Optimum Lining Performance for Particular Process Slags in Metallurgical Vessels Supported by Thermochemical Modeling Marcus Kirschen RHI AG, Process Technology and Systems Solutions Wienerbergstrasse 9, A-1100 Vienna, Austria
[email protected] Ronald Lanzenberger, Bernd Petritz RHI AG, Product Development and Management Basic Mixes Magnesitstrasse 2, A-8700 Leoben, Austria
[email protected],
[email protected] Thomas Prietl RHI AG, Product Marketing Nonferrous Metallurgy Wienerbergstrasse 9, A-1100 Vienna, Austria
[email protected]
Key words: Corrosive wear, thermochemical modeling, slag, EAF, BOF, tundish
ABSTRACT A thorough understanding of refractory lining corrosion processes in metallurgical reactors is crucial to improve lining lifetime and to decrease lining maintenance costs. Besides laboratory tests and plant trials, the modeling of corrosion processes provides an efficient and cost-effective tool to define the appropriate refractory material for a particular process slag composition. Thermochemical modeling was applied to: (1) study the interaction between covering slags of very distinct chemical composition with the tundish lining material, (2) understand the performance of repair mixes in contact with basic oxygen furnace (BOF) and electric arc furnace (EAF) slags, and (3) investigate the stable phases of infiltrated lining materials in a melting furnace with a process slag of very special chemical composition. The advantages and limitations of a phase equilibrium approach to examine corrosion processes in industrial metallurgy are discussed. MOTIVATION Mechanisms that lead to premature wear in metallurgical process vessels are complex due to the harsh process conditions in steel making, including heat transfer by electric arcs with ultra-high power intensity in the order of 30–100 MW and intensive heat generation by exothermal oxidation reactions with maximum temperatures above 1750 °C. Batch processes require melt tapping and furnace vessel opening for charging or lining repair, resulting in periodic, transient thermally-induced stress fields in the linings. Intensive bath and slag layer movement in the BOF, EAF, and treatment ladle is typically an important process target to increase melting and chemical reaction rates, and to decrease process time. Increased mass transfer in the steel melt and slag by increasing melt movement, for example with bottom gas purging plugs or by using supersonic jets for gas injection also increases refractory lining
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erosion. In addition to these inevitable process-related reasons for vessel lining wear, a third and constant factor for a finite lining lifetime is corrosive wear by the steel melt and slag, as components of the refractory material are soluble in the slag phase, especially at high temperature. In most cases, dissolution of refractory components into the process slag is more pronounced than dissolution into the steel or nonferrous melt. The solubility of refractory lining components in the slag is directly related to the difference in chemical potentials of the components in the refractory phases and slag, which depends at a given temperature solely on the composition of the contacting phases. Therefore, minimizing chemical potential differences between the refractory lining and process slag is considered one option to increase lining lifetime, besides measures to decrease thermal overload and thermal cycling of the lining. This can be achieved by an appropriate choice of lining materials for a given process slag composition or by modifying the process slag chemical composition to within certain ranges that are given by particular process requirements, for example slag viscosity, sulphur capacity, slag mass, and costs. In line with this strategy, examples are provided in this paper to demonstrate how the performance and lifetime of some metallurgical vessel linings was improved through an analysis of the refractory–slag interactions based on thermochemical modeling. APPLICATION OF THERMOCHEMICAL MODELING TO CORROSION PROCESSES The term thermochemical modeling is used in this paper as a synonym for applying the phase equilibrium approach to calculate the amount and compositions of phases in thermal and chemical equilibrium when the Gibbs free energy of the system is at a minimum. The FactSage1 program was used to determine the stable phase assemblages as the software is based on precise and widely accepted models for the pure and multicomponent solution phases in the CaO-MgO-SiO2-Al2O3-FeOx system. Additional components including MnO, CrOx, Na2O, K2O, P2O5 were also incorporated and validated for lower order systems. As the basic assumption of this approach is the thermal and chemical equilibrium of the system, the first limitation is related to thermal gradients in the linings and suppressed mass transfer, for example by carbon layers in pitch or resin-bonded materials. The detailed physics of mass transport from the slag bulk to the reaction layer and vice versa by turbulent flow and diffusion was considered a more or less constant factor in the corrosion processes. As these effects depend on the slag bulk chemistry, namely the saturation concentration, they are neglected when the corrosion capacity of one customer’s slag to different lining materials is discussed in this paper. Whereas thermal gradients can be assessed easily with thermal simulations, detailed information about diffusion processes and mass transport coefficients in infiltrated layers of the lining is restricted to some systems investigated under well-defined laboratory conditions2,3. Extrapolation of diffusion coefficients in oxide mineral systems to refractory linings of higher complexity is difficult. To handle this essential restriction to a more comprehensive corrosion model local equilibrium on a microscopic scale was assumed by carefully selecting the chemical composition of the system examined. When considering the local equilibrium in the lining in contact or infiltrated with slag, there are three principally approaches to represent the bulk composition of the individual system investigated (Figure 1): 1. 2. 3.
