Operations in Steel Teeming Ladle in Steel Melting Shop

Operations in Steel Teeming Ladle in Steel Melting Shop

The treatment of steel in the ladle has started around 1940s when the first ladle-to-ladle and ladle-to-ingot mould vacuum degassing processes for hydrogen removal has appeared on the scene. In the late 1950s, more efficient vacuum degassing units such as the Dortmund Hoerder (DH) and Ruhrstahl-Heraeus (RH) processes have become popular. In the middle of 1960s, degassing processes such as vacuum arc degassing (VAD), the ASEA-SKF process, and the vacuum oxygen decarburization (VOD) process for treating high-chromium steels were successfully implemented.

Converter processes such as the argon oxygen decarburization (AOD) process were introduced in the early 1970s. The AOD process is now the preferred route in the production of several specialty steels and stainless steels. Granulated flux injection into the liquid steel, combined with argon stirring, started in the early 1970s. This was soon followed by the application of cored-wire feeding of alloying elements for better control of composition and for inclusion morphology.

All the above developments have had a marked effect on the process of steelmaking, mainly with respect to the furnace in which the steel is produced. As an example, the implementation of ladle metallurgy (also known as secondary metallurgy) and its related aspects enabled the steelmakers to reduce the heat time in the primary steelmaking furnaces since there have been no need to perform any refining of the steel in the furnace. In addition, ladle refining and degassing have made it possible for the steelmaker to exert much tighter control over the properties of the final product through improved accuracy in composition as well as its cleanliness and by being able to control inclusion morphology.

Tapping of the liquid steel – During tapping of the liquid steel, air bubbles are entrained into the steel where the tap stream enters the bath in the steel teeming ladle. The quantity of air entrained into the steel increases with the increase of the free-fall height of the tap stream, as has been established with the help of water model studies. The entrainment of a gas such as air into a falling stream of liquid steel has been the subject of a number of studies. However, a reliable prediction of the quantity of air entrained into a stream of liquid steel during tapping is difficult because of the assumptions which are needed to be made.

The nitrogen contained in the air entrained by the steel is absorbed by the liquid steel depending on the extent to which the reaction ‘N2 (g) = 2[N]’ (equation 1) proceeds to the right, the symbol within the square brackets refers to the nitrogen dissolved in the liquid steel. It is well known that surface-active solutes such as oxygen (O2) and sulphur (S) hinder the kinetics of nitrogen absorption by the liquid steel. The higher the concentration of dissolved oxygen and / or sulphur, the lower is the extent of nitrogen (N2) absorption. This is shown in Fig 1a, where the effect of deoxidation practice on the nitrogen pickup during tapping of an electric arc furnace (EAF) is shown. For deoxidized steels, the average nitrogen pickup during the tapping is considerably higher than for non-deoxidized steels. The same effect is shown in Fig 1b, where the nitrogen pickup during tapping of 220-ton basic oxygen furnace (BOF) heats is shown as a function of the dissolved oxygen content for steels containing around 0.01 % sulphur.

Fig 1 Nitrogen pickup during furnace tapping  

The data in Fig 1a and Fig 1b are in complete agreement. Other sources contributing to the nitrogen pickup during or shortly after the tapping are (i) petroleum coke, when used for recarburization, and (ii) different ferro-alloys, particularly ferro-titanium (Fe-T), ferro-vanadium (Fe-V), and low-carbon and medium-carbon ferro-chromium (Fe-Cr).

Additions in the steel teeming ladle additions contain frequently moisture, which reacts with the liquid steel as per the reaction is ‘H2O = 2[H] + [O]’ (equation 2). It is seen from this equation that the extent of hydrogen (H2) pickup is more pronounced for a fully deoxidized steel in which the dissolved oxygen content is low. Amongst the different ferro-alloys, ferro-manganese (Fe-Mn) is probably the major contributor of hydrogen.

Carry-over of the furnace slag – It is normally unavoidable since a quantity of furnace slag is always carried over into the steel teeming ladle during the liquid steel tapping. The furnace slag normally contains a high concentration of FeO (ferrous oxide) and MnO (manganese oxide) and hence (in an untreated form) is not suitable for use as a refining slag. Accordingly, methods to minimize the quantity of furnace slag carry-over have been developed and implemented. These include locating the taphole in the barrel of the BOF, sometimes in conjunction with the use of a ceramic sphere to block off the taphole toward the end of the tapping.

In an EAF, the first improvement has been the use of submerged tapholes, later followed by the widespread implementation of eccentric bottom tapholes (EBT), which are now common in modern EAFs. In older shops, inadequately equipped for the control of furnace slag carryover, slag raking is frequently practiced for the removal of the furnace slag. A good slag-free surface is attainable by careful raking. However, raking is normally accompanied by a temperature loss of around 2.5 deg C per minute of treatment time and a metal loss of around 0.2 %. Raking times are typically of the order of 10 minutes. For good results, it is desired that the steel is tapped open, then is raked and covered with the synthetic ladle slag and finally deoxidized. As a result of furnace slag carry-over into the steel teeming ladle, oxidation of aluminum (Al) and silicon (Si) present in the ladle additions occurs through reactions with less stable oxides (e.g., FeO and MnO) present in the furnace slag.

Another consequence of furnace slag carry-over is phosphorus (P) reversion from the slag to the steel, particularly when the steel is fully deoxidized. For being able to predict the aluminum and silicon losses as well as the anticipated degree of phosphorus reversion, it is necessary to know the quantity of furnace slag carried over into the steel teeming ladle. A sensor can be used for measuring the depth of the slag layer in the ladle. The data provide accurate feed-back to the furnace operators and assist them in controlling the quantity of furnace slag carry-over.

Losses of aluminum and silicon – The reactions of aluminum and silicon dissolved in the liquid steel with the iron oxide and manganese oxide in the slag and with the fallen BOF skull can be represented by the general reaction ‘Fe(Mn)Ox + Al(Si) = Fe(Mn) + Al(Si)Ox (equation 3).

Using average molecular masses and assuming 80 % Fe3O4 (ferric oxide) for the composition of the skull, the empirical relation for the percentages of aluminum and silicon lost to the ladle slag for a 200-ton steel bath in the ladle is ‘[%Al + %Si]s = 0.0000011 x delta(%FeO + %MnO)Wfs = 0.0001 x Wsk’ (equation 4), where ‘Wfs’ and ‘Wsk’ are the weights (in kilograms) of the carried-over furnace slag and the fallen BOF skull respectively, and delta(%FeO + %MnO) is the decrease in the oxide contents of the furnace slag during tapping.

Phosphorus reversion – Furnace slag carry-over results normally in phosphorus reversion from the slag to the steel, particularly when the steel is fully deoxidized. The general relationship for the increase in steel phosphorus content ‘delta[%P]’, as a result of reversion from the slag is ‘delta[%P] = (%P) x (Wfs/Wb)’ (equation 5) where ‘(%P)’ is the phosphorus content of the furnace slag and ‘Wfs’ and ‘Wb’ are the weights of the carried-over furnace slag and the steel bath in the ladle respectively.

For bottom blown BOF and low-carbon heats made in EAFs with oxygen injection, typical values for ‘(%P)’ and ‘delta[%P]’ are around 0.3 and 0.003 respectively. Substitution of these values into equation 5 gives ‘Wfs/Wb = around 0.01’. In other words, when proper measures are taken for the prevention of excessive furnace slag carry-over, the average quantity of carried-over furnace slag is around 1 % of the steel tapped. This quantity of slag carried over during tapping corresponds to a slag thickness in a 200-ton ladle of 55 +/- 30 millimetres (mm), in general agreement with the plant observations.

In a study, it has been observed that phosphorus reversion is more likely to occur when the basicity, ‘%CaO / %SiO2’ of the carry-over slag is around 2 or less, and its FeO content is around 17 % or less. For FeO (ferrous oxide) contents of around 25 % or higher, the phosphorus reversion is noticeably less, provided the slag basicity exceeds 2 to 2.5. For the heats which are tapped open, to which only ferro-manganese and a small quantity of aluminum are added, the liquid steel is not sufficiently deoxidized to cause reversion of phosphorus. In fact, in some cases of tapping open heats to which 0.3 % to 0.6 % manganese is added, the phosphorus content of the steel decreases by around 0.001 % because of the mixing of the carried-over furnace slag with the steel during tapping.

Chilling effect of ladle additions – Ferro-alloys and fluxes added to the steel in the steel teeming ladle affect the temperature of the steel in the ladle, resulting normally in a decrease in temperature. The effect of different alloying additions, including coke, on the temperature of the liquid steel for an average bath temperature of 1,650 deg C is given in Tab 1.

Tab 1 Effect of alloying additions on the change in temperature of the steel in the steel teeming ladle for an average bath temperature of 1,650 deg C
Addition to give 1 % of alloying element at 100 % recoveryChange in steel temperature, Delta t
Coke-65 deg C
Ferro-chrome (50 %), high carbon-41 deg C
Ferro-chrome (70 %), low carbon-28 deg C
Ferro-manganese, high carbon-30 deg C
Ferro-silicon (50 %)Nil
Ferro-silicon (75 %)+ 14 deg C
Silico-manganese-6 deg C

The data in Tab 1 have been calculated from the heat capacities and heats of solution of the different solutes. It can be seen from Tab1 that ferro-silicon is the only ferro-alloy which, upon addition, does not result in a decrease in steel bath temperature. In fact, the use of ferro-silicon (75 %) results in an increase in temperature. This is because of the fact that the dissolution of silicon into liquid iron is exothermic, i.e., heat is liberated. Although ferro-silicon (75 %) normally costs more per unit weight of contained silicon than ferro-silicon (50 %), the use of the former material is justified under certain conditions, particularly when relatively large quantities are to be added and when the steel melting shop has no or limited reheating facilities such as ladle furnace.

