Effect of Hot Metal on Decarburization in the EAF and Dissolved Sulfur, Phosphorous, and Nitrogen Content in the Steel

Abstract

The effect of hot metal additions on the decarburization and dissolved sulfur, phosphorous, and nitrogen content in the steel of a DC EAF was investigated. The addition of hot metal of maximum 36% provided heat into the EAF allowing faster melting and reduced power on times of the furnace. The increased melting rate lowered the FeO in the slag and seems to allow faster kinetics for CaO dissolution into the slag. Hot metal utilization into the slag resulted in higher phosphorous distribution ratios compared to hot metal free heats achieving comparable phosphorous levels at the ladle metallurgical furnace even though the initial input phosphorous was much higher than hot metal free heats. The effect of hot metal on the desulfurization was not pronounced and no apparent difference could be ascertained compared with the 100% scrap charge. With the addition of carbon saturated iron units in the EAF, the evolution of CO and foaming was promoted, which inhibited the infiltration of nitrogen and lowered the overall partial pressure of nitrogen resulting in lower nitrogen levels in the steel. In addition, the dilution effect of the tramp elements Cu and Sn with hot metal ensured the critical defect index to be less than 10% resulting in a 50% reduction of the quality defect index.

1. Introduction

With the on-going demands for energy efficiency, environment, low cost, and flexibility for steelmaking operations, while allowing the production of high value-added steels, there has been significant development of the mini-mill process route.¹⁾ This operation unlike the integrated steel mills typically utilizes an electric arc furnace (EAF) with secondary steelmaking facilities and a thin slab caster connected to a continuous hot strip mill, where the iron feed is from scrap. However, due to increased consumer demands for steel-cleanliness and low-cost, there has been increased pressure on the EAF steel producers to provide effective means for satisfying customer needs, as existing feed materials of scrap continues to degrade in quality. In addition, thin slab casters, which can be highly productive when optimized, require better internal quality of the steels compared to conventional slab casters due to its high casting speed and high heat extraction directly being charged into the continuous hot strip mill. In order to meet these critical demands, the EAF steelmakers have attempted to utilize iron sources from various virgin iron sources including direct reduced iron, pig iron, and also hot metal. In particular, as electricity costs begin to rise, there has been increased emphasis in possible utilization of hot metal for facilities having excess hot metal especially during downturns in the steelshop, which has significant sensible heat for reduced power consumption in the EAF.

Zinurov et al.²⁾ described the effects of hot metal utilization in 180 ton/ch arc furnace, where substitution of 1% of scrap by hot metal at 1553 K (1280°C) lead to power savings of 3.5–4.5 kWh/t. However, hot metal introduction into the furnace was an issue as problems arose from the radiant heat loss and graphite liberation from the hot metal, which deposited around the plant equipments increasing maintenance and particulates in the plant. Duan et al.³⁾ showed 35 to 40% hot metal additions into a 50 ton/ch EAF after pre-treating the carbon saturated iron source reduced tap-to-tap times of the furnace by 5 to 10 minutes and decreased the electricity consumption by 35–50 kWh/t while maintaining the chemical compositional requirements and quality of the steel. In addition, although desulfurization potential was not significantly affected by hot metal additions, the de-phosphorization was enhanced with hot metal. Others^4,5,6) have also indicated significant costs savings for plants that have hot metal flexibility to add into the EAF as a source for both high quality iron and energy. Gottardi et al.⁷⁾ described oxygen injection practices to increase hot metal usage in the EAF without significant increases in the tap-to-tap times, which can affect the productivity of the furnace. However, few studies have described in detail the chemical composition changes with hot metal additions as carbon saturated iron units are added into the EAF to the knowledge of the present authors.

In this study, the effect of hot metal additions to the EAF in terms of chemical composition changes has been studied. Effect of hot metal additions up to approximately 36% in the 130 ton/ch DC-EAF on the decarburization rate, phosphorous, sulfur, and nitrogen content has been specifically examined. In addition, the inclusion chemistry with hot metal additions has also been discussed.

2. Experimental Method and Procedure

2.1. Hot Metal Supplement Procedure into the EAF

The schematic of the charging sequence of raw materials input into the EAF is shown in Fig. 1. EAF cover is opened and the scrap bucket containing 90–105 tons is placed above the EAF. The scrap is dropped into the furnace containing the hot heel of the previous charge with the light scrap initially loaded into the furnace to protect the refractory. The heavy scrap follows the light scrap and the small scrap pieces including the shredded scrap are charged into the furnace at the end to protect the boring electrode during the initial power on time. With the start of arcing the electrode, additional fluxes of 2 tons of quick lime and 0.5 tons of dolomite is added at the top of the scrap and the electrode bores into the scrap as it descends into the furnace. At the end of the boring cycle, the electrode is raised and 50–65 tons of hot metal is added into the crater created during the boring cycle of arcing.

Fig. 1.

Schematic of the scrap and hot metal charging into the 130 ton/ch DC-EAF.

