure above ladle [atm] Pbottom = Pressure at bottom of ladle = Ptop + ρgh [atm]

In general the correlations show that increased rate of desulfurization can be achieved by:

• Higher temperature (usually increases kinetics) • Higher gas stirring rate (V) • Decreasing the pressure above the bath by vacuum (Ptop)


The other parameters are related by the size of the ladle, for example increasing the planar surface area by increasing the ladle diameter, would decrease the ladle depth and consequently Pbottom also, for constant metal weight. The effect of slag mass (Wslag) in equation (15) might be contrary to ‘gut feeling’ at first glance, as higher slag mass seems to hinder desulfurisation by decreased rate (k). However, the slag mass also affects the equilibrium %S, so that the result is that it would take longer to reach the equilibrium sulfur content (lower k), but the equilibrium sulfur content is also much lower (according to (10)). In the second correlation, the kinetic constant increases much more drastically with stirring intensity above a critical value of 60 W/ton - equation (18), than below this critical intensity as in equation (17). This is due to emulsification of the metal-slag interface at higher flow rates (intensity of stirring), which increases the mass transfer at the boundary layer and the surface area significantly. At lower flow rates the metal-slag interface is relatively stagnant and mass transfer is slower. This is another reason why a fluid slag (creamy, but no major solid content) is important as the emulsification is important to achieve high rates of desulfurisation demanded by current production schedules. Another interesting aspect of the equations, is that decreasing the pressure above the vessel (e.g. in degasser) increases the stirring intensity for identical dimensions and gas flow rates – thus increases mass transfer and the rate of desulfurisation. However, it does not affect the sulfur distribution ratio or equilibrium sulfur content, just the rate at which it is achieved. For electro-magnetic stirring (EMS) emulsification of the slag-metal interface does not occur as readily as for gas stirring (physical effect of bubble is important) and consequently sulfur transfer is expected to be slower. The controlled variable for EMS would be the induction current (as gas flow rate is for gas stirring), but no correlation for sulfur removal could be found in literature. Data for inclusion removal for different EMS-current setting is available, but would be plant specific and does not allow correlation with stirring intensity (W/t) and desulfurization. The position of the purge plugs in the ladle can have a significant effect on the sulfur removal kinetics and thus accuracy of these correlations. Intermixing or formation of a slag-metal emulsion at the interface is promoted by the stir plug in the center of the ladle – this provides increased surface area and less boundary layer resistance for the desulfurisation reaction to occur. However, a central position also promotes the formation of dead zones in the lower part of the ladle and metal mixing is not efficient. Mixing efficiency is improved (mixing/homogenizing time reduced) by moving the purge plug away from the center of the ladle, also enabling the opening of an eye for alloy additions. The closer the plug is to the refractory lining however, the more slag line erosion is expected. Modeling work and practical experience suggests that the best position for a compromise between the above effects be with the stir plug on one third of the ladle diameter. (Gas stirring should also not be directly underneath an electrode position). The kinetic equations have been incorporated into the Baker Kinetic de-S model for estimation of desulfurization rates and correlation with plant data.


4.3 Change of %S with time. With knowledge of the overall kinetic constant, the sulfur content variation with time toward equilibrium can be determined according to equation (19) derived from equation (12).

kteqeq e)S%S(%S%S% −−°+= (19) The typical form of the exponential function in equation (19) is shown graphically in Figure 7 for different k-values, representing different stirring gas flow rates. The other parameters, the initial and equilibrium sulfur content are the same for both graphs.

0

0.005

0.01

0.015

0.02

0.025

0.03

0 5 10 15 20 25 30 35 40 45

Time (minutes)

%S

in m

etal

k=0.05 k=0.10

Figure 7: Example of %S versus time for different k-values.

Kinetic evaluation is useful to determine the how the important parameters (LS, slag volume, stirring rate, etc.) affect the time required to lower the sulfur content of the metal as tapped, to a certain aim value (specification). The time available for ladle treatment is limited and different scenarios with different the parameters can be evaluated theoretically with this calculation, before testing it in the plant. The sulfur change with time can also be evaluated for changed conditions – i.e. vigorous/hard stirring for 10 minutes, followed by alloy trimming and ‘softer’ rinsing stir for 15 minutes for inclusion flotation. The different conditions for each period can be determined and the expected sulfur change with time calculated for the two successive periods. This would be useful to determine the minimum required “hard stir time” to reach the specified %S after soft rinsing. The


conditions to limit sulfur removal (i.e. re-sulfurized grades), can also be evaluated with these relations. It is important to realize that the desulfurisation reactions continue to occur after treatment at the ladle furnace, even if at a lower rate due to reduced stirring. For a long delay before casting, the sulfur content can still reduce significantly after ladle treatment – can be critical for close spec or lower limit sulfur grades. This of course depends on how far from equilibrium the sulfur distribution is – i.e. the remaining driving force.