Infiltrating slag with the binding material of the brick or monolithic lining Infiltrating slag with the binding material and a proportion of the coarse grains in the brick or monolithic lining Infiltrating slag with the entire fine-grained lining material
A
B
C
Figure 1. Illustration of the three approaches to determine the bulk composition in local equilibrium in infiltrated bricks or EAF hearth and repair mix layers (A and B) and in fine-grained tundish slurry mixes (C). The first simulation approach is based on the assumption that the lining is destabilized and worn when the binding material is dissolved in the slag4. In this case the saturation limits of the infiltrating slag in relation to the ceramic binding components, for example Mg2SiO4, CaO, Ca2SiO4, are crucial to model the corrosion process (Figure 1A). However, it was determined in mineralogical postmortem analyses of BOF and EAF mixes that a precise prediction of the observed phase mass ratios required a proportion of the coarse grain fraction to be considered in equilibrium, mainly the MgO grain rims with a certain defined reaction depth of the coarse grain fraction (diameter up to 6 mm) in ANKERHARTH, ANKERREP, and ANKERJET mixes in contact with the slag (Figure 1B)5. However, despite the higher precision of the predicted phase amounts, this approach is expensive to apply to phase assemblages in other studies as the typical rim thickness that reacts with the infiltrated slag has to be determined by new laboratory trials and mineralogical studies. The easiest way to determine the bulk composition in local equilibrium is to assume that the total refractory material reacts with the infiltrating slag (Figure 1C). However, this approach is only valid for very fine-grained material, where the grain surface to grain volume ratio is sufficiently high to provide high reaction rates in the infiltrated layer and to minimize chemical composition gradients in one phase (e.g., Figure 26). Then the diffusion process time scale is not significantly longer than the
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timescale of the infiltration and corrosion processes resulting in rather low concentration gradients in the phases. This approach was applied to ANKERTUN tundish slurry, dry-setting, and self-hardening mixes as the typical maximum grain diameter is 0.5 mm. With regard to phase equilibrium studies in magnesia-based linings with basic slags in EAF, BOF, and tundish vessels, the investigated phase assemblages comprised only a few solids and solid solutions (ss) in the MgO-CaO-FeOx-SiO2-Al2O3 system: mainly periclase, forsterite, C2S, CaO, and MA spinel at process temperatures > 1500 °C and additionally a few Ca-Mg-silicates and Ca-Al-silicates at lower temperatures (Figure 2). Phase assemblages from postmortem analyses were used to check the thermochemical simulation results. The phase assemblages considered in simulations for lining concepts for nonferrous EAFs and smelters ranged from magnesia-based to MgO-Cr2O3-based and Al2O3-based bricks. The phase assemblages are far more complex when, for example alumina-based materials are combined with basic slag. In addition, the variety of process slag compositions in nonferrous metallurgy, for example the Cr2O3, SiO2, MnO, NiO, PbO, ZnO, and As2O3 content, is significantly higher than in the steel industry.
Figure 2. Postmortem analysis of a slag infiltrated CaO-rich ANKERTUN CS10 tundish mix showing periclase (1), monticellite (2), and spinel (3) as stable phases and an oxidic liquid indicating very small grain sizes6. The tundish cover slag was rice husk ash. The original structure between the periclase grains was destroyed. An absence of chemical gradients indicates conditions close to equilibrium In order to model corrosion of the refractory lining material in contact with slag with a local equilibrium approach the mass ratio between attacking slag and lining material (Figure 3) was varied from x = 0 to 30% and the stable phase assemblage was determined. This range corresponds to the mass ratio that was estimated from partially filling the initial open porosity of the lining with slag. For example, the open porosity of a densified ANKERHARTH lining decreased from 20% to 5–10% when infiltrated with slag, indicating a slag mass ratio of approximately x = 10–15%. The open porosity of tundish ANKERTUN linings is between 20%–40% depending on preparation and treatment of the material as well as the process temperature.