The hot metal is added slowly for about 3–4 min. into the furnace through a charging ladle after the EAF cover is removed. Excessively fast pouring can cause splashing and unnecessary vaporization resulting in metal losses. After the hot metal pouring is completed, cover is returned onto the furnace and arcing resumes and additional fluxes of 1.5 tons of quick lime and 0.5 tons of dolomite is added through the charging chute. Further melting and refining is continued until tapping the furnace.

The typical target temperature range of the steel at tap is between 1853 to 1903 K with an average of 1878 K using an R-type (Pt/13mass%Rh-Pt) thermocouple inserted into a combined sensor, which is discussed later. Steel temperature is measured after complete meltdown of the input materials and is taken twice. Once at approximately 80% of the tap-to-tap time and before final tap. Temperature as high as 1943 K have been recorded for heats requiring additional heat or due to overshooting. But these cases are anomalies that been filtered in most cases. In the present study, the temperature condition for the thermodynamic evaluations has been set at 1873 K for simplification, which is close to the average tap temperature observed.

2.2. Chemical Composition of the Materials Input and Analysis

The hot metal chemical and thermal characteristics of the major components in the present study are given in Table 2. The hot metal is de-siliconized and hot metal pre-treated to lower excess slag formation in the EAF and lower the overall input sulfur into the steel melt. The incoming temperature of the metal was on average 1595.5 K (1322.5°C). Typical scrap types vary with market conditions and operational parameters such as power input and oxygen lancing are slightly modified depending on the size, density, and the chemistry of the incoming scrap. Although the present study has attempted to unify the type of scraps used in the trials, it was not possible to completely control the type of scraps used in the span of over 200 heats. Table 3 provides the average composition of the scrap used. The cold pig iron units have also been applied as a secondary source of carbon, which have been compared with heats utilizing hot metal. The chemical analysis of the hot metal, scrap, and steel samples was analyzed by an OES (optical emission spectrometer; ARL4460, Thermo-scientific, FL, USA)⁸⁾ and the carbon and sulfur through the LECO C/S analyzer (CS600, LECO, MI, USA). Nitrogen analysis of steel samples was carried out using the N/O analyzer (TC500, LECO, MI, USA). The soluble oxygen and the temperature within the steel melt was measured using the combined oxygen and temperature sensor of Heraeus Electro-Nite, which is based on a stabilized zirconia electrolyte cell and an R-type thermocouple, respectively. Details of the working mechanism of the oxygen sensor can be found in past literature.⁹⁾

Table 1. General characteristics of the 130 ton/ch DC-EAF used in the present study.

Item	Details	Remarks
Furnace type	DC-EAF	Twin shell
Capacity	130 ton/ch
Transformer capacity	102 MVA	(51 MVA·2sets)
Shell dimensions	Dia. = 7000 mm, H. = 3950 mm
Tapping type	EBT	Eccentric bottom tapping
Typical tapping time	50 min

Table 2. Hot metal chemical and thermal characteristics of the major components in the present study.

Element	C	Si	Mn	P	S	Ti	Cu	Sn	Temp (K)
Range	4.2–4.5	0.22–0.59	0.20–0.36	0.077–0.113	0.012–0.034	0.023–0.174	0–0.005	Trace	1554–1637
Avg.	4.35	0.405	0.28	0.095	0.023	0.0985	0.0025	Trace	1595.5

Table 3. Average composition of the input scrap, cold pig iron units, and the tapped steel composition for the present study.

Element	Fe	C	P	S	Cu	Ni	Cr	Sn	N	Balance
Scrap	98.10	0.11	0.02	0.03	0.26	0.10	0.07	0.02	0.0075	1.30
Pig iron	94	4.35	0.1	0.035	–	–	–	–	0.0045	1.515
Tapped steel	99.6	0.03	0.013	0.015	0.058	0.026	0.018	0.003	0.0050	0.237

2.3. Slag and Inclusion Chemistry with Hot Metal Additions

From the slagmaking operation, a typical CaO–SiO₂–FeO system was formed within a CaO/SiO₂ averaging 2.4±0.56, when scrap alone was used, and a CaO/SiO₂ averaging 3.3±0.69, when hot metal of approximately 36% was used. The slag composition was verified using an XRF. The approximate chemical composition of the slag is provided in Table 4. Figure 2 shows the changes in the slag composition from its initial scrap melt down composition and oxygen blowing projected onto the CaO–SiO₂–FeO ternary phase diagram obtained from slag atlas, when scrap and hot metal is used in the charge of the EAF. The initial melt down slag composition is within region ‘A’ and as the FeO content increases with oxygen lancing the CaO additions dissolve into the molten slag resulting in a composition of region ‘B’, when only scrap is used. With hot metal utilization, the carbon content in the steel is significantly higher and the slag contains less FeO during the oxygen lancing period and more efficient dissolution of CaO, which rapidly increases the basicity (CaO/SiO₂) of the slag towards region ‘C’.

Table 4. Average and standard deviation of the slag composition with no hot metal and hot metal utilization of approximately 36%.