For aluminum-killed steels, the exothermic heat of the deoxidation reaction is to be taken into account when calculating the effect of the aluminum addition to the steel teeming ladle on the steel temperature. For example, when a steel containing 600 ppm (parts per million) of dissolved oxygen is deoxidized with aluminum, the heat generated by the deoxidation reaction results in a change in steel temperature of +19 deg C. In other words, when the steel in the steel teeming ladle is deoxidized with aluminum, the decrease in steel temperature as a result of tapping is less by 19 deg C.

Flux and slag conditioner additions decrease the temperature of the liquid steel in the ladle. The effect of these additions on the change in the steel temperature as determined from the heat capacity data is given in Tab 2. When tapping aluminum-killed steels to which typical ladle additions of 10 kilograms of lime and pre-fused calcium aluminate (CA) per ton of steel are made, the decrease in steel temperature during tapping to a pre-heated ladle is 55 deg C to 75 deg C. The heat loss because of the flux addition is around balanced by the heat generated by the deoxidation reaction with aluminum. Hence, in this particular practice the heat losses are almost entirely from the radiation and the conduction into the ladle lining.

Tab 2 Change in steel temperature in the steel teeming ladle as a result of different flux additions at a rate 1 kg/ton
Flux added (1 kilogram per ton)Change un steel temperature Delta T,
SiO2 (Silica)-2.5 deg C
CaO (Lime)-2 deg C
MgO (Magnesia)-2.7 deg C
CaO.MgO (Calcined dolomite)-2.3 deg C
CaO.Al2O3 (Calcium aluminate)-2.4 deg C
CaF2 (Calcium fluoride)-3.2 deg C

Pre-heating of steel teeming ladle – In the majority of the BOF and EAF steel melting shops, the ladle lining consists of high-alumina (Al2O3) bricks (70 % to 80 % Al2O3), while the slag line consists of magnesia (MgO) bricks, normally containing around 10 % carbon (C) and small quantities of metallic additions such as aluminum (Al), magnesium (Mg), or chromium (Cr) for minimizing the oxidation of carbon. For several EAF shops, the lining is based on sintered dolomite.

The thermal properties of different refractories used in the steel teeming ladles are given in Tab 3 and Fig 2a. It is seen from Fig 2a that magnesia has a considerably higher thermal conductivity than the other refractory materials. Magnesia also has a higher heat storage capacity than the other materials, as shown in Tab 3. The properties of dolomite are similar to those of magnesia. Because of the relatively high thermal conductivity of the refractories used in the steel teeming ladles, pre-heating of the ladle prior to its use is needed in order to avoid excessive heat losses during tapping and during subsequent refining operations.

Tab 3 Heat storage capacity per unit volume of different refractory materials at 1,200 deg C and 1,500 deg C relative to magnesia (= 100 %)
Material 1,200 deg C1,500 deg C
70 % alumina (Al2O3)83 %85 %
50 % alumina (Al2O3)79 %82 %
Fireclay73 %74 %

A study has been carried out on the effect of ladle refractories and practices on the control of the steel temperature. The study has developed a mathematical model to simulate a ladle which is used to supply liquid steel to a continuous casting machine (CCM). The model has been used for evaluating the temperature increase of the hot face and the cold face of a ladle during pre-heating. This is shown in Fig 2b, from which it is seen that the temperature of the hot face increases rapidly with time. However, the cold face temperature does not exceed 100 deg C after pre-heating for around 5.5 hours when the ladle is initially cold and dry.

Fig 2 Thermal conductivity of refractory bricks and lining temperatures

The rate of rise of temperature of the hot face depends on the distance from the ladle top to the burner wall as well as on the thermal input from the pre-heater. A rapid heating rate is to be avoided since (i) rapid heating results in a non-equilibrium temperature profile, i.e., a steep temperature gradient adjacent to the hot face, (ii) rapid heating causes extreme shell stresses, and (iii) the thermal shock resistance of the brick is not high enough for withstanding a rapid heating rate.

Another measure of the thermal condition of the ladle is the total heat content of the ladle brick. As shown in Fig 3a, the brick continues to absorb heat at a considerable rate for up to 12 hours. At this time, the ladle brick contains in excess of 90 % of the maximum heat content which can be achieved once steady state has been reached after 17 hours to 18 hours of pre-heating. After pre-heating, the ladle is moved to the steelmaking furnace. This causes a decrease in the temperature of the hot face. The temperature decrease is to be taken into account when adjusting the tapping temperature. The needed adjustment in the tapping temperature is shown in Fig 3b as a function of the time elapsed between the end of pre-heating and the start of tapping for two hot face temperatures.

Fig 3 Total heat content of a lining and effect of cooling time

Ladle free open performance – Upon arrival of the ladle at the tundish of the continuous casting machine, it has to be opened to allow the steel to flow into the tundish. When the ladle slide gate is stroked open and steel starts to flow without operator assistance, the procedure is classified as ‘free open’. If poking or oxygen lancing is required to open the ladle, the process is classified as ‘assisted open’. When all attempts to open the ladle are unsuccessful, it is classified as ‘non-free open’. Several factors determine the opening performance of a ladle. The most dominant factor is the residence time of the steel in the ladle, followed by the ladle pre-heat practice, the elapsed time between the end of stirring and the opening of the ladle at the CCM, the cycle time for an empty ladle, and the argon stirring practice.

The residence time in the ladle is defined as the elapsed time between tapping and the opening of the ladle at the tundish. However, longer residence times than normal can be there whenever problems are encountered at the CCM. When the residence time in the ladle increases and become more than normal, then a considerable decrease in free open performance takes place. This is since long residence times in the ladle cause partial sintering of the nozzle fill material because of the extended exposure to the liquid steel temperatures.

On examining the data of ladle pre-heat practices for ladles returned to service following repair, it can be seen that the free open performance is worse for repaired ladles compared with those which follow the normal cycle, indicating an inadequate pre-heat practice for the repaired ladles.

The time elapsed between the end of stirring at the trim station and the opening of the ladle at the tundish is important. The longer the time without stirring, the worse is the free opening performance.

The empty ladle cycle time is the elapsed time between closing the ladle at the CCM at the end of a casting and the time when the ladle is filled again at the tapping of the furnace. The normal practice in the steel melting shops is not to pre-heat the ladles when they are being rotated between the CCM and the primary steelmaking furnace. However, the rate of rotation can vary considerably because of the delays. The data from steel melting shops show that an empty ladle cycle time longer than two hours leads to a decrease in free open performance.

These days, majority of the ladles are equipped with bottom plugs for argon bubbling. Whenever the plug is out of commission, it becomes necessary to provide the needed stirring through a top lance. The data from different steel melting shops show that top stirring with argon results in a slightly decreased free open performance.

Stirring in ladles – For achieving a homogeneous bath temperature and composition, the liquid steel in the ladle is normally stirred by means of argon gas bubbling. For moderate gas bubbling rates, e.g., less than 0.6 N cubic metre per minute (cum /min), porous refractory plugs are used, normally mounted in the bottom of the ladle. A schematic illustration of a porous plug assembly in the ladle bottom is shown in Fig 4.

 Fig 4 Porous plug assembly and different types of porous plugs

Fig 4 also shows some typical examples of different plugs. As can be seen in Fig 4, porous plugs have either a conical or a rectangular shape. The conically shaped plug is easier to change in case the plug wears out before the lining. Rectangular plugs are geometrically compatible with the surrounding bricks and can be used to advantage in cases where the plug life is comparable to that of the lining. The performance and life of isotropic plugs can be improved by producing the element in two or three components stacked together with metal inserts.

The main advantage of the so-called directional-porosity or capillary plug, shown at 5 and 6 in Fig 4, is that the plug can be made of the same dense refractory as the lining brick, or even denser. This results in increased hot compression strength, higher resistance to erosion, and a longer service life. Disadvantages of capillary plugs are that they are more prone to infiltration by liquid steel because of the loss of argon gas pressure.

Some steel melting shops utilize electro-magnetic induction stirring in the ladles. Some reported features of induction stirring in the ladle include better stirring homogeneity (especially near the ladle bottom), the ability to reverse the direction of the stirring forces (useful for alloy additions), and stirring without breaking the slag cover and exposing steel to the ambient oxidizing atmosphere. These benefits are offset by the high capital cost, including ladles equipped with stainless steel panels comprising at least one-third of the ladle shell and the need for auxiliary gas stirring for adequate removal of hydrogen.

Stirring power and mixing times – Homogenization of bath temperature and composition by gas bubbling is mainly caused by the dissipation of the buoyant energy of the injected gas. The thermodynamic relationship describing the effective stirring power of a gas has been derived by Pluschkell. The equation for the stirring power is derived from Pluschkell’s relationship which is given by the equation ‘e = 14.23(VT/M) x log[(1+H)/1.48Po]’ (equation 6), where ‘e’ is the stirring power in W/ton, ‘V’ is the gas flow rate in N cum/min, ‘T’ is bath temperature in  Kelvin, ‘M’ is the bath weight in ton, ‘H’ is the depth of gas injection in metre, and ‘Po’ is the gas pressure at the bath surface in atmosphere.

The stirring time for achieving 95 % homogenization is defined as the mixing time ‘t’. Several experimental and theoretical studies have dealt with mixing phenomena in gas-stirred systems. The relationship which expresses the mixing time ‘t’, in terms of the stirring power ‘e’ (W/ton), ladle diameter ‘D’ (metre), and the depth of injection ‘H’(metre), is given by ‘t(second) = 116(e)to the power -1/3 x [(D)to the power 5/3 x 1/H)]’ (equation 7).  Mixing times calculated from equation 6 and equation 7 are shown in Fig 5 for the simplified case of D = H. The mixing times shown in Fig 5a are in good agreement with those calculated from other correlations. It can be seen from Fig 5a that a 200-ton heat is homogenized in 2 minutes to 2.5 minutes after bubbling with argon at a flow rate of 0.2 N cum/min.

