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Metallurgia Italiana

International Journal of the Italian Association for Metallurgy

n. 11/12 Novembre-Dicembre 2018 Organo ufficiale dellâ&#x20AC;&#x2122;Associazione Italiana di Metallurgia. Rivista fondata nel 1909

GIORNATE NAZIONALI SULLA

xiii edizione

PALERMO - 3-5 LUGLIO 2019

PRESENTAZIONE

LE EDIZIONI

La XIII edizione delle Giornate Nazionali sulla Corrosione e Protezione si terrà a Palermo dal 3 al 5 luglio 2019, presso il Campus dell’Università degli Studi di Palermo. Le Giornate rappresentano l’evento di riferimento a livello nazionale per la discussione ed il confronto sulle questioni scientifiche, tecnologiche e produttive, nell’ambito della corrosione e protezione dei materiali. In particolare, il Convegno prevede la presentazione dei risultati raggiunti da vari gruppi di studio e da numerose aziende del settore, in forma orale e poster. Il Convegno si aprirà nel pomeriggio di mercoledì 3 luglio con la Cerimonia d’apertura e due plenary lectures, a cui seguiranno nei due giorni seguenti le sessioni tecniche. La Cena del Convegno sarà offerta la sera di venerdì 5 luglio.

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presentazione di memorie

Gli interessati a presentare memorie scientifiche dovranno inviare entro il 31 gennaio 2019, il titolo della memoria, i nomi degli autori con relative affiliazioni ed un breve riassunto. Le memorie potranno essere proposte: - compilando il form online presente sul sito dell’evento: www.aimnet.it/gncorrosione - inviando il riassunto e tutte le informazioni richieste a mezzo e-mail: info@aimnet.it

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La Metallurgia Italiana

Metallurgia Italiana

International Journal of the Italian Association for Metallurgy

n. 11/12 Novembre Dicembre 2018 Organo ufficiale dell’Associazione Italiana di Metallurgia. Rivista fondata nel 1909

International Journal of the Italian Association for Metallurgy Organo ufficiale dell’Associazione Italiana di Metallurgia. House organ of AIM Italian Association for Metallurgy. Rivista fondata nel 1909

n. 11-12 Novembre-Dicembre 2018 Direttore responsabile/Chief editor: Mario Cusolito Direttore vicario/Deputy director: Gianangelo Camona Comitato scientifico/Editorial panel: Livio Battezzati, Christian Bernhard, Massimiliano Bestetti, Wolfgang Bleck, Franco Bonollo, Bruno Buchmayr, Enrique Mariano Castrodeza, Emanuela Cerri, Lorella Ceschini, Mario Conserva, Vladislav Deev, Augusto Di Gianfrancesco, Bernd Kleimt, Carlo Mapelli, Jean Denis Mithieux, Marco Ormellese, Massimo Pellizzari, Giorgio Poli, Pedro Dolabella Portella, Barbara Previtali, Evgeny S. Prusov, Emilio Ramous, Roberto Roberti, Dieter Senk, Du Sichen, Karl-Hermann Tacke, Stefano Trasatti Segreteria di redazione/Editorial secretary: Valeria Scarano Comitato di redazione/Editorial committee: Federica Bassani, Gianangelo Camona, Mario Cusolito, Carlo Mapelli, Federico Mazzolari, Valeria Scarano Direzione e redazione/Editorial and executive office: AIM - Via F. Turati 8 - 20121 Milano tel. 02 76 02 11 32 - fax 02 76 02 05 51 met@aimnet.it - www.aimnet.it

Anno 110 - ISSN 0026-0843

Produzione d'acciaio con convertitori ad ossigeno / Oxygen steelmaking A new approach for modelling and control of dephosphorization in BOF converter S. Khadhraoui, H.-J. Odenthal, S. Das, M. Schlautmann, K. Hack, B. Glaser, R. Woolf

Conventional steelmaking route being subject to transformation P.Argenta, M. Marcozzi, M. Dorndorf, P. Duarte, S. Maggiolino, O. Demir 17 Estimation of trace elements in unalloyed scrap R. Ammer, P. Reisinger, M. Egger, R. Tober, G. Salzmann