Component	T.Fe		SiO2		CaO		Al2O3		MgO		MnO
	Avg.	Stdev.	Avg.	Stdev.	Avg.	Stdev.	Avg.	Stdev.	Avg.	Stdev.	Avg.	Stdev.
All scrap	40	8.5	10.88	2.86	25.23	5.51	2.9	0.88	6.28	2.32	3.8	0.83
36% hot metal	35.8	7.35	9.54	2.38	30.43	5.15	2.88	1.04	7.2	4.03	3.72	0.83

Fig. 2.

Slag compositional changes with hot metal utilization in the CaO–SiO₂–FeO ternary phase diagram adapted from slag atlas.

Inclusions from selected steel samples taken from the tundish were analyzed by the slime extraction method.^10,11) The resulting oxide inclusions were then analyzed using EPMA for elemental Ca and Al to determine the calcium-aluminate inclusions present with use of hot metal.

3. Results and Discussions

3.1. Effect of Process Parameters with HMR (Hot Metal Ratio)

From over 200 heats of hot metal additions from 20 to 36%, the power-on-time (min/ch) in Fig. 3(a) shows a decrease with a slope of 0.34 min/%HM assuming a linear regression of the data. Power on time is the total amount of time within the tap-to-tap time, where power through the transformer is provided and excludes the loading, preparation and tapping times of the EAF.¹²⁾

Fig. 3.

Effect of hot metal additions on the (a) power on time in the EAF and (b) oxygen flowrate (Nm³/hr).

This decrease in the consumed power is expected since there is significant sensible heat within the hot metal.^13,14,15) Possible reasons for the scatter of the data is speculated to be caused by the difference in hot metal chemistry and also the in-coming hot metal temperature. The chemical composition of the steel likely changes the thermal conductivity and heat capacity of the melts, which have been documented elsewhere.^16,17,18) The in-coming hot metal temperatures are between 1554–1637 K, as shown in Table 2. In addition, unexpected process issues such as breakouts at the continuous caster, cold ladles, or prolonged LMF (ladle metallurgical furnace) processing can result in extended power-on-times that can exacerbate plant data scatter.

During the hot metal utilization in the EAF, the power on time is significantly decreased, but the decarburization of the steel increases and thus the oxygen lancing time determines the total steelmaking time of operation. Furthermore, the hot metal additions allows rapid formation of the steel melt and at the final stages of boring approximately 4000 Nm³/h of oxygen is lanced for 8 min. and depending on the foaming behavior, the oxygen flowrate is slowly increased. When the melt is sufficiently formed and slag begins to foam, the oxygen lance is submerged deeper into the slag layer and the flowrate is increased to 6000 Nm³/h for 5 min. This allows an optimal foaming condition of the slag and the oxygen flowrate is again increased to 8000 Nm³/h for 8 min. followed by an oxygen increment to 9000 Nm³/h for the next 5 min. Near the target decarburization content, slag foaming is significantly decreased and the oxygen flowrate is decreased to 5000 Nm³/h. The soluble oxygen content at the final stages of decarburization is targeted at 400 ppm oxygen. The oxygen flowrate pattern is illustrated in Fig. 3(b).

Typical EAF operations have approximately a 10 min. refining stage and must operate in an excess oxygen operation for increased kinetics. This results in more oxygen dissolution in the steel compared to the C/CO equilibrium of reaction (1). The 1 mass% standard state was taken for the calculation of the activities of the dissolved carbon and oxygen in iron. The interaction parameters used for reaction (1) are the calculated values derived from the experimental results of Banya et al.¹⁹⁾   

$$\begin{array}{l}@1\hspace{0.17em}873\hspace{0.17em}K\\ \underset{\_}{C}+\underset{\_}{O}=COlogK=\frac{1\hspace{0.17em}160}{T}+2.003\end{array}$$(1) ¹⁵⁾

The above thermodynamic calculation is carried out at 1873 K for simplification considering the average tap temperature at the EAF was close to 1873 K.

Figure 4(a) shows the dissolved carbon and oxygen before tap for heats with and without hot metal. The dissolved oxygen for the heats with 100% scrap is typically higher than the equilibrium value, but the heats utilizing hot metal is slightly closer to the equilibrium value than heats with all scrap. It would initially seem that the utilization of hot metal would require higher decarburization rates compared to all scrap heats in order to maintain a constant steelmaking process time, which would require higher oxygen flowrates and thus possibly increased oxygen dissolution in the steel. However, if optimized for the input carbon content from the hot metal, this increased decarburization of the melt will increase the formation of CO and CO₂ gases providing better slag foaming and increased mixing in the furnace resulting in higher reaction kinetics and thus moves towards the equilibrium line much faster than scrap alone. And as long as excessively high amount of oxygen, much higher than the necessary input carbon, is avoided the dissolved oxygen content will be closer to the equilibrium value. It should be mentioned that the number of data utilized in Fig. 4(a) and subsequent figures do not correspond completely to Fig. 3(a). This is due to the availability of plant data and in some cases with defective sensor measurements or sampling errors during operation. All available data pertaining to the present study have been utilized for analysis.

Fig. 4.

(a) Dissolved oxygen content as a function of carbon content and (b) decarburization rate with hot metal additions into the EAF specified as input carbon.