Fig 5 Gas injection rate and effect of gas flow rate

In a study on the effect of the location of the bottom stirring plug on mixing time, it has been found that the mixing time is decreased by placing the bottom plug off-centre, e.g., at mid-radius. A stirring plug placed in the centre of the ladle bottom generates a toroidal loop of metal flow in the upper part of the bath while a dead zone is created in the lower part, resulting in longer mixing time. Eccentrically located bottom plugs give rise to extensive circulation of metal through-out the entire bath, avoiding dead zones and leading to shorter mixing times.

Slag-metal reaction rates in gas-stirred liquid steels – A number of experimental studies have been done for investigating slag-metal gas reactions in gas-stirred ladle systems under a variety of experimental conditions. For the majority of the slag-metal reactions, the rates are controlled mainly by mass transfer of the reactants and products across the slag-metal interface. In stirred systems such as a liquid steel bath in a ladle stirred by argon, the slag-metal interfacial area is affected by the degree of agitation in the bath, which, in turn, is determined by the stirring power.

Desulphurization – During desulphurization of the liquid steel in the ladle, the mixing of slag and metal is achieved by argon bubbling, and the rate of desulphurization is described by ‘ln{1 + 1/Ls x (Wm/Ws) x [%St/%So] – 1/Ls x (Wm/Ws)} / {1 + 1/Ls x (Wm/Ws) = (m x Ap/Ws)t’ (equation 8).

The overall rate constant is related to the average mass transfer coefficient ‘ms’, the slag-metal interfacial area ‘A’, and the steel bath volume ‘V’.  by the following expression ‘ks = ms x (A/V)’ (equation 9). It has been shown that ‘ks is proportional to (e) to the power n’ (equation 10). where the exponent ‘n’ can vary between 0.25 and 0.30, depending on the specific system under consideration.

From pilot plant tests with 2.5-ton heats for studying desulphurization, it has been found that at moderate gas bubbling rates, corresponding to ‘e’ less than 60 W/ton, there is little or no slag-metal mixing, hence, the rate of desulphurization is slow. For higher stirring rates corresponding to ‘e’ higher than 60 W/ton, better mixing of slag and metal is achieved, and the rate constant for desulphurization is increased accordingly. The results of these experiments are given in Fig 5b, from which the approximate relationships between the overall rate constant and the stirring power are derived are (i) ks per minute = around 0.13(e) to the power 0.25, for the value of ‘e’ higher than 60 W/ton, and (ii) ks per minute = around 0.00000008(e) to the power 2.1, for the value of ‘e’ higher than 60 W/ton (equation 11).

It is to be noted that these are empirical correlations. The value of ‘ks’ depends on the energy dissipation per unit area at the slag-metal interface, the properties of the slag, and the quantity of the slag. The abrupt change in overall rate constant for values of the stirring power at around 60 W/ton is explained by the fact that an increase in the energy input rate results in increased emulsification of slag and metal, leading to an increase in interfacial area ‘A’ which, in turn, increases the value of ‘ks’.

Dephosphorization – The removal of phosphorus from the liquid steel by the ladle slag is governed by the same rate equation as that for sulphur removal. Hence, the overall rate constant for dephosphorization is expected to have the same form as per the equation 9. That this is, in fact is with the case in which a study has been done for the dephosphorization in the ladle with CaO – CaF2 – FeO slags in a 50-ton VAD/VOD as well as in a 250-ton ladle furnace facility.

The overall rate constant can be represented by the approximate relationship ‘kp per minute = around 0.019 (e) to the power 0.28’ (equation 12). This expression is similar to that for ‘ks’, valid for ‘e’ less than 60 W/ton (equation 11). It is not clear why no abrupt change in ‘kp’ has been observed for ‘e’ higher than 60 W/ton, as in the case of desulphurization (Fig 5b).

It is interesting to note that in a study, it has been observed that mass transfer between metal and slag is hindered when the stirring plug in the ladle bottom is located off-centre. A stirring plug located in the centre results in increased slag-metal emulsification with increasing gas flow-rate. Eccentrically-located stirring plugs create a slag-free zone, the so-called eye, close to the ladle wall. This affects the detachment of slag particles from the main slag phase and results in decreased emulsification. The ultimate choice of location of the stirring plug in the bottom of the ladle is, hence, appear to be determined by which aspect of stirring is more important for a given operation i.e., good mixing characteristics, or the ability to achieve rapid desulphurization and / or dephosphorization. In majority of the cases a compromise is needed to be struck.

Effect of stirring on inclusion removal – One of the objectives of stirring the steel in the ladle is the removal of non-metallic inclusions. In a study done on the deoxidation kinetics of aluminum-deoxidized steels in 20-kilogram melts as well as in a 50-ton ASEA-SKF furnace, a model has been developed for inclusion removal based on the postulate that the decrease in total oxygen content is determined by the coalescence of oxide particles as the rate-controlling step. The model is in essence a combination of a coalescence theory and an algorithm for turbulent recirculatory flows.

In another study, a fluid mechanical model for inclusion removal from liquid steel has been presented. As per this model, the total oxygen content after stirring time ‘t’ is given by ‘(Ct-Cf)/(Ci-Cf) = (e) to the power -at’, (equation 13), where ‘Ct’ is the total oxygen content after the stirring time ‘t’, ‘Cf’ is the final total oxygen content after long stirring times (steady state), ‘Ci’ is the initial total oxygen content, and ‘a’ is the time constant for inclusion removal.

It is to be noted that the equation 12 is an extremely simplified expression. The rate of inclusion removal depends on several factors, including the inclusion type, refractory type, and exact stirring conditions etc. From experiments with an inductively-stirred 140-ton liquid steel employing a range of values for the specific stirring power (e), the approximate relationship which can be obtained is ‘ ‘a’ per minute = around e/27’ (equation 14). This expression is an approximation and is valid only for moderate induction stirring. If it is assumed that the final steady-state total oxygen content, ‘Cf’, is small compared to ‘Ci’ and ‘Ct’, the combination of equation 13 and equation 14 gives ‘a’ per minute is equal to around (e) to the power -et/27’.

Reheating of the liquid steel bath – The ever-increasing pressure for lowering of the operating costs and increasing the efficiency has made it necessary to make effective use of primary steelmaking furnaces, and implement sequential continuous casting. These factors have prompted the installation of facilities for steel reheating, necessary for the additional time needed for steel refining and the adjustment of the temperature of the steel for uninterrupted sequential casting. The two methods for reheating steel in the ladle are (i) arc reheating, and (ii) reheating by injecting oxygen and aluminum or silicon.

Over time, several types of furnaces for arc reheating have been developed and commercialized. Some of the designs also have the capability for degassing of the steel. An important issue in arc reheating of a steel bath is whether the thermal energy which is supplied at, or near, the surface of the melt can be dispersed rapidly enough such that no significant temperature gradients are created within the steel in the ladle.

In a study, it has been estimated that, in the absence of agitation, the Biot number (the ratio of the thermal resistance for conduction inside a body to the resistance for convection at the surface of the body) is of the order of 300, indicative of significant temperature gradients in the steel bath. The Biot number is defined as ‘NBi = hL / keff’, where h is the heat transfer coefficient between the arc and the bath, keff is the thermal conductivity of the steel and L is the bath depth. In systems agitated either by induction stirring or gas bubbling, the Biot number is estimated to be of the order of 5, indicative of small temperature gradients even in agitated systems. Once heat supply is discontinued, the temperature in gently agitated baths is expected to become uniform quite rapidly. In arc reheating the primary cost factor is electric power, followed by electrode and refractory costs.

The heating efficiency ‘n’ of arc heating is defined as ‘n = delta Tact / delta Tth = 0.22(delta Tact / E’ (equation 15), where ‘delta Tact’ is the actual temperature increase of the bath in deg C, ‘delta Tth’ is the theoretical temperature increase of the bath for 100 % thermal efficiency in deg C, and ‘E’ is the energy consumption in kWh/ton. The heat capacity of liquid steel is 0.22 kWh/ton deg C; i.e., for 1 ton of liquid steel ‘delta Tth = E/0.22’.

The heating efficiency increases with increasing bath weight, as shown in Fig 6a. The data represent overall averages comprising a range of heating times. For minimizing refractory consumption, heating times in ladle furnaces are kept as short as possible, typically around 15 minutes. Further measures for shortening the reheating time and hence minimizing refractory erosion are (i) the use of a large-capacity transformer, e.g., 35 MW to 40 MW for a 200-ton to 250-ton heat, (ii) submerged arcing in the slag layer, (iii) argon stirring through a bottom porous plug at a flow-rate of around 0.5 N cum/min, and (iv) a slag layer thickness of around 1.3 times the length of the arc.

Fig 6 Heating efficiency and electrode consumption

Normally, electrode consumption in ladle furnaces increases with increasing cross-sectional current density and heating time. The trends are shown in Fig 6b. For one of the steel plants, the average electrode consumption is 0.2 kg/ton for typical total reheat times of 20 minutes. and average current densities of around 35 amperes per square centimetre (35 A/sq-cm).

Refractory consumption – The refractory materials used in ladle furnace linings are similar for the majority of the steel melting shops and are applied in the same configurations, i.e., slag-line, bottom, and barrel. Hence, a comparison of refractory consumption data can be made. Fig 7a shows that the ladle life, expressed as the number of heats processed in a given ladle, increases with increasing bath weight. It is to be noted that the ladle refractory consumption in any given shop is strongly affected by the specific operating practice, as reflected by the considerable scatter in the data in Fig 7a. The submerged arc heating in the slag layer results in lower refractory consumption.