4.4 Additional comments on kinetics: • The conditions in the ladle will change with time (temperature, slag composition, activity of

oxygen) and will affect the driving force or Ls value and thus also the kinetics. If the changes are minor, average or constant values can be used for the calculations.

• The difference between the aO at the slag-metal interface and in bulk of the metal (where it is usually measured) can significantly affect sulfur removal.

• Excessive stirring and opening of a large “bubble eye” can lead to significant re-oxidation – affecting both the aO at the interface and slag composition (also by inclusion flotation).

• The effect of reducible oxides (FeO & MnO) is difficult to accurately include in the calculations and is expected to lower the sulfur distribution ratio by way of its effect on the interfacial oxygen activity. Significant reducible oxide content will result in fading of the deoxidizer and resulting increase in the bulk oxygen activity.

• Although not common, it is possible at high total sulfur in the system (sulfur load) and high Ls, that the sulfur content of the slag exceeds the maximum solubility and solid CaS is formed. It is believed that this will lead to a lowering of the effective Ls, due to lowering of the sulfide capacity (Cs) as the lime content of the liquid slag is reduced. It is not clear whether experimentally determined values for Cs are valid for conditions that would exceed S-solubility in the slag. This can also be a localized effect if poor mixing conditions exist in the slag phase.

• Initial sulfur content in EAF transfer slag and in ladle skulls or remaining slag can be important, as they contribute to the total sulfur content of the system (sulfur load) – especially if both low S and re-S grades are produced in the same shop (ladles).

5 Sulfur behavior in the EAF

5.1 EAF conditions and Cs. The basic theory for sulfur removal (Ls as in equation (9)) indicates that reducing conditions (low aO) and highly basic slags are best to transfer sulfur from metal to slag. These are the conditions during secondary metallurgical treatment in the ladle. This implies that sulfur removal in the EAF at oxidizing conditions and less basic slags will be less effective. The advantage in the EAF is the relatively large slag volume (in the region of 200 lb per ton of metal for EAF, vs. 20 lb per ton of metal in the ladle). Due to the large slag volume, the sulfur removed can still be significant, even at low LS. The ideal slag composition in the EAF is not dictated by desulfurisation, but by more important process aspects like foaming. Sulfur removal in the EAF


is limited by other process requirements and the best method of controlling sulfur will still be selection of input materials and good desulfurisation in the ladle.

Not much experimental data on sulfide capacity (Cs) of EAF type slags is available, as most investigations have concentrated on ladle slags. Generally the log Cs of an EAF slag will be much lower than that of a synthetic ladle slag due to lower CaO content and significant SiO2 and MgO content. The classic relations of logCs to optical basicity are not necessarily valid, as the correlations were developed for ladle slags. There is also uncertainty regarding the optical basicity values to be used for FeO and MnO, which are present in much large quantities than in ladle slags and can significantly affect the optical basicity (e.g. values of either 0.51 or 1.0 for FeO can be used). Recent research results have indicated that sulfur capacity increases with increasing FeO or MnO content, but this also depends on which specie is replaced in the slag (e.g. CaO versus SiO2). However, the oxygen activity of the system will also increase with increasing FeO content and the change in Ls as a result of the opposing changes is not quite clear.

5.2 Log Ls Turkdogan showed a correlation of sulfur data in terms of distribution ratio Ls for oxidizing processes, without consideration of sulfide capacity. The graphical correlation shows that the parameter kso = Ls*%O or ks = Ls*%FeO decreases with increasing (%SiO2 + 0.84*%P2O5). (For some of the data Al2O3 is included in the SiO2 term, with a coefficient of 1.18). The %FeO term is used as a measure of the oxygen activity in the EAF system. This agrees with the general theory that more basic slags are better desulfurizers or acidic slags are worse, however, the FeO content is also important. By application of these relations to plant data from a particular EAF-shop, Turkdogan showed that similar to the relations in the BOF, the conditions just before tap are still below equilibrium. A curve was fitted through Turkdogan’s plots of ks and the equations are:

For equilibrium: ks = 481.2 exp (-0.0939 * (%SiO2 +0.84*%P2O5)) (20)

For typical EAF: ks’ = 434.3 exp (-0.1077 * (%SiO2 + 0.84* %P2O5)) (21) As for the case in ladles, the practical slag composition is however, restricted by other process issues and in the EAF slag foaming is of major importance, as is refractory compatibility. By limiting the slag compositions to those known to be ideal for foaming properties in the EAF, the sulfur distribution is shown to vary between 1 and 3. For comparison, the slag compositions close to dual saturation were chosen and the ks and Ls values calculated and the results are shown in Table 4. Even though the ks value increases with increasing basicity, the FeO content also has to increase to maintain a liquid slag. More FeO is required to flux the slag, when the SiO2 content reduces as a result of the increase in basicity. These effects almost cancel out and the resulting Ls values do not increase much as shown in the table. The values between 1 and 3 are low compared to typical values of 50 to over 500 for ladle slag systems.


Table 4: Variation of Ls for dual saturation in EAF slags (* ’SiO2’ = %SiO2 + %Al2O3)

Basicity %FeO %MgO %CaO %'SiO2'* EAF-ks LS 1.5 15 13 43.2 28.8 20 1.3 1.65 20 11 43.0 26.0 26 1.3

2 26 9 43.3 21.7 42 1.6 2.5 31 8 43.6 17.4 66 2.1 3 34 7 44.3 14.8 89 2.6 3 45 6 36.8 12.3 116 2.6

The last two rows show a high basicity, defined as CaO/[SiO2 +Al2O3], with low and high FeO content. The correlation (21) shows an increase in ks for the lower SiO2 content, but the higher FeO content off-sets the change and the Ls does not change. The higher FeO content results in increase yield losses and probably departure from the ideal foamy slag composition. Therefore, in the practical range for EAF foaming slag compositions, the Ls will be between 1 and 3, with a value of 5 being a possible extreme case. The fraction of sulfur remaining for different Ls-values is indicated as a function of slag volume in Figure 8. At a slag rate of 0.15, which relates to 150 kg slag per tonne (1000 kg) of metal or alternatively 300 lb of slag per ton (2000lb) of metal, the fraction of sulfur remaining at equilibrium would be 0.7 for an Ls=3. This means that 30% of the sulfur was removed from the metal, which is still significant. Due to the constantly changing conditions in the EAF, equilibrium is rarely achieved and the application of theoretical calculations is less accurate, than in the ladle. Sulfur distributions determined in the plant, based on actual metal and slag samples can deviate to either side of the equilibrium calculations, due to kinetic limitations and late addition of sulfur via scrap melting or carbon injected.


0.4

0.5

0.6

0.7

0.8

0.9

1.0

0 0.05 0.1 0.15 0.2 0.25

Slag rate (ton slag per ton metal)

Frac

tion

S re

mai

ning

LS=1 LS=3 LS=5

Figure 8: Influence of Ls and slag rate on sulfur remaining/removed.

5.3 Additional comments - EAF: • Consider the sulfur content of the slag remaining in the furnace after tapping for sulfur

balance of the following heat. • Most carbon sources for charging and injection (especially coal) contain sulfur, which can be

a significant contributing factor. • The conditions in the EAF change drastically over time compared to the ladle furnace

(temperature, oxygen activity and slag composition), resulting in a “constant change” of the thermodynamic driving force for sulfur removal.

• The EAF (or other primary steel making process) is principally oxidizing and thus best suited to the removal of Si, Mn, C and phosphorous from the melt. The sulfur removed is important, but changing conditions to maximize sulfur removal, will negatively impact other process requirements.

6 References

"Ladle Metallurgy Principles and Practices", Fruehan RJ, ISS, Bookcrafters Inc. 1985

"Physicochemical properties of molten slags and glasses", Turkdogan ET, Metals Society London. 1983.

“Fundamentals of Steelmaking”, Turkdogan ET, The Institute of Materials, London, 1996.


“The Making, Shaping and Treating of Steel”, 11th edition, Steelmaking and Refining Volume, Fruehan RJ (Editor), 1998, AISE Steel Foundation, Pittsburgh PA.

“Slag Atlas”, Verein Deutscher Eisenhüttenleute, prepared by the Committee for Fundamental Metallurgy. 1981


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