The decarburization rate with the input carbon in the EAF charge is shown in Fig. 4(b). Higher input carbon required higher decarburization rates for a final carbon target of 300 ppm.

Figure 5 shows the soluble oxygen content of the steel melt and T.Fe in the slag at tap. Hot metal additions into the furnace allowed lower turndown FeO in the slag compared to the hot metal free heats. The additional carbon available in the melt provided better FeO/C reactions at the slag/metal interface resulting in lower FeO in the slag. This is also evident from the vicinity of the T.Fe results of the hot metal with the equilibrium line compared to the heats without hot metal utilization. The T.Fe (%) and O (ppm) equilibrium curve at 1873 K was calculated from the Fe/FeO equilibrium using the thermodynamic data provided in reference databook¹⁹⁾ The oxygen activity in liquid Fe equilibrated with FeO in slag was calculated from the recommended interaction parameters and the activity of FeO in slag was estimated by assuming an ideal liquid solution. The T.Fe in slag was recalculated by dividing the mass percent of FeO with 1.286. The mathematical formulation is beyond the scope of the present study. A more complex equation derived by Basu et al.²⁰⁾ for the activity of FeO in slag is available, but is not significantly different from the calculation of the activity assuming an ideal slag solution in the range of FeO in the present study.

Fig. 5.

The T.Fe in slag as a function of the soluble oxygen in the steel for hot metal free and hot metal added heats.

3.2. Effect of HMR on the Phosphorous Content

By utilizing hot metal, the time to obtain the molten steel pool is decreased and allows the operator much longer steel refining times before tap, while maintaining the past tap-to-tap times. Depending on the oxygen partial pressure, phosphorous removal from steels through the slag-metal interface can occur by either phosphate ( $P{O}_4^{3-}$ ) at high O₂ or phosphide (P^3–) at low O₂, as expressed in reaction (2) and (3).^21,22,23) Considering the oxygen lancing during decarburization and the amount of FeO present in the slag, the phosphate is likely dominant. It should also be mentioned that for most steel decarburization processes at 1873 K result in an oxygen partial pressure of higher than 10^–16 atm and thus phosphorous removal by phosphate is dominant.   

$$\underset{\_}{P}+\frac{5}{4}{O}_2+\frac{3}{2}{O}^{2–}=\left(P{O}_4^{3–}\right)$$(2)

  

$$\underset{\_}{P}+\frac{3}{2}{O}^{2–}=(P^{3–})+\frac{3}{4}{O}_2$$(3)

With 36% HM input into the EAF, the initial input amount of P increases by 0.02%, but due to longer refining times and better CaO dissolution kinetics into the slag, the distribution ratio of P between the metal and slag (L_P) increases with HM additions, which result in comparable P contents in the final tapped steel with 100% scrap feed EAF processes. If the de-phosphorization percentage (De-P%) is defined as Eq. (4).   

$$De–P\%=\frac{\left(P_{Input}–P_{@LMF}\right)}{P_{Input}}\times 100$$(4)

where P_Input is the overall total P input into the EAF from the hot metal and the scrap feed and P_@LMF is the measured P content at the LMF (ladle metallurgical furnace). Figure 6 shows the De-P% to be on average 50% for 100% scrap and with 36% HM utilization the De-P% to be on average 75%.

Fig. 6.

Phosphorous content at the ladle metallurgical furnace with phosphorous input into the EAF.

The HM utilization into the EAF can also effect the T.Fe in the slag due to the higher carbon content and its subsequent reaction with the FeO, as previously mentioned. According to Fig. 5, the use of HM allows better control of soluble oxygen dissolved in the steel and is closer to the oxygen content achieved from the C/CO equilibrium at steelmaking temperatures resulting in lower T.Fe in the slag. This lower T.Fe and subsequently FeO content and improved reaction efficiency towards the thermodynamic equilibrium value increases the P distribution of the slag comparable to the results of Balajiva and Vajragupta.²⁴⁾ Figure 7 describes the effect of T.Fe in the slag on the phosphorous distribution plotted together with the results of Balajiva and Vajragupta.²⁴⁾ Thus, lower FeO in the slag with hot metal utilization likely improves the phosphorous removal similar to the results obtained from past literature.^23,25) Furthermore, comparison of the phosphorous removal in the present study at basicity of 3.3 with Balajiva and Vajragupta²⁴⁾ at basicity of 2.4 shows the basicity in the present slag compositional range to be insignificant for phosphorous, which has also been suggested by others.^23,25)

Fig. 7.

Effect of T.Fe in slag on the dephosphorization distribution ratio (L_P) in the EAF.