Fig 7 Ladle life and steel temperature increase

Reheating by oxygen Injection – Liquid steel can be reheated by oxidizing aluminum and / or silicon by means of oxygen injection through a lance. The heat generated for the reactions (i) ‘4Al (RT) + 3O2 (RT) = Al2O3 (1,630 deg C)’ (equation 16), and (ii) ‘Si (RT) + O2 (RT) = SiO2 (1,630 deg C)’ (equation 17) are 27,000 kJ/kg of aluminum for the reaction at equation 16, and 28,500 kJ/kg of silicon (in Fe-75%Si) for reaction at equation 17. The enthalpies are calculated from the thermodynamic data, taking into account that the reagents, aluminum and oxygen, are to be heated from room temperature (RT) to the temperature of the bath 1,630 deg C. On the basis of 100 % thermal efficiency, the bath temperature can be raised by 50 deg C when 1 N cum O2 / ton of steel is injected together with 1.46 kg of aluminum per ton or by injecting 1.2 N cum O2 / ton together with 1.85 kg Fe-75%Si per ton.

Reheating of steel in the ladle with submerged oxygen injection is being practiced at several steel plants. In a study on reheating by submerged injection of oxygen into 270-ton heats, the data shows that the temperature increase as a function of the specific quantity of oxygen (N cum/ton of steel) injected can be obtained. This is shown in Fig 7b, line (b). A comparison of the presented data with the maximum attainable temperature-increase for 100 % thermal efficiency, line (a), indicates that reheating by means of submerged oxygen injection is around 70 % efficient. The data of temperature-increase in a 160-ton RH-OB vessel is indicated by line (c) in Fig 7b and it indicates an average thermal efficiency of 20 % to 30 %. Fruehan has quoted a reheating efficiency of around 80 % for the RH-OB operation at the Oita works of the Nippon Steel Corporation. Data for a 245-ton RH-KTB vessel, in which the oxygen is supplied through a top lance instead of through submerged tuyeres as in the RH-OB, are indicated by line (d) in Fig 7b. The thermal efficiency for the RH-KTB process appears to be similar to that for submerged oxygen injection into the ladle.

A comparison of total oxygen contents measured in slabs cast from oxygen-reheated heats and heats which have not been reheated has shown no significant differences between the two sets of values. In addition, a study has compared the inclusion ratings in rail-grade steels produced from oxygen reheated steels and heats which have not been reheated and has found no significant differences between the two. These studies have recommended the addition of a synthetic ladle slag after reheating, followed by a thorough argon stirring for floating out the alumina inclusions so that they can be dissolved in the slag.

A study on the effect of reheating with aluminum and oxygen in the RH-OB process on the cleanliness of the final product, an increase in the total oxygen content during and shortly after oxygen blowing has been observed. However, the total oxygen content in the final product has been found to be similar for steels treated with oxygen and aluminum when compared with those treated without oxygen blowing in the RH, provided the total bath circulation time in the RH-OB is long enough.

Refining in the ladle – The refining of steel in the ladle is broadly defined as comprising of several operations namely (i) deoxidation, (ii) desulphurization, (iii) dephosphorization, (iv) controlled additions of alloying elements, and (v) inclusion modification.

Deoxidation – The first step in the refining sequence in the ladle is normally the deoxidation of the steel with ferro-manganese, ferro-silicon, silico-manganese, and aluminum. There are three categories of steel deoxidation namely (i) steel deoxidized with ferro-manganese to yield 100 ppm to 200 ppm of dissolved oxygen (these are normally resulphurized steel grades), (ii) semi-killed steels deoxidized with (a) Si / Mn to yield 50 ppm to 70 ppm of dissolved oxygen, (b) Si / Mn / Al to yield 25 ppm to 40 ppm of dissolved oxygen, and (c) Si / Mn / Ca to yield 15 ppm to 20 ppm of dissolved oxygen, and (iii) killed steels deoxidized with aluminum to yield 2 ppm to 4 ppm of dissolved oxygen.

Deoxidation in the presence of synthetic slags – The practice of refining steel in the ladle has made it possible to deoxidize the steel partially with ferro-manganese, ferro-silicon, and / or silico-manganese followed by a final deoxidation with aluminum. Such a practice has several advantages, including minimization of nitrogen pickup during tapping, minimization of phosphorus reversion from the carried-over furnace slag, and minimization of aluminum losses because of the reaction with carried-over furnace slag. These days, the use of synthetic slags in the ladle is an integral part of ladle metallurgy because of the needs necessary to produce ultra-clean steels, frequently combined with a demand for extra low sulphur contents.

The concept of using synthetic slags in the ladles dates back to the 1930s, when the Perrin process was developed for the improved deoxidation of open hearth or Bessemer steel with ferro-manganese or ferro-silicon by tapping the steel on a molten calcium alumino-silicate slag placed on the bottom of the tap ladle. The dissolution of the deoxidation products such as Mn(Fe)O or manganese silicates in the calcium alumino-silicate slag lowers their thermodynamic activity, hence increasing the extent of deoxidation.

Partial deoxidation with ferro-manganese – A study has described the results obtained for deoxidation with ferro-manganese in several plant trials. When deoxidizing with ferro-manganese, the deoxidation product is Mn(Fe)O, the activity of which is lowered in the presence of a calcium aluminate slag. The change in dissolved oxygen and manganese contents during tapping of a 200-ton heat of steel, to which 1,800 kg of lime-saturated calcium aluminate and ferro-manganese is added when the tap ladle is 12.5 % full, is shown schematically in Fig 8a. Upon addition of the ferro- manganese, the small quantity of steel present in the ladle is almost completely deoxidized, resulting in around 1.6 % manganese in the steel. As the ladle is filled, the dissolved manganese is consumed by the deoxidation reaction and decreases to around 0.32 % when the ladle is full, and the residual dissolved oxygen content is reduced to around 300 ppm from the original 650 ppm at the beginning of tapping.

Fig 8 Change in dissolved manganese and oxygen contents and partial deoxidation of steel

The results obtained using this deoxidation practice in EAF and the bottom blown BOF steel melting shops are shown in Fig 8b for steels containing less than 0.003 % of aluminum and silicon each. Without calcium aluminate slag addition to the tap ladle, i.e., deoxidation with manganese and iron only and pure Mn(Fe)O as the deoxidation product, the concentration of dissolved oxygen in the steel follows the broken line in Fig 8b. In the EAF trial heats, there has been no argon stirring in the ladle during furnace tapping. Yet it has been found that the slag-aided partial deoxidation of the steel achieved during tapping is close to the levels determined by the slag–metal equilibrium. This observation has led to the conclusion that there is sufficient mixing of slag and metal for promoting relatively rapid deoxidation in the ladle during tapping.

Deoxidation with silico-manganese – It is well known that the deoxidation of steel with manganese and silicon together leads to lower dissolved oxygen contents than the deoxidation with either of these elements alone. This is since the activities of the oxides in the deoxidation reaction [Si] + 2(MnO) = 2[Mn] + (SiO2) (equation 18) are less than unity. The symbols within the square brackets refer to species dissolved in the steel, and those within parentheses refer to species in the manganese silicate phase. By making use of the oxide activity data in the MnO-SiO2 system together with the thermodynamic data for reaction as per equation, a study has computed the equilibrium state pertaining to the deoxidation with silico-manganese.

When the deoxidation with silico-manganese takes place in the presence of a small quantity of aluminum dissolved in the steel, the deoxidation product is molten manganese alumino-silicate, and the resulting dissolved oxygen content is around 50 ppm for a steel containing roughly 0.8 % manganese and 0.2 % silicon. This is around half the value in a steel deoxidized with silico-manganese and not containing aluminum. This is since the activities of MnO and SiO2 are lowered further in the presence of the alumino-silicate phase. As an example, it is possible to decrease the dissolved oxygen content to around 20 ppm by means of deoxidation with silico-manganese together with the addition of 1,000 kg of pre-fused calcium aluminate to a 200-ton heat of steel.

For ladle slags containing a high percentage of alumina, there is some reduction of the alumina by the silicon in the steel. The data for steel containing aluminum and silicon in equilibrium with calcium aluminate slags containing around 5 % silica are shown in Fig 9a. It can be seen from Fig 9a that considerable pickup of aluminum from the slag can be expected if the steel is initially low in aluminum, e.g., 0.01 % Al, and contains around 0.2 % silicon. The final stage of deoxidation of the steel in the ladle is determined by the quantity of aluminum recovered from the slag.

Fig 9 Aluminum and silicon in equilibrium in steel and dissolved oxygen content in 0 05 % C steel

Deoxidation with calcium / silicon – Semi-killed steel deoxidized with silico-manganese can be deoxidized further with calcium / silicon, e.g., by injection of calcium silicide (Ca-Si) in the form of cored wire. The deoxidation of 50-kg steel melts at 1,620 deg C with silico-manganese and varying quantities of Ca-Si is the subject of one study. Some of the results are shown in Fig 9b, from which it is seen that the dissolved oxygen content in a low-carbon steel deoxidized with silico-manganese can be lowered from around 85 to around 55 ppm by adding 2.5 kg Ca-Si per ton.

Similar observations have been made in a series of plant trials in which Ca-Si cored wire has been injected into 60-ton heats of steel. The lower dissolved oxygen content achieved after the addition of Ca-Si is a result of the formation of calcium manganese silicate as the deoxidation product, further decreasing the activities of MnO and SiO2.

Deoxidation with aluminum – A number of experimental data, got by the emf (electro-motive force) technique, exist on the solubility product of Al2O3 in pure liquid iron. The reported values for 1,600 deg C are in the range ‘[%Al]square × [%O]cube = 9.77 × (10) to the power -15 to 1.2 × (10) to the power -13 (equation 19). The higher values are from earlier study when the interference with the emf readings caused by partial electronic conduction in the electrolyte of the emf cell has not been well recognized. In the most recent study, the emf readings have been corrected for electronic conduction, leading to the value of 9.77 x (10) to the power of -15 for the Al2O3 solubility product at 1,600 deg C.