Attualità industriale / Industry news Manifestazioni AIM

Dissolution behaviour of Ulc steel in carbon saturated hot metal edited by: F. Markus Penz, J. Schenk, R. Ammer, K. Pastucha, B. Maunz 36 Effect of converter dart design aspects on its settling behaviour edited by: J. van Oord, H.J. Visser 46

siderweb LA COMMUNITY DELL’ACCIAIO

Gestione editoriale e pubblicità Publisher and marketing office: Siderweb spa Via Don Milani, 5 - 25020 Flero (BS) tel. 030 25 400 06 - fax 030 25 400 41 commerciale@siderweb.com - www.siderweb.com La riproduzione degli articoli e delle illustrazioni è permessa solo citando la fonte e previa autorizzazione della Direzione della rivista. Reproduction in whole or in part of articles and images is permitted only upon receipt of required permission and provided that the source is cited. Reg. Trib. Milano n. 499 del 18/9/1948. Sped. in abb. Post. - D.L.353/2003 (conv. L. 27/02/2004 n. 46) art. 1, comma 1, DCB UD Siderweb spa è iscritta al Roc con il num. 26116 Stampa/Printed by: Poligrafiche San Marco sas - Cormòns (GO)

Scenari / Experts' Corner Reaction near hotspot in refining of molten iron during supplying top-blown oxygen and powder additives edited by: Yu-ichi UCHIDA 58 Development and application of high efficiency blowing technology of converters in shougang edited by: Pan GAO, Xinhua WANG, Haibo LI, Yuming GUO, Yong ZHANG, Xiaodong ZHAO, Huixiang YU 67 Atti e notizie / Aim news Calendario eventi internazionali

Rubrica dai Centri

In ricordo di Gianfranco Banzato - Acciaierie Venete CleanTech4

78 79

Giornata di Studio “Ottimizzazione dei processi termochimici e meccanici nell’industria meccanica” 81 Indice analitico Indice autori

82 89

l’editoriale La Metallurgia Italiana

Ing. Federico Mazzolari Presidente AIM

E’ dicembre anche per AIM E’ tempo di bilanci: quello di AIM è positivo, grazie all’impegno dei nostri Centri di Studio che hanno saputo organizzare Corsi e Giornate di Studio dedicate per un totale di 29 eventi ed alla struttura di AIM che si è prodigata per realizzare al meglio i 4 Convegni. L’attività di AIM per il 2019 sarà rivolta prevalentemente ad iniziative di carattere nazionale con oltre 20 manifestazioni già messe a programma, essendoci stata destinata, al momento, quale sola iniziativa internazionale la European Conference Heat Treatment & Surface Engineering che avrà luogo a Bardolino in giugno. L’unica nota non positiva è il lieve calo del numero dei soci che ammonta a circa 1.500 unità; su questo punto specifico stiamo già lavorando perché la nostra missione è quella di diffondere al meglio la cultura metallurgica. Anche per il settore della metallurgia nel complesso l’anno è stato positivo, dazi americani permettendo. La produzione ed i consumi si sono mostrati robusti. In Italia la produzione di acciai si è riportata ad un valore di 25 milioni di tonnellate, mentre per altri metalli come l’alluminio ed il rame e sue leghe dobbiamo parlare solo di quantità trasformate poiché la produzione primaria è oramai assente dall’Italia da alcuni anni; ciò non toglie che tramite il riciclo dei rottami e le importazioni del metallo primario i produttori italiani arrivino a trasformare circa 2 milioni di ton di alluminio e 1 milione di ton di rame e sue leghe, e questa non è che una delle tante eccellenze degli imprenditori metallurgici nazionali. Tornando all’acciaio, i consumi italiani si sono mostrati in tensione a circa 28 milioni di ton con importazioni nette per altri 3 milioni di ton, e ciò anche per la mancata produzione dello stabilimento di Taranto che vale almeno 4 milioni di ton. Sull’onda di una fase espansiva dell’economica mondiale l’Italia è vissuta al traino dei mercati europei grazie alla capacità di cogliere le opportunità offerte e alla competitività, tipica espressione dell’industria italiana. E’ più che doveroso un ringraziamento da parte mia ai nostri soci, la vera ricchezza della nostra associazione, e a tutti coloro che hanno contribuito alle attività di AIM, con la pubblicazione della nostra rivista e alle svariate attività sia di studio sia di convegno. Un augurio ambizioso (non vorrei osare troppo): che il 2019 possa trasformare i desideri in realtà.