Lining mix
Process slag x
Figure 3. Illustration of the corroded layer modeling approach that considers an increasing amount of slag mass in equilibrium at the hot side. x denotes the local mass ratio between the slag and refractory material APPLICATION TO TUNDISH LININGS Using the presented modeling approach, the calculated phase assemblages of an infiltrated tundish lining are shown in Figure 4. The phase assemblage of pure lining material is shown at the ordinate x = 0. RHI provides three classes of tundish lining mixes: High-MgO ANKERTUN mixes (MgO > 88 wt.%) SiO2-rich ANKERTUN mixes (12–17 wt.% SiO2) CaO-rich ANKERTUN mixes (8–22 wt.% CaO) based on alpine magnesia sinter as, for example the ANKERHARTH mixes
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The stable phase assemblages of a CaO-rich ANKERTUN 280, a SiO2-rich ANKERTUN DS20T, and a high-MgO ANKERTUN DS70 lining in contact with a customer CaO-MgO-silicate cover slag were compared (Figures 4A, 4C, and 4E). In order to simulate the lining performance when ladle slag is carried-over, the equilibrium phase assemblage with the ladle slag consisting of CaO-MgOSiO2-Al2O3 was also calculated (Figures 4B, 4D, 4F). In all six simulation cases, MgO saturation was immediately achieved in the infiltrated layer, which is not surprising as MgO was in excess (also see Figure 2). More significant was the dissolution or saturation behavior of the secondary phases that represents the binding material between the MgO grains. For this particular tundish cover slag, the forsteritic content of the ANKERTUN DS20T mix decreased slightly with the infiltrated slag mass in favor of an oxidic liquid (Figure 4C). This corrosion effect was even more pronounced when Al2O3-containing ladle slag was carried over into the tundish (Figure 4D). The use of a CaO-rich ANKERTUN mix in the particular customer tundish resulted in a C2S saturation of the infiltration by tundish cover slag (Figure 4A). The formation of any additional solid decreased the amount of destabilizing oxidic liquid and possibly decreased the mass transport due to the formation of an accretion layer. This effect was maintained even when ladle slag was carried over to the tundish (Figure 4B). 100%
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Oxide liquid
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Oxide liquid
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CaO ss
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MgO ss
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CaO ss
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Slag to refractory mass ratio [%]
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Oxide liquid
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Slag to refractory mass ratio [%]
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Slag to refractory mass ratio [%]
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Oxide liquid
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Oxide liquid
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5
10
15
20
Slag to refractory mass ratio [%]
25
30
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5
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15
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30
Slag to refractory mass ratio [%]
Figure 4. Stable phases calculated for CaO-rich ANKERTUN 280 (A and B), SiO2-rich ANKERTUN DS20T (C and D), and highMgO ANKERTUN DS70 (E and F) tundish mixes infiltrated with CaO-SiO2-rich tundish cover slag (A, C, and E) and ladle slag (B, D, and F), T = 1570 °C
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Using the presented simulation approach, the combinations of various ANKERTUN mixes with a series of commercially available synthetic tundish cover slags (Table I), customer tundish slags, and ladle slags were examined. The surprising variation of appropriate and less appropriate combinations of cover slags and tundish lining mixes was due to the varying chemical composition of infiltrated oxidic liquid in the lining material. The results indicate the application of one tundish lining mix may be appropriate for one customer process whereas the use of a different tundish slag requires the application of another tundish lining mix for optimum lifetime. In addition, the influence of ladle slag on tundish lining lifetime depends specifically on the composition of the carried-over ladle slag. This analysis has been confirmed by in service experience, for example high-MgO ANKERTUN 228 showed the best results at a producer of wire, automotive, and construction steels in Eastern Europe. Table I. Matrix of more and less appropriate combinations of various ANKERTUN mixes with synthetic tundish cover slags (i.e., SiO2-rich: 12–17 wt.% SiO2, CaO-rich: 8–22 wt.% CaO, and high-MgO: MgO > 88 wt.%)