3.3. Effect of HMR on the Sulfur Content

Figure 8 shows the effect of desulfurization with HM of 36% utilization for approximately 200 heats analyzed at tap. The removal of sulfur at the slag-metal interface can be described by two mechanisms of sulfide (S^2–) and sulphate ( $S{O}_4^{2-}$ ), as expressed in reactions (5) and (6).^23,26,27) From the CaSO₄–CaS equilibrium relationship, the critical oxygen partial pressure can be estimated to be approximately 10^–5 atm assuming CaO flux is the dominant desulfurizer in the slag. At the oxygen partial pressure of the primary steelmaking process, the dominant desulfurization mechanism is well known to be the sulfide reaction of (5) and the sulfate reaction of (6) to be negligible, but the low basicity of the present slag seems to counteract the oxygen partial pressure effects and inhibit significant removal of sulfur in the metal. Furthermore, the lower carbon after decarburization results in reduced activity of the sulfur through the changes in the interaction coefficient of sulfur, which decreases the driving force for desulfurization.   

$$\underset{\_}{S}+O^{2–}=S^{2–}+\frac{1}{2}{O}_2$$(5)

  

$$\underset{\_}{S}+\frac{3}{2}{O}_2+O^{2–}={(S{O}_4)}^{2–}$$(6)

Fig. 8.

Comparison of the sulfur content at the ladle metallurgical furnace with sulfur input into the EAF with and without hot metal. Note the resulfurization of the elements in some of the heats.

According to the results, significant correlation of the sulfur with HM additions could not be observed. It is speculated that desulfurization in the EAF is not efficient due to the counteracting effects of basicity and oxygen partial pressure pressures during hot metal utilization. Although faster dissolution of CaO in the slag is observed for hot metal utilization, the high oxygen flowrates counteract the effects of the basicity. Thus, hot metal utilization of the present slag system does not promote the removal of sulfur at the final target carbon content of less than 0.03%.

For some of the heats, excessive slag carryover from the EAF during tapping and un-optimized CaO dissolution into the slag can cause resulfurization at the LMF. This reversion of sulfur from the slag to the steel can take place if the S-rich elements interact with a slag composition with comparatively high FeO, MnO, SiO₂, and low basicity, which have lower desulfurization potential.²⁸⁾

3.4. Effect of HMR on the Nitrogen Content

Although typical EAF processes utilize a top cover and apparently sealed slag door during steelmaking, its effect to seal atmospheric gases from entering the EAF is not absolute. Thus, significant infiltration of nitrogen through the cover and crevices of the slag exit door is possible. As the arc is formed between the electrode and the scrap, the nitrogen gas is ionized and rapidly travels towards the surface of the molten pool, which can dissolve into the steel increasing the nitrogen content.²⁹⁾ A schematic of the typical nitrogen content variations in steels during the EAF operating stages are provided in Fig. 9.³⁰⁾ During the 1st melting period in region B, where the small molten steel pool forms, nitrogen significantly increases above 170 ppm N. As the final charge is melted and a slag layer is formed, N pickup is inhibited. Heating to the steelmaking temperatures with a fully covered slag protects the steel from atmospheric nitrogen pickup. When region E is reached, significant boiling and foaming of the slag occurs, which drives the nitrogen out of the steel and also significantly lowers the partial pressure of N₂, resulting in significantly lower N. During tapping, nitrogen pickup typically occurs through the stream of the steel.

Fig. 9.

Schematic of the nitrogen content of the steel during the operating practice of the EAF adapted from Pilliod.³⁰⁾

The nitrogen content in steels is a function of temperature, steel composition and nitrogen partial pressure according to Eq. (7).   

$$\begin{array}{l}log\left[N\right]=–\frac{1\hspace{0.17em}050}{T}–0.815–0.125\times \left[\%C\right]–0.065\\ \hspace{1em}\hspace{1em}\hspace{0.17em}\hspace{0.17em}\hspace{0.17em}\times \left[\%Si\right]+0.02\times \left[\%Mn\right]+0.5log\left[P_{N_2}\right]\end{array}$$(7)

Equation (7) was derived from the equilibrium reaction between gaseous nitrogen (N₂) and dissolved N in liquid Fe. The activity of the dissolved nitrogen was calculated using the 1 mass% standard state. The recommended interaction parameters from the steelmaking data sourcebook¹⁹⁾ with respect to the steel composition of C, Si, and Mn for calculation of the activity coefficient have been utilized.