La Metallurgia Italiana - n. 11/12 2018

Produzione d'acciaio con convertitori ad ossigeno

A new approach for modelling and control of dephosphorization in BOF converter S. Khadhraoui, H.-J. Odenthal, S. Das, M. Schlautmann, K. Hack, B. Glaser, R. Woolf

In the BOF process, the accurate control of Phosphorous removal up to ultra-low values is very important to ensure a high quality of the end–product. This has proven to be associated with several challenges, leading to a preferred usage of high priced low P iron ores. Thus, in the scope of the European BOFdePhos project, important thermodynamic and kinetic aspects of the dephosphorisation reaction such as the effect of solid phases on Phosphorous distribution and lime dissolution in a foamy slag were investigated. It was found that BOF slags are heterogeneous during a large period of the blow and also at the end of blow in most cases. The type and amount of solid phases is strongly affected by temperature and minor oxides content such as MgO-, MnO- and Al2O3-content. The consideration of solid phases formation in the slag, especially the P dissolving C2S_C3P phase, is crucial for a successful modelling and control of dephosphorisation. However, most of the Phosphorous distribution equations available in the literature were developed for homogeneous slags. Thus, a new approach for thermodynamic modelling of the P-distribution between a heterogeneous slag and liquid iron covering the total blowing period in converter was developed and incorporated in a kinetic dephosphorisation model. It was found that while the P distribution in a fully liquid slag was a strong function of temperature, CaO- and FeO-content, the P distribution in a heterogeneous slag depended further on temperature but also on the amount of solid phase as well as the basicity of the liquid slag phase. Even though the C2S_C3P phase can dissolve high amount of Phosphorous, the P-dissolution in the solid phase in the industrial process is associated with strong kinetic limitations. New strategies for the enhancement of dephosphorization control in the BOF converter, focusing on using the potential of the solid phase in removing Phosphorous, were developed.

KEYWORDS: BOF CONVERTER – DEPHOSPHORISATION – PHOSPHROUS DISTRIBUTION – C2S_C3P SOLID PHASE – HETEROGENEOUS SLAG – PHASE DIAGRAMS – FOAM – LIME DISSOLUTION– FACTSAGE – JET VELOCITY INTRODUCTION The achievement of ultra-low-P values in steel during oxygen converting of hot metal is crucial for high quality steel production. Lowering the Phosphorous content has become a critical requirement for steels used in thin sheets for deep drawn applications, automobile exteriors and pipelines for transportation of natural gas and petroleum products [1]. The oxygen converter process provides excellent conditions for phosphorous removal from steel. The dominance of a high oxygen potential in the metal-slag interface throughout the blow and the presence of a basic slag cause favored thermodynamic conditions for P removal from the metal phase. Moreover, the high mixing intensity induced by top and/or bottom gas injection at high speed as well as the CO/CO2 gas generation resulting from the decarburization reaction further contribute to the enhancement of the dephosphorization kinetics. However, the challenges are high for steelworkers especially those with a large variation in hot metal composition and subsequently also in the final slag composition. The non-achievement of final P contents at the end of the oxygen blowing requires a reblowing, that is continuation of oxygen blow instead of the scheduled tapping under addition of fluxing agents, which result in production time loss, decrease in iron yield, excessive lime consumption and increased La Metallurgia Italiana - n. 11/12 2018

refractory wear. Within the scope of the European BOFdePhos project, a comprehensive study of the dephosphorization reaction has been carried out with focus on the thermodynamic