ANKERTUN mix Synth. tundish cover slag 1 Synth. tundish cover slag 2 Synth. tundish cover slag 3 Synth. tundish cover slag 4 Synth. tundish cover slag 5 Synth. tundish cover slag 6 Synth. tundish cover slag 7 Synth. tundish cover slag 8
SiO2-rich A -++ ++ + + o --
SiO2-rich B -+ + o o --
CaO-rich A + -+ --
CaO-rich B o + ++ o -
High-MgO A o + + + o o o -
Table II. Matrix of more and less appropriate combinations of various ANKERTUN mixes with customer slags
ANKERTUN mix SiO2-rich A SiO2-rich B Customer tundish slag 1 Customer ladle slag 1 --ANKERTUN mix Competitor mix Customer tundish slag 1 ANKERTUN mix SiO2-rich C Customer tundish slag 2 Customer ladle slag 2a -Customer ladle slag 2b -ANKERTUN mix SiO2-rich D Customer tundish slag 3 + ANKERTUN mix SiO2-rich D Customer tundish slag 4 + Customer ladle slag 4 -ANKERTUN mix SiO2-rich E Customer tundish slag 5 Customer ladle slag 5 -*: Mix applied by the customer with very effective results
CaO-rich A ++ CaO-rich C ++ CaO-rich B o o + CaO-rich E o* CaO-rich E o CaO-rich E o +
CaO-rich B ++ o
CaO-rich D o o ++
CaO-rich F ++ ++
High-MgO A + o
High-MgO A + ++ ++ High-MgO D +* High-MgO D + + High-MgO E ++ +
High-MgO B + High-MgO C ++ * High-MgO C o ++ ++
High-MgO F + +
Besides the corrosion resistance of tundish lining mixes to tundish cover slag, sintering of the lining mix is important as complete sintering results in unacceptable shell temperatures. In order to model the sintering process, the phase equilibrium simulation was combined with thermal simulations of the tundish lining. It was assumed that the formation of the intrinsic oxidic liquid resulted in the observed sintering of the lining material. Formation of the oxidic liquid as a function of temperature is shown in Figure 5. The calculations indicate sintering of SiO2-rich ANKERTUN 217 occurs at lining temperatures higher than 1340 °C, producing a very thin sintered layer (approximately 10 mm thick) at lining temperatures > 1340 °C close to the hot side. The calculated sintered layer of ANKERTUN 280, for example, was approximately 40–50 mm due to occurrence of oxidic liquid above > 1210 °C. These results are in close agreement with observations at a reinforced bar steel producer. The temperature field of the ANKERTUN 217 lining of a 4strand tundish is shown in Figure 6 and the results indicate a reasonably thin sintered layer.
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80
60 50
40
A
60
30 20
MgO ss CaO ss Oxide liquid Liquid, model C2S Ca2Fe2O5
70
Phase [%]
70
Phase [%]
80
MgO ss Mg2SiO4 Oxide liquid Liquid, model CaMgSiO4 Ca3MgSi2O8
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40 30 20
→ 10 mm sintered layer thickness
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→ 40-50 mm sintered layer thickness
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Temperature [°C]
1.400
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Temperature [°C]
Figure 5. Calculated stable phases of SiO2-rich ANKERTUN 217 (A) and CaO-rich ANKERTUN 280 (B)
B
A
Impact pad
Well block
Figure 6. Calculated temperature field in a SiO2-rich ANKERTUN 217 lining of a 4-strand tundish (A) and sintered layer of a SiO2-rich ANKERTUN 217 wear lining (B) GUNNING AND REPAIR MIXES FOR BOF AND EAF APPLICATIONS Similar thermochemical simulations were performed for EAF hearth mixes (i.e., ANKERHARTH) and gunning and repair mixes for EAFs and BOFs (i.e., ANKERREP, ANKERGUN, and ANKERJET), although the very special process conditions at very high temperatures in EAFs and BOFs as well as significant bath movement may significantly override corrosion processes based on refractory component dissolution. In addition, the BOF and EAF process slag compositions are in a much narrower range than for example the tundish cover slags or slags in the nonferrous industry. Typically the phase compositions of the mixes play a more important role when applied as hot repair mixes to the EAF or BOF lining. Effective adhesion of the mix to the lining wall is crucial to avoid unnecessary material losses to the process slag. The gunning mix producer has three basic approaches to adjust this property: Careful selection of the raw material Optimizing the grain size distribution Optimizing the binder combination RHI favors a low iron sintered magnesia with more than 90 wt.% MgO and a CaO/SiO2-ratio of more than 1.8 in the primary raw material component for high quality gunning mixes. Along with the precondition of a high CaO/SiO2 ratio, attention must also be paid that the sum of Fe2O3 and Al2O3 oxides does not exceed a limit of 1 wt.%. In addition, these mixes need lime-rich additives to enable the binders to completely form high melting point phases such as CaO, phosphates, or Ca2SiO5, without the formation of low melting