Thus to effectively lower the nitrogen in steels, lower temperatures, active CO boiling and subsequent decrease in the nitrogen partial pressure, and an increase in the slag foaming to decrease the possible infiltration into the bath melt should be pursued. Figure 10 shows the comparison of the effect of N in steels between hot metal and cold pig iron unit additions into the EAF. A baseline N content for typical EAF operations would be between 80 to 130 ppm N during tap as can be verified from the plant trials with no carbon saturated iron units within the furnace. For region 1, the additions of carbon saturated cold pig iron into the furnace did not significantly lower the nitrogen content up to approximately 22%. Although the higher C content in the cold pig iron was expected to improve CO gas formation and subsequent lowering of the N content in the steel, results showed that the nitrogen content did not deviate much from the 100% scrap operation at low amounts of additions. This was speculated to be the increased power input in Region 1 of the dense cold pig iron units compared to 100% scrap operation due to its higher density and slower melting rate during operation. Increased additions of carbon saturated containing iron units showed comparable lowering effects of nitrogen content due to the significantly higher CO evolution and subsequent slag foaming that would suppress the infiltration and dissolution of nitrogen into the steel. This critical ratio of 22% carbon saturated iron unit additions was obtained from the intercept of the average nitrogen value with no carbon saturated iron units at 105 ppm N extended towards the addition of carbon saturated iron units and the linear regression line of data points with carbon saturated iron units. As the amount of carbon saturated iron units exceeds 22% of the charge, the overall trend seems to indicate that the addition of carbon containing iron units into the EAF decreases the nitrogen content in the tapped steel. In particular, the addition of hot metal slightly decreases the N content compared to the cold pig iron unit additions due to the decrease in the power on time of the EAF and subsequent decrease in the nitrogen ionization through the plasma arc. In particular the region B and C in Fig. 9 would significantly decrease with hot metal additions. However, it should also be noted that the nitrogen dissolution into the metal is likely more affected by the evolution of CO gases during decarburization and its subsequent lowering of the nitrogen partial pressure from comparison between the cold pig iron and hot metal heats. Furthermore, the addition of carbon containing iron units into the steel melt provides a more rapid slag formation and foaming practice that can inhibit the atmospheric infiltration into the melt and also provide a more steady arcing stream, which can lower the nitrogen dissolution into the melt. The negligible effect in region 1 is speculated to be due to the counteracting effect of the improved CO evolution and the increased power consumption of the slow melting of the pig iron units.

Fig. 10.

Effect on the nitrogen content by additions of carbon saturated iron units including cold pig iron and hot metal at steel tap from the EAF.

3.5. Effect of HMR on the Inclusion Chemistry

For HMR of 36% utilization, the steel is affected in three aspects of internal quality. First, the input material contains significant amounts of carbon forming large amounts of CO during the melting and refining stages of the steelmaking process, which prevent nitrogen infiltration into the furnace and also improved foamy slags inhibiting atmospheric gas penetration into the steels. This result in about 1/3 of soluble nitrogen in the hot metal utilized steels compared to hot metal free heats and possibly decreased nitride inclusions. Second, sufficient carbon in the steel at the end of the melt down cycle of the EAF allow lower free oxygen in the steels after the refining stages lowering the oxygen content by about 100–150 ppm. Third, the hot metal dilutes the amount of tramp elements such as Cu, Ni, Sn within the scrap,¹³⁾ which significantly lowered visible crack defects of 2 mm and longer (defect ratio %) by 50% on the hot coil during casting and rolling.

Due to the mini-mill operation of the present study, where a thin slab caster and continuous hot strip mill process is adopted, the process characteristic requires significant control of inclusions for production of highly value-added steel products. Figure 11 shows the quality comparison of identical steel products produced through 100% scrap and 36% HM utilization in the DC-EAF with comparable secondary metallurgy practices. The defect index is typically obtained from the visual inspection results of the surface of the product, which is determined from the length, depth, and number of the defect. The combination of these variables determines the defect index (%), which is a comparison to the reference value set-up for each individual steel grade from ultra-low carbon to high carbon steels. The details of the reference value and the equation are proprietary information to the specific company, but have become more stringent as customer demands become more critical. For example, the reference value for a crack sensitive peritectic grade may be 10 cracks with 4 mm length by 2 mm depth, but a low carbon grade may be 5 cracks with 3 mm length by 2 mm depth within the plant. Typically, the total content of (Cu+10Sn) is an essential control criterion for product quality. Increased amounts of Cu and Sn in the steel tend to result in significant downstream quality issues such as surface hot shortness cracks^31,32) and are constrained below a total of 0.1 (Cu+10Sn). With hot metal additions of 36%, the tramp elements of Cu and Sn is correspondingly diluted and ensures the critical defect index to be less than 10%. The critical defect index is the limit, where quality codes are changed on the product of the steel. Below the critical defect index of 10%, products are not downgraded and above this critical value downgrades of the products are typically required. For heats that have been added with hot metal of 36%, the Cu+10Sn value is less than 0.1% resulting in critical defect index values lower than 10%. For typical all scrap heats with no hot metal, the Cu+10Sn values exceed 0.1 and some of the inspections of products resulted in defect index above the critical defect index of 10% resulting in product downgrades and subsequent loss in quality. The critical defect index of 10% at Cu+10Sn is an empirical value that has been developed within the plant, which has been found to be free of steel product downgrades. Not all identical steel grades in different heats have exactly the same steel chemistry since there is a range of chemical compositions acceptable for qualification. The direct cause of the steel samples exceeding Cu+10Sn of 0.1 for all scrap heats, but lower than the critical defect index of 10% has yet to be fully understood. It may be speculated that other elements in the steel such as Ni could have provided increased Cu-solubility in the austenite to lower the propensity for hot cracks to form. And if the Ni content was relatively high, but within the range of acceptable steel chemistry, it may have provided beneficial effects to the steel. It has been well-defined in past publications^32,33) that Ni content approximately half of the copper can inhibit surface defects in steels. Shibata et al.³⁴⁾ has also shown that higher Si content and P up to 200 ppm can lessen the impact of the surface hot shortness and crack susceptibility of the steel sample. Thus, various steel chemistries may have assisted with lowering the defect index, although the Cu+10Sn exceeded 0.1. According to the quality results, a 50% reduction in the average quality defect index was observed, when hot metal is used as previously mentioned.