In addition, the study has done a number of emf measurements in inductively-stirred iron melts in contact with CaO-Al2O3 slag mixtures in which the activity of alumina has been less than unity. The study has observed that the aluminum–oxygen relationship in these melts has indistinguishable from that in iron melts in equilibrium with pure alumina. Similar observations have been reported in another study. These experimental findings indicate that even in the presence of a calcium aluminate slag, the dissolved oxygen content in steel is controlled by the alumina inclusions always present in the interior of the bath.

In ladle metallurgy operations, where the steel is frequently covered with a calcium aluminate slag containing minor quantities of magnesia and silica, it is hence to be expected that the final dissolved oxygen content is controlled by the alumina inclusions dispersed throughout the bath. A measurable decrease in dissolved oxygen content as a result of treatment with calcium aluminate slag is not to be expected. A decrease in dissolved oxygen content can be expected only if substantially all the alumina inclusions have been modified to calcium aluminates, e.g., by calcium treatment.

Desulphurization In certain steel grades, such as those used in line pipe applications, a very low sulphur content is needed, i.e., 20 ppm or less. These low sulphur contents can be achieved only by steel desulphurization in the ladle in the presence of a calcium aluminate slag when the steel is fully killed. The governing reaction is ‘2/3[Al] + [S] + (CaO) = (CaS) + 1/3(Al2O3)’ (equation 20) where the symbols within brackets denote species dissolved in the steel and those within parentheses refer to species dissolved in the slag phase.

The change in free energy for reaction, based on the recent data on the solubility product of alumina in liquid iron, is given by ‘delta G0 = -319,343 + 111.3T, J/mole’ (equation 21). This gives for the equilibrium constant of reaction (equation 20) where ‘log K = (16,680/T) – 5.813’, and ‘K = [aCaS x (aAl2O3 to the power 1/3] / [aCaO x aS x (aAl) to the power 2/3]’ (equation 22).

The oxide and sulphide activities are relative to the respective pure solid phases, and the activities of aluminum and sulphur dissolved in liquid iron are defined such that for dilute solutions, below around 0.5 %, the activities can be replaced by their respective concentrations in weight percent. For a given slag composition, the activities of the oxides are fixed and can be incorporated in the equilibrium constant while the sulphur content of the slag, (%S), is proportional to the sulphide activity, ‘aCaS’. Hence, the equilibrium constant can be replaced by a pseudo-equilibrium constant ‘Ks = (%S) / [%S] x [%Al] to the power -2/3 (equation 23).

For a given aluminum content of the steel, the ratio LS = (%S) / [%S] is a function of the ladle slag composition at a given temperature. This is shown in Fig 10a, where lines of equal Ls-values are projected on the Al2O3-CaO-SiO2 phase diagram for slags in equilibrium with steels containing 0.03 % aluminum. For different aluminum contents, the LS -values in Fig 10a are to be multiplied by the factor (%Al / 0.03) to the power 2/3.

Fig 10 Iso-sulphur and Iso-phosphorus distribution ratios

Depending on the specific operating conditions, the range of compositions of ladle slags normally used is the mixture of Al2O3 – 20 % to 40 %, CaO – 35 % to 55 %, MgO – 8 % to 15 %, and SiO2 – 10 % to 15 %, together with minor quantities of FeO and MnO. In some cases, CaF2 is added to the ladle slag.

Desulphurization rate – For reaction (equation 20) to proceed fast so that the needed degree of desulphurization can take place within a practical time span, good mixing of steel and slag is necessary. The rate at which sulphur can be removed is, hence, strongly affected by the gas flow-rate or the stirring power density. As the capacity of typical ladle slags to absorb sulphur is high, the rate of desulphurization is controlled by mass transfer in the liquid steel, and the rate of desulphurization is described by equation 8.

In a study, the relative sulphur removal ‘R’ has been calculated for different conditions. The results are shown in Fig 11a.

Fig 11 Desulphurization of liquid steel

This figure can be used for estimating the specific quantity of ladle slag ‘r’, needed to give the desired degree of sulphur removal, as shown by the following example. In the example, a shop is considered where hard stirring, equivalent to a stirring power density of 100 W/ton, is practiced for desulphurization with a ladle slag with a composition equivalent to ‘Ls’ = 500. From Fig 5b the value ‘ks’ of around 0.13 per minute is achieved for ‘e’ = 100 W/ton. Assuming that the total time to be reserved for desulphurization is 15 minutes for achieving 80 % desulphurization, i.e., from 0.01 % to 0.002 % S. From Fig 11a, it is found that ‘R’ = 0.8 can be achieved in the given time, provided ‘rLs’ = 10 or ‘r’ = 0.02, equivalent with a specific quantity of ladle slag of 20 kg/ton of steel. For a stirring power density of around 50 W/ton the value of ‘ks’ is estimated to be around 0.03 per minute (Fig 5b), and a treatment time of 65 minute is needed for achieving 80 % desulphurization. This example shows the importance of hard stirring for effective desulphurization to low sulphur levels.

The rate of desulfurization depends on (i) stirring rate, (ii) slag chemistry which affects ‘Ls’, and (iii) aluminum content which also affects ‘Ls’. Application of equation 8 indicates that desulphurization equilibrium in a well-stirred ladle using 0.85 N cum to 10 N cum argon occurs in around 10 minutes to 15 minutes. For achieving very low sulphur contents, the injection of fluxes into the ladle is frequently practiced. A study has described results achieved by injecting 70 % CaO-30 % CaF2 powder mixtures into powder injection results in a desulphurization rate which is, on average, around 15 % faster than desulphurization with a top slag only, combined with gas stirring. This implies that the contribution of the so-called transitory reaction with the powder mixture as it ascends the bath is minor compared with the reaction with the top slag.

Desulphurization of steel in the ladle is accompanied by a decrease in the temperature of the steel bath. Today majority of the steel melting shops are equipped with facilities for reheating the steel, either by electric arc or by injection of oxygen and aluminum. If no facilities for reheating are available, an exothermic mixture consisting of 58 % burnt lime, 30 % hematite and 12 % aluminum powder can be used for the desulphurization.

Dephosphorization – Normally, it is preferable to remove phosphorus from steel under the oxidizing conditions prevalent in the BOF or in EAFs with oxygen injection. In older EAF shops equipped with furnaces with inadequate or no oxygen injection capability, the need for steel dephosphorization in the ladle can arise. Also, ladle dephosphorization can be necessary in BOF shops in which hot metal with a high phosphorus content is charged and where there is no capability of dephosphorizing the hot metal prior to charging to the BOF. Steel made in induction furnace can also need dephosphorization.

Removal of phosphorus from the steel in the ladle is achieved by treating the steel with lime-based oxidizing slags containing iron oxide. For a steel with a given dissolved oxygen content, the ratio Lp = (%P) / [%P] is a function only of the slag composition and the temperature. This is shown in Fig 10b, where lines of equal ‘Lp’ values are projected on the CaO-FeO-(SiO2 + P2O5) phase diagram, the dissolved oxygen contents being controlled by the Fe-FeO equilibrium.

Dephosphorization in the ladle during tapping of the BOF was the subject of a study, In the study, varying quantities of a mixture of 50 % CaO, 30 % iron oxide (FeOx) and 20 % CaF2 have been used. Around 30 % to 40 % of the mixture has been placed on the bottom of the steel teeming ladle while the remainder has added during tapping. The phosphorus content of the steel tapped from the BOF varied between around 0.01 % to 0.035%. The results of these plant trials are given in Fig 12a. It is seen from the figure that around 75 % of the phosphorus is removed when 12 kg/ton of the above-mentioned mixture has been used.

Fig 12 Degree of dephosphorization and inclusion modification

In the practice, the high-phosphorus slag is removed by reladling followed by reheating in a ladle furnace. In another practice of steel dephosphorization in the ladle at tapping, the steel is tapped open while, depending on the anticipated quantity of the BOF slag carry-over, varying quantities of lime, ore, and sometimes fluorspar are added during tapping for producing a lime-saturated dephosphorizing slag high in iron oxide. The results of these practices indicate 60 % to 70 % phosphorous removal.

Because of the oxidizing conditions prevailing during dephosphorization, manganese and chromium losses are to be expected. The relationship between the loss in chromium or manganese ‘Ex’, and the degree of dephosphorization ‘EP’ is Ex = 100 Ep / (100 Kx – Ep (Kx – 1)’ (equation 24) where x is chromium or manganese and ‘Ex’ = {[%x]i – [%x]f } / [%x]i, and ‘Ep’ = {[%P]i – [%P]f } / [%P]i. From the plant data, the values for Kx:KCr = 6.6 and KMn = 2.2. For example, for 50 % phosphorus removal (Ep = 50) the chromium loss is 13 % (ECr = 13), while the manganese loss is 31 % (EMn = 31), as found from equation 24.

Alloy additions – Metals and alloys can be added to liquid steel at different stages in the steelmaking process, e.g., inside the furnace, during furnace tapping, in the ladle furnace, and during vacuum treatment etc. The timing of the additions depends on the process route, the logistics and certain characteristics of the addition in question such as its melting point, volatility, and its susceptibility to oxidation. As an example, nickel can be added to the EAF at any time as nickel oxide, which is easily reduced. In the oxygen steelmaking process route, alloying additions such as ferro-silicon and ferro-manganese are made during furnace tapping, while the other alloys are added in subsequent stages of secondary steelmaking.

A study has defined two broad categories of ferro-alloys namely (i) class I ferro-alloys with melting points below the temperature of liquid steel, and (ii) class II ferro-alloys with melting points higher than the liquid steel temperature. When a ferro-alloy is added to liquid steel, a solidified shell of steel forms around the alloy particle as a result of the local chilling effect. As time progresses, the shell melts while the ferro-alloy inside the shell is heated to its melting point. The complete dissolution is governed by convective heat transfer processes in the bath as well as the size of the ferro-alloy added.