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silicates. Lime-rich additives to react with the binder are most effective if distributed as fines within the mix. Furthermore, highquality gunning mixes contain additional free lime. This additional lime raises the basicity of an infiltrating slag in order to avoid, as far as possible, the supplementary formation of low melting silicates by foreign phase infiltration. The additional free lime could be added in the form of sintered dolomite fines, however the associated volume expansion after application could damage the gunned layer. A more appropriate material is a lime-rich alpine sintered magnesia with approximately 25 wt.% CaO. The lime distribution in this material is inhomogeneous resulting in a relatively high slaking resistance. With regard to the grain size distribution, care must be taken that the dry mix flows evenly in the machine and that it does not segregate in the hose. The fines improve the plasticity required for adherence, however they must not exceed a certain limit as the mix will spall during drying or the gunned layer will become too unstable. The water slurry incorporates with the grains and the coarse grain stabilizes and densifies the gunned layer, whose bulk density should be approximately 2.6 g/cm³ after drying, and increases corrosion resistance. However, too many coarse grains cause increased rebound rates and thus material loss. The correct grain size distribution is the most important precondition for an optimum gunning mix. The ideal grain range is 0–3 mm, following as closely as possible the grain size recommendation according to Dinger and Funk. If the percentage of smaller grains less than 0.063 mm is too high, there will inevitably be an inferior, uneven output from the machine. In addition, spalling during gunning increases due to an inadequate vapor permeability of the mix. In contrast, if the percentage of fines is too low, plasticity and thus primary adherence during gunning is diminished. The selection of binding agents has to comply with the following criteria: Optimal adherence due to quick stiffening and setting Sufficient mix strength over the entire application temperature range due to appropriate binder selection Minimized reduction of wear resistance resulting from low melting residues from the binding agents
Oxide liquid in infiltrated layer [%]
Basically there are two binder combinations used in high quality gunning mixes: (1) the silicate bond based on the use of watersoluble alkali silicates and (2) the phosphate bond based on the use of various phosphates of varying acidity and different degrees of polymerization. Generally, in top-grade gunning mixes RHI uses the phosphate-bonding system because of better adherence and wear resistance. Nevertheless, the silicate bond is favored for certain processes. It is relatively cheap and can be employed universally, for example in special steel producing furnaces where phosphorisation of steel melt by corroded refractories must be avoided. In addition, it is also more handleable during application as a result of its lower setting rate. Phosphates are known to react more rapidly with basic raw materials if they are more acidic. For example, very acid phosphates like mono-aluminum-phosphate Al(H2PO4) or urea phosphate CO(NH2)2-H3PO4 react and set within a few seconds. This timeframe is too short for gunning mixes because the surface of the gunned layer must remain soft for long enough to incorporate and embed the coarse grains. Furthermore, very short setting reaction times increase the risk of lance clogging. With a combination of different phosphates, stiffening of the wet gunning mix occurs within an optimum period. This is achieved not by thixotropy which is a reversible process, but by an irreversible chemical reaction which subsequently also results in an increased material strength. However, once the mix is applied, infiltration of process slag occurs due to the residual porosity in the applied repair mix layer. If the viscosity of the infiltrated layer decreases drastically, the applied ANKERJET layer may dissolve more rapidly into the slag. Therefore, the amount of liquid as a measure of the tendency for dissolution and wear was calculated for a series of converter mixes infiltrated with a typical converter slag (i.e., 51 wt.% CaO, 15.7 wt.% SiO2, 17.5 wt.% Fe2O3, 1.6 wt.% Al2O3, and 2.1 wt.% MgO) (Figure 7). The resulting amount of oxidic liquid was strongly dependent on the Al2O3 content of the mix, as the Al2O3 content of the BOF slag is usually low. 50 45