Sabrine Khadhraoui, Hans-Jürgen Odenthal SMS group GmbH, Germany

Satyajit Das

SMS India Pvt. Ltd, India

Martin Schlautmann

VDEh-Betriebsforschungsinstitut GmbH, Germany

Klaus Hack

GTT Technologies, Germany

Björn Glaser

Royal Institute of Technology, Sweden

Rhian Woolf

Tata Steel Europe, UK

Oxygen steelmaking modelling of the state of metal-slag-equilibrium and consideration of metal droplets circulating through a metal-slag emulsion for modelling of transport kinetics. In this paper, the main results with respect to the investigations of thermodynamics and kinetics of dephosphorization will be presented. Evaluation of the state of BOF slags The achievement of the aim steel composition at end of the blowing process, especially with respect to P and C-content, depends strongly on controlling the slag evolution. A nonsuccessful slag control may lead to slopping, that is slag overrunning from converter mouth with considerable economic withdraws for steelworks. Even though a lot of effort has been made to investigate the thermodynamic state of BOF slags since the 1930s, their behavior under various process conditions is still not completely understood. This is due to their complex structure, combined with the fact that the slags undergo large variations during the process in terms of composition and temperature, mainly as a result of the simultaneous oxidation and reduction processes during the blow. Also, the slag is foamy during a large period of the blow, as a result of CO and CO2 gas evolution which are the products of the main refining reaction, decarburization. Thus, the state of converter slags as a function of temperature, oxygen partial pressure and minor oxides addition was evaluated within the current project. Some of the main results will be briefly introduced in the current section whereas

a detailed description can be found in [2]. Effect of various process conditions In order to investigate the thermodynamic behavior of slags in equilibrium with metallic phase under various process conditions, a thermodynamic database for the system CaOFeOx-P2O5-SiO2-MnOx-MgO-Al2O3 was developed by GTT. The Gibbs energy of the liquid phase was modelled by the nonideal associate solution model [3] and the Gibbs energy of non-stoichiometric solid compounds based on the Compound Energy Formalism [4]. Due to the importance of dephosphorization reaction, special attention was given to modelling the P-containing phases. The database was critically assessed using all experimental data available, mostly for binary and ternary subsystems [2, 5]. The effect of temperature, oxygen partial pressure and composition was investigated by means of coupling the database to the thermodynamic software package Factsage TM [6]. A. Effect of temperature The effect of temperature on the ternary system CaO-FeO-SiO2 in equilibrium with liquid iron is illustrated in Fig.1 by means of the isothermals presentation at 1550 °C (Fig.1-a) and 1700°C (Fig.1-b) for the range relevant to the BOF converter slags (low SiO2 range).

Fig. 1 – Liquidus isothermals of the system CaO-FeOx-SiO2 in equilibrium with liquid Fe a) 1550°C b) 1700°C It can be seen that the system is dominated by solid phases precipitation regions: The dominant region is zone 1, which is the saturation region on the 2.CaO.SiO2 solid phase, denoted as C2S, followed by zone 2, the double saturation region on both 2.CaO.SiO2 and 3CaO.SiO2, denoted as C3S. The double saturation region on both C3S and the CaO-based monoxide solution is marked as zone 3. It can be seen that the lime only saturation region (zone 4), usually considered as the target zone for optimal dephosphorization, has a small stability range in the main oxide system compared to zone 1 and zone 3. For example, at a temperature of 1700°C, the lime saturation only region (zone 4) can be reached only when the SiO2 contents of the total slag is less than 10 wt. % and this value decreases with a decrease 6

in temperature. Also, the liquid slag only region (zone S) can be reached only at very high (%FeO)-contents in the total slag: At a temperature of 1700°C, (%FeO) should be higher than 40% to maintain a purely liquid slag, and this value is higher for lower temperatures. Thus, converter slags are expected to be heterogeneous at the end of blowing process, where the typical temperatures are in the range of 1600-1700°C and saturated either on C2S or double saturated on both C3S phase and the CaO-based monoxide solution. B. Effect of MgO and Al2O3 The effect of MgO- and Al2O3-addition on the main slag system CaO-FeOx-SiO2 is illustrated in Fig. 2-a and Fig. 2-b respectively. La Metallurgia Italiana - n. 11/12 2018