The class II ferro-alloys have melting points higher than the temperature of the liquid steel. These alloys dissolve at a slower rate than the class I alloys, their dissolution rate being controlled by mass transfer in the liquid steel, even in agitated baths. It is, hence, important to ensure that their size is within 3 mm to 10 mm in order to get good mixing, fast dissolution, and high recovery rates.

Compacted powder mixtures of ferro-alloys such as ferro-vanadium, ferro-tungsten and ferro-molybdenum dissolve faster than solid pieces of similar size. Auto-exothermic alloys, which generate heat upon melting, can also be used for faster melting and dissolution as well as improved recovery.

Several methods of alloy addition are practiced. Examples are throwing of filled bags, adding with a shovel or through mechanized chutes, wire feeding, powder injection, and bullet shooting etc. A special process for making alloy additions is the so-called CAS process (Composition Adjustment by Sealed argon bubbling). In this process a refractory-lined snorkel is partially immersed in the steel bath in such a manner that it envelops the ascending gas plume created by the injection of argon through the porous plug in the ladle bottom. Alloy additions are made onto the liquid steel surface within the area covered by the snorkel. The plume eye within the snorkel is filled with argon, hence it has a low oxygen partial pressure, preventing oxidation of the alloy addition. Melting and distribution rates are high as a result of the agitation brought about by the ascending gas bubbles.

A study has made water model studies for investigating the sub-surface motion of both buoyant and sinking additions in the CAS process. The study showed that buoyant additions such as aluminum and ferro-silicon dissolve more readily into the steel bath rather than react partially with the slag as in conventional addition methods, hence giving improved recovery. The study has further recommended that ferro-manganese and ferro-niobium be crushed to an average size of around 5 mm to achieve better control.

Wire feeding of alloys by means of the cored-wire techniques, developed mainly for the addition of calcium to steel, is practiced for adding elements which are less dense than steel or have a limited solubility, high vapour pressure, and high affinity for oxygen. Wire feeding is also used in cases where the element to be added is toxic or when very small additions are needed. The cored-wire technique permits the quantity of alloy or elements being fed into the steel to be adjusted with high precision and to trim the composition of the steel within narrow limits. As an example, ferro-boron or tellurium additions can be made in precise and minute quantities by wire feeding. Excessive additions of these elements can cause hot-shortness.

It is also possible to feed aluminum wire with the same wire feeding equipment used for cored-wire. Advantages of aluminum wire additions include higher recovery, better control of aluminum content, and improved cleanliness. A study has given examples of improved control of the steel aluminum content by wire feeding.

The dissolution characteristics of cored-wire additions of ferro-molybdenum, ferro-niobium and ferro-chromium modified with minor quantities of silicon, so-called micro-exothermic alloys, has been studied for achieving improved dissolution into liquid steels. The exothermicity shown by these modified alloys is based on the formation of an inter-metallic compound (a silicide) which is accompanied by the release of heat. The enthalpy released is sufficient to melt the compound, hence allowing rapid dissolution of the ferro-alloy into the liquid steel. The improved dissolution rate of these modified ferro-alloys makes them well-suited for tundish additions where the relatively short residence time needs a rapid dissolution of additions.

Calcium treatment and inclusion modification – The addition of calcium to steel goes back a long time, Watts being the first to add Ca-Si to a steel melt. The wide-spread practice of calcium additions to steel melts has not started until the 1960s with the development of improved addition methods and composite calcium-bearing alloys. Today calcium treatment of steel is a normal practice, with particular emphasis on the modification of alumina inclusions in aluminum-killed steels for preventing nozzle clogging during continuous casting operations. As a result of the treatment with calcium, the alumina and silica inclusions are converted to liquid calcium aluminates or calcium silicates. These liquid inclusions are globular in shape because of surface tension effects. This change in inclusion composition and shape is normally known as inclusion morphology control or modification. The effect of calcium treatment on inclusion morphology is shown schematically in Fig 12b.

It is seen from Fig 12b that few or no sulphide stringers are expected to be present after rolling steel which has been successfully treated with calcium. This phenomenon is known as sulphide shape control by calcium treatment. Examples of other metallurgical advantages brought about by the modification of oxide and sulphide inclusions by calcium treatment of steel are (i) improvement of castability in continuous casting operations through minimization or prevention of nozzle clogging, (ii) decreasing inclusion related surface defects in billet, bloom and slab castings, (iii) improved machinability of the final product at high cutting speeds, and (iv) minimization of the susceptibility of high-strength low-alloy (HSLA) line pipe steels to hydrogen-induced cracking (HIC) in sour gas or oil environments.

Addition of calcium to steel melts – The boiling point of calcium is 1491 deg C, accordingly, calcium is a vapour at steelmaking temperatures. Hence, when adding calcium to liquid steel, special measures are to be taken for ensuring its proper recovery in the steel bath. Recently-developed processes for adding calcium to a liquid steel bath are all based on the principle of introducing the calcium or calcium alloy into the bath at the highest possible depth so as to make use of the increased pressure from the ferro-static head and hence prevent the calcium from evaporating.

Ototani has given details regarding the TN (Thyssen Niederrhein) process for injecting calcium with argon as a carrier gas as well as the SCAT process, also known as the bullet shooting method. Today the majority of steel producers add calcium by wire feeding. The principle is similar to wire feeding of ferro-alloys and aluminum. When wire feeding calcium in the conventional manner, there is a possibility that the wire does not travel in a straight downward line after entering the bath, hence causing the calcium to be released at a shallow bath depth and decreasing the calcium recovery in the steel. For preventing this, the so-called wire lance (WL) method for adding calcium has been developed. A schematic diagram is shown in Fig 13a.

Fig 13 Wire lance method and comparison of calcium recoveries

 The wire lance method ensures that the calcium wire travels in a straight downward line after entering the bath. It is claimed that the dispersion of calcium throughout the bath is improved by the argon which is injected simultaneously. The calcium recovery or yield observed for additions with the wire lance process is compared in Fig 13b with that observed for conventional wire feeding. The better recovery of calcium achieved with the wire lance method is especially pronounced for calcium addition rates less than around 0.2 kg/ton.

Calcium usage efficiency – The material balance for calcium consumption is expressed as ‘Wi = Wb + Wo + Wo’ + Ws + Wv’ (equation 25), where ‘Wi’ is the quantity of calcium injected, ‘Wb’ is the quantity of calcium dissolved in the bath, ‘Wo’ is the quantity of calcium present in aluminates and sulphides,  ‘ Wo’ ‘ is the quantity of calcium reacted with alumina and subsequently floated out, ‘Ws’ is the quantity of calcium reacted with the slag, and ‘Wv’ is the quantity of calcium escaped through the vapour phase and subsequently burnt at the bath surface.

It is normally accepted that ‘Wb’ is very much less than ‘Wo’, hence giving for the efficiency of calcium usage ‘E(Ca)u = [(Wo – Wo’)/Wi] x 100 %’ (equation 26), while the efficiency of calcium retention in the steel is given by ‘E(Ca)r = (Wo/Wi x 100 %’ (equation 27).

Experience from several plant trials has shown that the calcium retention efficiency decreases with increasing quantity of calcium injected. The quantity of calcium to be injected is needed to be adjusted as per the degree of cleanliness of the steel or its total oxygen content. Clearly, injecting more calcium than can react with the available inclusions leads to a low calcium retention efficiency. Also, it is to be expected that the calcium retention efficiency in the continuous casting mould or in the teeming ingot is lower than the retention efficiency in the ladle because of the flotation of calcium-containing inclusions out of the bath in the time interval prior to casting or teeming.

Reactions of calcium in steel and inclusion modification – Regardless of the form in which calcium is added to liquid steel, e.g., as Ca-Si, Ca-Al, or as pure calcium admixed with nickel or iron powder, the subsequent reactions taking place in the bath are the same. The series of reactions expected to occur to varying extents in aluminum-killed steels containing alumina inclusions and sulphur are (i) Ca(l) = Ca (g)’ (equation 28), Ca(g) = [Ca]’ (equation 29), ‘[Ca] + [O] = CaO’ (equation 30), ‘[Ca] + [S] = CaS’ (equation 31), ‘[Ca] = (x + 1/3) Al2O3 = CaO.xAl2O3 + 2/3[Al]’ (equation 32), and  ‘(CaO) = 2/3[Al] + S = (CaS) + 1/3(Al2O3)’ (equation 33).

The symbols within brackets refer to species dissolved in the steel, and those within parentheses are dissolved in the aluminate phase. Observations in a number of studies have indicated that the extent to which reaction at equation 30 occurs is negligible. For steels with sufficiently low sulphur contents, reaction at equation 32 take place first, followed by the reaction at equation 33.

The critical question is for which sulphur content reaction at equation 31 predominates such that, for a given quantity of calcium added, there is insufficient calcium available to modify the alumina inclusions as per the reaction at equation 32. Adequate modification of the solid alumina inclusions into liquid calcium aluminates is necessary for preventing nozzle clogging during continuous casting operations.

By combining the data for the activities of the oxides in aluminate melts with the equilibrium constant for reaction at equation 33 as given by reaction at equation 32, the critical sulphur content for the formation of liquid calcium aluminates can be evaluated. The results are shown in Fig 14a and Fig 14b, where the critical sulphur and aluminum contents for the formation of liquid alumina-saturated calcium aluminate and liquid 12CaO.7Al2O3 are shown. Whenever the aluminum and sulphur contents in the steel fall below a curve for a given temperature, the formation of a liquid calcium aluminate is favoured.