Converter slag, T = 1750 C
40 35
30
Top grade mixes
25
Standard quality
20
15 10
ANKERJET mixes Competitor mixes
5 0 0,0
0,5
1,0 1,5 2,0 Al2O3 in converter repair mix
2,5
3,0
Figure 7. Calculated amount of oxidic liquid as measure of wear potential in infiltrated ANKERJET mixes
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MODELING SUPPORT FOR APPROPRIATE LINING MATERIAL SELECTION IN NONFERROUS EAF PROCESSES Corrosion reactions in the infiltrated layer of alumina, high-magnesia, or magnesia-chrome based bricks in nonferrous EAFs and smelters are more complex due to numerous reactions with process slags of diverse compositions ranging from SiO 2-rich to basic slags with multiple secondary components including Cr2O3, MnO, P2O5, and PbO. Therefore, initially thermochemical simulations were evaluated with corrosion trials in an experimental scale induction furnace (IF). It became apparent from the results that whilst the calculated amount of stable oxidic liquid is an initial indication of corrosion processes in brick linings, the infiltrate viscosity also has to be considered to predict wear. Bricks with a highly viscous liquid phase in the infiltrated layer show a higher refractoriness in the IF tests than bricks with a low viscosity infiltrate. Therefore, the viscosity of the liquid infiltrate from the calculated chemical equilibrium composition was determined using the viscosity model from Decterov et al.,7 and Grundy et al.8 Calculated phase assemblages of an alumina-chrome brick and a magnesia-chrome brick partially infiltrated with an 18 wt.% Al2O3, 18 wt.% CaO, and 35 wt.% SiO2 slag from a primary nonferrous smelting process are depicted in Figure 8. In Figure 10 the calculated phase assemblages are shown for an alumina-chrome brick and a magnesia-chrome brick infiltrated with an olivine-type slag (20 wt.% MgO, 12 wt.% FeO, and 48 wt.% SiO2) from another nonferrous EAF smelter. The open porosity of these bricks was 15–17%. The simulated phase assemblages agree well with observations from postmortem analyses of infiltrated bricks (Figure 109). The corresponding viscosities of the infiltrated liquid and the corresponding results of the wear tests are given in Table III. The infiltrate of an alumina-chrome brick showed the highest viscosity of the brick series, even higher than the viscosity of the pure process slag oxidic liquid. For all other bricks considered for the customer furnace, the infiltrate viscosity was lower than the process slag. The observation that an alumina-chrome brick10 had the highest refractoriness is therefore explained by the increased viscosity of the infiltrated oxidic liquid. The viscosity depends on the chemical composition of the infiltrate, which differs significantly depending on whether the brick is alumina- or magnesia-based. From the calculated viscosity data in Table II it follows that superheating the lining results in a significant decrease of the infiltrate viscosity. In this case, superheating is deleterious in terms of the lining wear since mass transport will significantly increase in the infiltrated layer. Subsequently, the use of an alumina-chrome based lining led to very satisfactory results during the customer’s nonferrous process. From these simulations a preliminary list of feasible refractory materials can be selected for laboratory tests or plant trials, thereby decreasing the number and cost of laboratory investigations. 100%
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Fe-Cu liquid
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Slag to refractory mass ratio [%]
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10 12 14 16 18 20 22 24 26 28 30
Slag to refractory mass ratio [%]
Figure 8. Calculated stable phases of an alumina-chrome brick (A) versus a magnesia-chrome brick (B) partly infiltrated with alumina-calcium-silicate slag, T = 1400 °C 100%
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Slag to refractory mass ratio [%]
Figure 9. Calculated stable phases of an alumina-chrome brick (A) versus a magnesia-chrome brick (B) partly infiltrated with olivinetype silicate slag, T = 1650 °C
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B
A
Figure 10. Observed phases in an alumina-chrome brick (A) versus a magnesia-chrome brick (B) partly infiltrated with olivine-type silicate slag following induction furnace tests [9, 10], T = 1650 °C, sample images approximately 2 mm from hot side (alumina-chrome brick A: consisting of Cr-corundum (1) with Cr-rich rims (2), Cr-corundum in the matrix (3), Mg-Mn-Cr-Al spinel (4) and glassy phase (5) [9]; magnesia-chrome brick B: consisting of magnesia fine grain (1), forsterite infiltrate (2), chromite (3), (resin filled pore 4) [10]) Table III. Calculated viscosities of oxidic liquid in infiltrated brick linings and results from wear tests with olivine-type slag