Fig. 11.

Changes in the defect ratio index with 36% hot metal utilization in the EAF.

The dominant inclusion chemistry from analysis of the total Ca and Al content in the steel was identified to be closer to the C₃A (3CaO·Al₂O₃) with the inclusions using hot metal or scrap alone as shown in Fig. 12. Using the thermodynamic data and relationships in reaction (8) and (9) from Turkdogan,³⁵⁾ the equilibrium line between Al and Ca for the calcium-aluminate type inclusions expected of forming at steelmaking temperatures of 1823 K could be ascertained and plotted in Fig. 12. At low solute concentrations of Al and Ca in the steel, the activities of Ca and Al were taken equivalent to the mass% of Ca and Al, respectively. Al₂O₃, CaAl₂O₄, and Ca₃Al₂O₆ were considered to be in the pure standard states.   

$$\begin{array}{l}\underset{\_}{Ca}+\frac{4}{3}A{l}_2{O}_3=CaA{l}_2{O}_4+\frac{2}{3}\underset{\_}{Al}\\ \mathrm{\Delta }{G}^o=57\hspace{0.17em}354–85.3\times T(J/mol)\end{array}$$(8)

  

$$\begin{array}{l}3\underset{\_}{Ca}+2A{l}_2{O}_3=C{a}_{3}A{l}_2{O}_6+2\underset{\_}{Al}\\ \mathrm{\Delta }{G}^o=207\hspace{0.17em}471.76–229.52\times T(J/mol)\end{array}$$(9)

Fig. 12.

Inclusion chemistry analysis of tundish samples after the LMF with hot metal utilization of 36% and no hot metal utilization in the EAF. Note the EPMA analysis of Ca and Al on the surface of separated inclusions using the slime method is shown.

From the results of the inclusion analysis, the impact of hot metal on the effect of inclusion chemistry does not seem to have significant impact as the inclusion modification treatment has been appropriately optimized for the dissolved oxygen content and Al in the steels before the steel is sent to the tundish for casting. The inclusion composition seems to indicate slightly excess Ca treatment for the steels, but is comparable for both the hot metal free and 36% hot metal added heats. The typical 3CaO·Al₂O₃ inclusion morphology observed after treating with the slime method is shown adjacent to Fig. 12 labeled C₃A. Although not thermodynamically stable for the mass% of Ca and Al in the steel sample, Al₂O₃ inclusions are also observed within the melt labeled A. It is likely that some Al₂O₃ inclusions could not effectively interact with the Ca injected at the LMF for the given processing time and fluid flow conditions during secondary refining and thus remained as Al₂O₃ in the melt. This suggests that the reactions involved during inclusion modification in secondary refining do not reach complete thermodynamic equilibrium in actual operations.

4. Conclusions

In this study, the effect of hot metal additions into a DC EAF was investigated. The dissolved oxygen for heats with 100% scrap was typically higher than the equilibrium value, but the heats utilizing hot metal was slightly closer to the equilibrium value. Hot metal additions into the furnace allowed lower turndown FeO in the slag compared to the hot metal free heats. With 36% hot metal input into the EAF, although initial P was higher than the 100% scrap heats, hot metal utilization resulted in longer refining times with better CaO dissolution kinetics into the slag increasing the distribution ratio of P between metal and slag, which attained comparable P contents in the final tapped steel. Hot metal utilization for the basicity and oxygen partial pressure of the present slag system did not particularly promote sulfur removal compared to the 100% scrap heat. The addition of carbon saturated iron units up to 22% into the furnace did not significantly lower the nitrogen content, but as the amount exceeded 22% of the charge, the overall trend seems to indicate the nitrogen content in the tapped steel to decrease. In particular, hot metal additions slightly decreases the N content compared to the cold pig iron unit additions due to the decrease in the power on time of the EAF. With hot metal additions of 36%, the tramp elements of Cu and Sn is correspondingly diluted and ensured the critical defect index to be less than 10% resulting in 50% reduction of the quality defect index. The existing inclusion chemistry of the heats was identified to be closer to the C₃A (3CaO·Al₂O₃), but was comparable for both the hot metal free and 36% hot metal added heats.

Acknowledgements

This study was partially supported by the Brain Korea 21 PLUS (BK21 PLUS) Project at the Division of the Eco-Humantronics Information Materials and the Ministry of Trade, Industry, and Energy project No. 2013-11-2008.