Fig 14 Critical steel sulphur and aluminum contents

The underlying assumption made in the calculations of the Fig 14a and Fig 14b is that the activity of calcium sulphide (aCaS) equals unity. However, the presence of manganese in the majority of the steels causes the sulphur to precipitate as calcium manganese sulphides, Ca(Mn)S, in which the activity of calcium sulphide in less than unity, hence decreasing the critical sulphur content for the formation of liquid calcium aluminate (CA).

In a study, it has been shown that the presence of up to 2 % manganese in the steel has only a small effect on the critical sulphur content. In another study, the modification of oxide inclusions by calcium in a number of samples got from laboratory melts as well as from steelmaking operations have been studied. The results from this study are given  in Fig 15a, from which it can be seen that the agreement with the theoretical predictions, indicated by the curves labelled CA and C12A7, is normally good. It is normally difficult to assess whether the injection of calcium into the steel has resulted in the desired degree of inclusion modification.

Fig 15 Composition of inclusions and calculated oxygen activity

The study has pointed out that, in theory, the degree of inclusion modification can be determined by measuring the activity of oxygen in the steel by means of an oxygen sensor. The activity of alumina in calcium aluminate inclusions is normally less than unity. Hence, the activity of oxygen in a steel in equilibrium with a calcium aluminate is less than that in a steel in equilibrium with alumina. As the inclusions are modified from alumina to calcium-rich calcium aluminate, the activity of oxygen in the steel decreases, provided that the aluminum content of the steel is essentially constant. This is shown schematically in Fig 15b, from which it can be seen that the decrease in oxygen activity (in ppm) is considerable whenever complete modification to liquid calcium aluminate has occurred.

Hence, an oxygen sensor measurement before and after calcium treatment, in principle, indicates how effective the treatment is in terms of inclusion modification. In several samples taken from the steelmaking operations, the oxide inclusions contain varying quantities of magnesia. A study has estimated the effect of the presence of magnesia in the oxide inclusions on the critical sulphur content for the formation of liquid calcium magnesium aluminate for a given aluminum content of the steel. It has been found that for inclusions containing less than 10 % magnesia, the critical sulphur content is somewhat higher than that for inclusions not containing magnesia.

Sulphide shape control – In steels not treated with calcium, the sulphur precipitates as finely dispersed manganese sulphide particles in the inter-dendritic liquid which freezes last. The manganese sulphides delineate the prior austenitic grain boundaries in the as-cast structure. During hot rolling, the manganese sulphide (MnS) particles are deformed, resulting in stringers in the rolled product. These stringers make the final product susceptible to, for example, hydrogen-induced cracking in sour gas or oil environments.

In calcium-treated low-sulphur steels the grain boundary precipitation of MnS during solidification is suppressed as a result of the precipitation of sulphur as a Ca(Mn)S complex on the calcium aluminate inclusions, as indicated by the following reaction ‘(CaO) + 2[S] + [Mn] + 2/3[Al] = (CaS.MnS) + 1/3(Al2O3)’ (equation 34).

The extent of sulphide shape control which can be achieved during solidification of calcium-treated steel depends on the total oxygen, sulphur, and calcium contents of the steel. This is described by a model based on the reactions occurring in the impurity-enriched inter-dendritic liquid during solidification. On the basis of this model, the criteria can be derived for the tundish compositions of aluminum-killed steels for giving adequate sulphide shape control in the final product.

In steels with a total oxygen content of 10 ppm or less and relatively high sulphur contents, i.e., higher than 100 ppm, sulphide shape control by means of calcium treatment is clearly not feasible. For minimizing the occurrence of sulphide stringers in such steels, the addition of tellurium (Te) or sometimes selenium (Se) has been found to be beneficial. Because of the strong effect of both these elements on the interfacial tension between sulphides and steel, the tendency of sulphide stringer formation during rolling is decreased. The result is that, after rolling, the sulphides are ellipsoidal in shape with a length-to-width ratio which depends on the Te / S-ratio in the steel. Tellurium is normally added to liquid steel either by powder injection or by wire feeding.

Vacuum degassing – Vacuum degassing of steel has an even longer history than the treatment of steel with calcium. Initially, vacuum degassing was used mainly for hydrogen removal. However, during the late 1970s or so, there has been an increased use of vacuum degassing for the production of ultra-lowcarbon (ULC) steels with carbon contents of 30 ppm or less. Also, a relatively new family of steel grades, the so-called interstitial-free (IF) steels with carbon and nitrogen contents of 30 ppm or less, has appeared on the scene. For achieving these low carbon and nitrogen contents, a treatment under vacuum is mandatory. Presently, almost every high-quality steel producer has installed a vacuum treatment facility.

The two principal types of degassing units are (i) recirculating systems such as RH, RH-OB, RH-KTB and DH, and (ii) non-recirculating systems such as ladle or tank degassing units, including VAD (vacuum arc degassing) and VOD (vacuum oxygen decarburization), and stream degassing units. In both recirculating and non-recirculating systems, argon is used as the lifting or stirring gas. In recirculating systems, the argon is used as the so-called lifting gas for lowering the apparent density of the liquid steel to be lifted up from the ladle into the vacuum vessel. In non-recirculating systems, argon is used as the stirring gas for promoting the removal of hydrogen and / or nitrogen and for homogenizing the bath.

The decision of which degassing system, recirculating or non-recirculating, for installing in a given shop is largely determined by the product mix to be produced. If a relatively large number of heats has to be decarburized to very low levels to produce ULC or IF steels, a recirculating system such as the RH or one of its modifications is normally preferred. For example, a carbon content of 25 ppm can easily be achieved in an RH or RH-OB (KTB) degassing unit, whereas in a tank degassing unit, such as a VOD, such low carbon contents cannot be achieved within a practical time span. There is not much difference between recirculating and non-recirculating systems in terms of the effectiveness with which hydrogen or nitrogen can be removed. Hence, if the primary function of the degassing unit is to remove hydrogen and sometimes nitrogen, the choice of system is to be determined mainly by the desired match between the steel melting vessel (BOF or EAF) and the CCM as well as by considerations with regard to capital and operating costs.

Vacuum carbon deoxidation – One of the purposes of treating steel in an RH or RH-OB (KTB) degassing unit is to lower the dissolved oxygen content of the steel by means of carbon deoxidation before adding aluminum to kill the steel completely. With such a carbon deoxidation practice there are considerable cost savings as a result of the decreased usage of aluminum. Vacuum carbon deoxidation is described by the reaction ‘[C] + [O] → CO (g)’ (equation 35), where the carbon and oxygen are dissolved in the steel bath. Fig 16a shows schematically the carbon–oxygen relationship during the vacuum decarburization treatment.

Fig 16 Different relationships in vacuum degassing

In the RH process, decarburization proceeds nearly to the stoichiometrically related decrease in carbon and oxygen contents, ‘Delta[O] = (16/12) Delta[C]’. This is also called the self-decarburization process. Since in the RH-OB (KTB) process oxygen is supplied from an outside source, decarburization initially takes place without a simultaneous decrease in the steel oxygen content, the so-called forced decarburization. In the later stages decarburization follows the path of self-decarburization. One of the advantages of an RH-OB (KTB) over a conventional RH process is that the steel can be tapped at a higher carbon content, hence decreasing the BOF processing time and increasing the iron yield (lower slag FeO).

Plant data got for the carbon and oxygen contents of the steel before and after RH treatment are shown in Fig 16b. Similar results are got with the tank degassing unit. Although the pressure in the vacuum vessel is around 0.1 kilopascal (kPa), the final carbon and oxygen contents correspond to CO pressures varying from 6 kPa. to 8 kPa. (Fig 16b). After around 20 minute of treatment time, the final oxygen content of the steel is always high when the initial content is high (low initial carbon content) in both the tank degassing unit and the RH unit (Fig 16c).

It can be seen from Fig 16c that the decrease in oxygen content of the steel as a result of the vacuum decarburization treatment is less than that expected from the stoichiometry of reaction at equation 35. This is since there is oxygen transfer from the ladle slag to the steel during vacuum decarburization. Another source of oxygen is the iron oxide-rich skull which builds up on the inside of the vacuum vessel as a result of previous operations. Hence, some decarburization through the reaction which also takes place is ‘(FeO) + [C] = Fe + CO (g)’ (equation 36).

From the stoichiometry of the reactions, the material balance which gives the relation for the quantity of oxygen transferred to the steel from the ladle slag and the oxidized skull inside the vacuum vessel is ‘Delta O(slag) = (16/12) Delta [C] – Delta [O] (equation 37), where Delta [C] = [%C]i – [%C]f and Delta [O] = [ppm O]i – [ppm O]f. The values of Delta O(slag) derived from the plant data using equation 37 are shown in Fig 17a. It is seen that during decarburization the quantity of oxygen transferred from slag to steel is higher the higher the initial carbon content. For initial carbon contents of 200 ppm or less there is no more oxygen pick-up from the slag during vacuum decarburization, and the carbon content decreases as per stoichiometry of reaction at equation 35.

Fig 17 Vacuum degassing different relationships

Rate of decarburization – The rate of decarburization is expressed by the relationship ‘Ln{[%C]f/[%C]i} = -kc x t’ (equation 38), where ‘[%C]I’ and ‘[%C]f’ are the carbon contents before and after decarburization, respectively, and ‘kc’ is the rate constant for decarburization. For RH degassers the rate constant is given by the relationship ‘kc = [Q/9Vb x d)] x [q/{(Q/d +q)}], per minute (equation 39), where ‘Q’ is the circulation rate of liquid steel in kg/minute, ‘Vb’ is the volume of the steel bath in the ladle in cubic metre, ‘d’  is the density of liquid steel in kg/cum, ‘q ‘ is volumetric mass transfer coefficient of decarburization in N cum/minute.