Tested brick linings
High-magnesia Alumina-chrome Magnesia-chrome Customer olivine slag
Main components of oxidic liquid SiO2 [wt.%] 39 52 37 52
MgO [wt.%] 31 8 31 22
Al2O3 [wt.%] 9 24 8 9
CaO [wt.%] 21 3 19 5
Calculated viscosity [Pas] T = 1530 °C
T = 1650 °C
0.165 17.7 0.155 0.458
0.094 5.30 0.089 0.247
Observed results in IF test [10] Wear area Wear depth [cm2] [mm] Not used in IF tests 6.0–6.4 19.9–21.3 6.1–6.9 25.0–25.3
SUMMARY In metallurgical processes with highly variable cover slag chemical compositions, for example in the tundish, nonferrous EAF, and smelter processes, the optimum choice of refractory material is a challenging decision that often requires laboratory or plant trials. In order to better understand the corrosion processes in general and the results of particular laboratory tests and plant trials, simulations based on thermochemical modeling assuming local equilibrium in the infiltrated linings were performed. Results of the simulations concerning phase assemblage and infiltrate viscosity were interpreted in terms of the expected lining stability when in contact with a particular customer’s process slag. These simulations have helped to refine and decrease the number of lining materials chosen for laboratory tests and plant trials, therefore decreased investigation costs and supported well-founded lining recommendations can be made by RHI for individual customer processes. It has also been demonstrated that a significant proportion of phase assemblages are stable that result from the numerous combinations of customer slag compositions with the available lining materials. Therefore, a detailed analysis of customer slag chemistry is beneficial in order to minimize lining wear through the appropriate choice of lining material. Due to the large variability of slag chemistries in the tundish and in nonferrous processes this analysis is particularly relevant for tundish cover slags and nonferrous slags. RHI provides a large range of refractory materials to provide customer-tailored solutions depending on both the process conditions and slag composition and thermochemical modeling provides a tool to aid the appropriate selection. In addition, such analysis can also help improve lining performance in BOFs and EAFs although the slag chemical composition is less variable in these processes.
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REFERENCES [1] C.W. Bale, P. Chartrand, S.A. Decterov, G. Eriksson, K. Hack, R. Ben Mahfoud, J. Melançon, A.D. Pelton and S. Petersen, “FactSage Thermochemical Software and Databases,” Calphad, Vol. 26, No. 3, 2002, pp. 189-228. [2] S. Amini, M. Brungs, S. Jahanshahi and O. Ostrovski, “Effects of additives and temperature on the dissolution rate and diffusivity of MgO in CaO-Al2O3 slags under forced convection,” ISIJ International, Vol. 46, No. 11, 2006, pp. 1554-1559. [3] E. Gardés and W. Heinrich, “Growth of multilayered polycrystalline reaction rims in the MgO-SiO2 system, part II: modeling,” Contributions to Mineralogy and Petrology, 2011, DOI 10.1007/s00410-010-0581-4. [4] V. Reiter, F. Melcher, H. Harmuth and W. Netzer, “Thermochemical Modeling of EAF Slag and Refractory/Slag Equilibria,” Unitecr 2007, Dresden Germany, Unitecr Conference Proceedings, pp. 170-173. [5] W. Eckstein, R. Lanzenberger and R. Neuböck, “Phase calculations for ANKERHARTH mixes,” RHI internal report no. 20020277, Leoben Austria, 2002. [6] D. Gregurek, RHI internal report no. 20070068, Leoben Austria, 2007. [7] S. A. Decterov, A. N. Grundy, I.-H. Jung and A. D. Pelton, “Modeling the Viscosity of Aluminosilicate Melts”, In: Computation in Modern Science and Engineering, (Proc. Int. Conf. on Computational Methods in Sciences and Engineering), AIP (American Institute of Physics) Conference Proceedings, Vol. 963, Issue 2, Pt B, 2007, Eds. T. E. Simos and G. Maroulis, pp. 404-407. [8] A.N. Grundy, H.-C. Liu, I.-H. Jung, S.A. Decterov and A.D. Pelton, "A Model to Calculate the Viscosity of Silicate Melts. Part I.: Viscosity of Binary SiO2-MeOx Systems (Me = Na, K, Ca, Mg, Al)". Int. J. Mat. Res., Vol. 99, No. 11, 2008, pp. 1185-1194. etc. [9] A. Retschnig and D. Gregurek, RHI customer report no. CR20080059, Leoben Austria, 2008. [10] C. Maijcenovic and A. Retschnig, RHI customer report no. CR20090140, Leoben Austria, 2009.
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