References

• 1)   J. P.  Birat: Rev. Métall., 97 (2000), 1347.
• 2)   I. Y.  Zinurov,  A. M.  Shumakov,  S. G.  Ovchinnikov,  S. N.  Pigin and  Yu. F.  Mamenko: Stal, 7 (2009), 35.
• 3)   J.-P.  Duan,  Y.-L.  Zhang and  X.-M.  Yang: Int. J. Min. Metall. Mater., 16 (2009), 375.
• 4)   P.  Argenta and  M. B.  Ferri: La Metallurgia Italiana, 97 (2005), 41.
• 5)   T. P.  Wandekoken,  B. T.  Maia,  P.  Hopperdizel and  J. R.  de Oliveira: 43rd Seminario de Aciaria – Internacional Belo Horizonte Brazil, Brazilian Metallurgical, Materials and Mining Association Brazil, Brazil, (2012), 75.
• 6)   H.  Liu,  J.  Li and  C.  Ma: Adv. Mater. Res., 572 (2012), 243.
• 7)   R.  Gottardi,  S.  Miani,  A.  Partyka and  M.  Suber: Iron Steel Technol., 9 (2012), 61.
• 8)   J.-Y.  Park,  S.-M.  Jung and  I.  Sohn: Metall. Mater. Trans. B, 45 (2014), 329.
• 9)   J. W.  Fergus: http://www.tms.org/pubs/journals/jom/0010/fergus/fergus-0010.html, (accessed 2014-10-10).
• 10)   J.  Wakita and  S.  Xizoguchi: Tetsu-to-Hagané, 66 (1980), 5862.
• 11)   H.  Hoef,  H.  Lessing and  G.  Masing: Stahl Eisen, 76 (1956), 1442.
• 12)  World Steel Association: http://www.steeluniversity.org/content/html/eng/default.asp? catid=25&pageid=2081271946, (accessed 2014-08-20).
• 13)   C. P.  Manning and  R. J.  Fruehan: JOM, 53 (2001), 20.
• 14)   A.  Orth,  N.  Anastesijevic and  H.  Eichberger: Miner. Eng., 20 (2007), 854.
• 15)   T.  Jiemin,  M. B.  Ferri and  P.  Argenta: Ironmaking Steelmaking, 32 (2005), 191.
• 16)   T.  El Gammal and  X.  Li: Mater. Sci. Forum, 73–75 (1991), 37.
• 17)   M.  Susa,  M.  Watanabe,  S.  Ozawa and  R.  Endo: Ironmaking Steelmaking, 34 (2007), 124.
• 18)   M. J.  Peet,  H. S.  Hasan and  H. K. D. H.  Bhadeshia: Int. J. Heat Mass Transf., 54 (2011), 2602.
• 19)  The Japanese Society for the Promotion of Science: Steelmaking Data Sourcebook Rev. Ed., Gordon and Breach Science Pub., New York, NY, (1988), 59.
• 20)   S.  Basu,  A. K.  Lahiri and  S.  Seetharaman: Metall. Mater. Trans. B, 39B (2008), 447.
• 21)   H.  Suito and  R.  Inoue: ISIJ Int., 35 (1995), 258.
• 22)   X.  Guo,  H.  Matsuura,  M.  Miyata and  F.  Tsukihashi: ISIJ Int., 53 (2013), 1381.
• 23)   N.  Sano,  W.-K.  Lu,  P. V.  Riboud and  M.  Maeda: Advanced Physical Chemistry for Process Metallurgy, Academic Press, San Diego, CA, (1997), 51.
• 24)   K.  Balajiva and  P.  Vajragupta: J. Iron Steel Inst., 155 (1947), 563.
• 25)   H.  Kishitaka: The 43rd Nishiyama Memorial Seminar, ISIJ, Tokyo, (1977), 50.
• 26)   Y.  Nakai,  N.  Kikuchi,  Y.  Miki,  Y.  Kishimoto,  T.  Isawa and  T.  Kawashima: ISIJ Int., 53 (2013), 1020.
• 27)   P.  Yan,  X.  Guo,  S.  Huang,  J.  Van Dyck,  M.  Guo and  B.  Blanpain: ISIJ Int., 53 (2013), 459.
• 28)   J.  Lehmann and  M.  Nadif: Rev. Mineral. Geochem., 73 (2011), 493.
• 29)   B.  Bowman and  K.  Kruger: Arc Furnace Physics, StahlEisen Communications, Düsseldorf, (2009), 5.
• 30)   C. F.  Pilliod: 46th Electric Furnace Conf. Proc., The Iron and Steel Society, Pittsburgh, (1988), 107.
• 31)   A.  Takemura,  Y.  Ugawa,  K.  Kunishige,  Y.  Tanaka,  S.  Hashimoto and  K.  Ootsuka: Mater. Trans., 52 (2011), 1905.
• 32)   N.  Imai,  N.  Komatsubara and  K.  Kunishige: ISIJ Int., 37 (1997), 224.
• 33)   ASM  International: ASM Metals Hand Book, 10th ed., Properties and Selection: Iron; Steels and High Performance Alloys, Vol. 1, Materials Park, OH, (1995), 389.
• 34)   K.  Shibata,  S.-J.  Seo,  M.  Kaga,  H.  Uchino,  A.  Sasanuma,  K.  Asakura and  C.  Nagasaki: Mater. Trans., 43 (2002), 292.
• 35)   E. T.  Turkdogan: Physical Chemistry of High Temperature Technology, Academic Press, New York, (1980), 5.