As per the circulation rate of the liquid steel in the RH vessel is given by ‘Q (ton per minute) = 11.43 x (V) to the power 1/3 x (D) to the power 4/3 x [ln(P1/P0)] to the power 1/3’ (equation 40), where ‘V’ is the flow-rate of argon injected into the up-leg snorkel in N litre/minute, ‘D’ is the inside diameter of the up-leg snorkel in metre, ‘P1’ and ‘P0’ are the pressure at the argon injection point and at the bath surface respectively.

The volumetric mass transfer coefficient of decarburization ‘q’ is proportional to the cross-sectional area ‘Av’ of the vessel, which is equivalent to the surface area of the bath. From plant observations on 240-ton to 300-ton RH vessels, the approximate empirical relationship which is found is ‘q = 0.26 (Q) to the power 0.64 x Av x [%C]’ (equation 41) valid for [%C] in the range of 0.0025 to 0.01.

It is to be noted that the actual rate is very complex. The reaction occurs at different sites, including the argon bubble surface, refractory surfaces, metal free surfaces, and homogeneously in the melt. Hence, equation 41 is to be used only for similar conditions for which it has been developed.

As per the above equations, the rate of decarburization increases with snorkel diameter and vessel diameter, which has been confirmed by actual plant data.

In the Fig 17b, the decarburization rate constant ‘kc’ is shown as a function of the specific flow rate (N litre/min per ton) of the stirring gas for recirculating systems such as RH and DH and for non-recirculating systems such as VAD. Because of the lower specific flow-rates for the stirring gas used in non-recirculating systems, the time needed for removing 50 % of the carbon is around 7 minutes, whereas in the RH, this time can be as short as 3 minutes to 4 minutes. (Fig 17b).

Several methods for improving the decarburization rate in the RH unit have been reported. In one of the studies, the decarburization rate has been increased by injecting argon through nozzles installed in the hearth of the RH vessel. By injecting argon at a rate of 400 N litre/minute to 500 N litre per minute, the carbon content in a 100-ton heat of steel has been decreased from 200 ppm to around 10 ppm in 20 minutes, corresponding to ‘kc’ less than 0.15 per minute as found from equation 38. This is around 50 % higher than ‘kc’ for a conventional RH unit (Fig 17b).

In one study, hydrogen and argon are co-injected into the up-leg of the RH unit for achieving 6 ppm hydrogen in the steel. The vessel is then evacuated while the injection of hydrogen continued. The evolution of hydrogen gas bubbles within the bath resulted in a final carbon content of 5 ppm to 10 ppm.

One of the studies has reported that top blowing of powders onto the surface of the bath in the RH unit, the so-called RH-PB (powder blowing) process, is effective for achieving ultra-low carbon, nitrogen and sulphur contents. For example, by blowing 20 kg/minute to 60 kg/minute of iron ore powder (minus 100 mesh) through a top lance positioned 2 m to 3 m above the bath surface, the final carbon content achieved has been less than 5 ppm. Whenever the initial carbon content of the steel is relatively high, the decarburization rate can be limited by the supply of oxygen. For remedying this, the RH-OB (oxygen blowing) and RH-KTB (Kawasaki top blowing) processes have been developed. In the RH-OB process the oxygen is supplied through tuyeres installed in the sidewalls in the lower part of the RH vessel. In the RH-KTB process the oxygen is supplied through a lance situated in the RH vacuum vessel. In these configurations of the RH process, skull formation inside the vessel has been minimized by post-combustion of the CO by the injected oxygen.

In one study, a reaction model has been developed for decarburization in the RH and RH-KTB processes which is based on a mixed control mechanism involving the mass transfer of carbon and oxygen in the liquid steel present in the vacuum vessel, as well as on the transport of carbon and oxygen by the recirculating steel. The model satisfactorily takes into account the effect of the concentration of oxygen on the decarburization rate in the conventional RH as well as in the RH-KTB process.

Hydrogen removal – The rate of hydrogen removal during degassing is controlled by mass transfer in the liquid steel, for which the rate equation is given by ‘ln([H]f – [H]e]/[H]i – [H]e) = -kh x t’ (equation 42), where ‘[H]f’ is the hydrogen content after degassing, ‘[H]I’ is the initial hydrogen content, ‘[H]e’ is the equilibrium hydrogen content as determined by the pressure in the system, ‘kh’ is the overall rate constant for hydrogen removal. In the majority of the modern degassers the achievable pressure is below 1 kPa (around 10 torr) and hence, ‘[H]e’ can be neglected with respect to ‘[H]I’ and ‘[H]f’. This gives the simplified rate equation ln([H]f/[H]i) = -kh x t’ equation (43).

In one of the studies, a mathematical model has been developed for hydrogen removal in a 185-ton tank degasser. The agreement between the hydrogen content measured after degassing and that calculated from the model is excellent, as shown in Fig 17c. The model is based on fundamental principles. The most critical parameter is the bubble size, which is extremely difficult to predict. Hence, the model is to be used only to make comparisons for similar operating conditions.

Using the model, in the study, hydrogen removal rates have been calculated for initial hydrogen contents varying between 3 ppm and 7 ppm and argon flow-rates of 0.9 N cum/minute and 1.8 N cum/minute. The results are given in Fig 18a, from which it can be seen that it takes 2 minutes to 3 minutes longer to achieve a given degree of hydrogen removal for steels initially containing 3 ppm hydrogen than for steels initially containing 7 ppm hydrogen. Also, it is noted that doubling the flow-rate of the argon stirring gas results in a marginally shorter treatment time, e.g., around 2 minute shorter for achieving 70 % hydrogen removal. Hydrogen removal rates for tank pressures ranging from 0.1 kPa to 100 kPa are shown as a function of the total volume of argon stirring gas flowing at 1.8 N cum/min in Fig 18b.

Fig 18 Removal of hydrogen and nitrogen from steel

It is seen that the final hydrogen content of the steel is essentially unaffected by the tank pressure for pressures up to 1 kPa. From the data depicted in Fig 18a. it is found that the overall rate constant for hydrogen removal ‘kh’, increases from around 0.09 per minute to 0.16 per minute, when the argon flow-rate increases from 0.9 N cum/minute to 1.8 N cum/minute. The higher ‘kh’ value is comparable to the value of 0.13 per minute observed for an RH vessel with a 600 mm diameter snorkel and a steel circulation rate of around 140 ton/minute. These values indicate that the efficiencies of recirculating and non-recirculating systems for the removal of hydrogen are similar.

Hydris probe – In steel melting shops where hydrogen-sensitive steel grades such as, for example, large bars, are produced, it is important to know the hydrogen content of the steel bath before it is delivered to the CCM or the teeming platform. Under such conditions, an in-site determination of the hydrogen content of the bath in the ladle is desirable.

Around late 1980s, a study described an immersion system, the Hydris probe, for the rapid in-situ determination of hydrogen in a bath of liquid steel. The principle of the measurement is based on equilibrating a known volume of argon, being passed through the liquid steel, with the hydrogen dissolved in the steel. Hence, the argon-hydrogen gas mixture leaving the steel after equilibration has a partial pressure of hydrogen which, through Sieverts’ law ‘K = [%X]/(Px2) to the power 1/2 (equation 44), can be related to the hydrogen content of the steel. The Hydris probe has been tested extensively, and it has been found that it gives reliable readings of the hydrogen content of the liquid steel with an uncertainty of around +/- 5 %.

The Hydris probe has been found to be rugged and reliable. The cost associated with the use of the probe has to be weighed against cost savings made possible by its use, for example, there is a considerable decrease in degassing time for achieving the needed hydrogen content in the steel.

Nitrogen removal – Some nitrogen removal from liquid steel during vacuum degassing is possible, provided the steel is fully killed and has a low sulphur content. A study has developed a rate equation for the nitrogen removal in a 185-ton tank degassing unit which has been based on a mixed-control model, i.e., liquid-phase mass transfer of nitrogen to the argon bubbles coupled with chemical reaction control at the liquid-gas bubble interface. As shown in Fig 18c, the nitrogen contents after degassing calculated from the model and indicated by the solid lines are in good agreement with the measured nitrogen contents indicated by the different symbols. Equally good agreement between calculated and measured values has found for steels containing between 20 ppm and 200 ppm sulphur.

The rate of nitrogen removal for different initial nitrogen contents as calculated from the model is shown in Fig 19a for a killed steel containing 2 ppm dissolved oxygen, 10 ppm sulphur and a tank pressure of 0.1 kPa. It can be seen that under these conditions around 50 % of the nitrogen can be removed in roughly 15 minutes, provided the initial nitrogen content is 50 ppm or higher. The effect of the steel sulphur content on the rate of nitrogen removal as calculated from the model is shown in Fig 19b. As with the hydrogen case, the model is to be used carefully because of the uncertainty in bubble size.

Fig 19 Removal of nitrogen from steel

A simplified form of the rate equation for nitrogen removal is ‘(1/ppm N – 1/ ppm No) = kN x (1- theta) x t’ (equation 45) where kN is the apparent rate constant for de-nitrogenization for the limiting case in which both the oxygen and sulphur contents tend to zero.

From the nitrogen removal rates shown in Fig 19b, together with equation 46 for (1 – theta), the apparent rate constant is estimated to be kN less than 0.0013 per ppm N min for 185-ton heats in a tank degassing unit at 0.1 kPa pressure. Equation 46 is ‘1/%N – 1/%No = 100(A /d x V) x (phi)r x (1- theta) x t’.

The overall rate constant for de-nitrogenization in an RH degassing unit is shown as a function of the sulphur content of the steel in Fig 19c. It is to be noted that the value of kN for sulphur contents approaching zero is very similar to the aforementioned value for a tank degasser. Hence, from these values it is concluded that recirculating and non-recirculating degassing units are equally effective in removing nitrogen as well as hydrogen. Recently, an on-line method for the determination of the nitrogen content of liquid steel has been developed. The system, called Nitris, is based on the same measuring principle as the Hydris probe.

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