Produzione d'acciaio con convertitori ad ossigeno The formation of C3S-phase was suppressed for the sake of enhanced readability of the diagrams. Thus, the C2S phase was considered representative of the silicate phases of the system. The addition of Al2O3 or MgO resulted in the shrinkage of the C2S-saturation region (marked blue in Fig. 2 and corresponds to zone 1 in Fig. 1) in favor of expansion of both liquid slag and lime only saturation regions (zones S and 4 in Fig. 1), with the effect of Al2O3 being stronger than that of MgO in this aspect. One major difference in the effect of both oxides is however

the massive change below lime saturation line (for both high and low FeO systems) in case of MgO addition: The saturation regions containing both silicates and monoxide solid solutions (zones 3 and 4 in Fig.1) expanded on the cost of the C2S-only saturation region (marked blue in Fig. 2 and corresponds to zone 1 in Fig. 1). For an MgO-content equal or higher than 7 wt.%, a miscibility gap was observed in the monoxide phase leading to formation of a CaO- and MgO-based monoxide solutions.

Fig. 2 â&#x20AC;&#x201C; Effect of a) MgO b) Al2O3 addition at 4% (top), 7% (middle) and 10% (bottom) to the CaO-FeOx-SiO2 system in equilibrium with liquid Fe at 1600Â°C: The dashed lines represent the corresponding isothermal in the original CaO-FeOx-SiO2 system In conclusion it can be said that the addition of Al2O3 and MgO results in more liquid slag formation and less silicates whereas MgO has the additional effect of increasing the amount of monoxide precipitations: For MgO values in the range of 7 % or higher, a lime-saturated slag became saturated on two monoxide phases: an MgO-based and a CaO-based solid solutions.

of more liquid slag compared to the low p(O2) case (solid lines). A small shift in the lime saturation line towards higher CaOcontents was observed, marking an increase in lime solubility of the slag. The effect of p(O2) became less significant at higher temperatures, especially for the silicates saturation zones, as can be seen from Fig. 3-b.

C. Effect of Oxygen partial pressure An increase in the Oxygen partial pressure (dashed line in Fig. 3), further noted p(O2), resulted in a small shrinkage of the silicates saturation regions (zone 1, 2 and 3 in Fig. 1) in favor

La Metallurgia Italiana - n. 11/12 2018

Oxygen steelmaking

Fig. 3 – Effect of p(O2) on the CaO-FeOx-SiO2 system: The solid lines represent the case of low p(O2) (equilibrium with liquid Fe) whereas the dashed lines represent the case of a high p(O2) = 10-5 bar, a) 1550°C b) 1750°C D. Conclusion Steelmaking slags are heterogeneous for a large period of the blowing process as well as at the end of blow as long as the final temperatures are equal or less than 1700°C. For low-MgO systems, the final slag is expected to be saturated on C2S or double saturated on C3S and CaO-based monoxide phase, while for the MgO-containing systems, the slag is expected to be either saturated on MgO-based monoxide phase or double saturated on MgO- and CaO- based monoxides. In the industrial process, MgO, Al2O3, MnO and further oxides co-exist in the slag and the resulting effect on slag may be different than in the case of separate addition. For example, while it was found that the addition of MnO does have a considerable effect on the main CaO-FeOx-SiO2 system, it was found that in combination with MgO, MnO stabilizes the monoxide-phase which is formed in the lime saturated part of the system due to its dissolution in both CaO- and MgO-based monoxide phases. As a result, a reduction in the C2S-saturation zone accompanied by an expansion in monoxide saturation zones was observed. Thus, the consideration of minor oxides presence and their combinations is crucial for a proper assessment of the slag state.

Effect of foaming Slag foaming leads to an increase in the droplets residence time compared to a non-foaming slag [7] which would eventually increase their refining rate, especially with respect to the decarburization reaction [8]. Further experimental investigations of foaming slags within the current project indicated that the foam had considerable effect on the slag structure and the lime dissolution rate. It was found that the gas bubbles are not distributed uniformly in the slag, but rather form a 2-phasemixture with layers dominated by foam where no liquid film was separating the gas bubbles and bulk slag layers consisting of pure liquid slags [9]. In addition, lime dissolution experiments in foamy slags (see Fig. 4) carried out by KTH showed that the foam slows down the lime dissolution rate by a factor of 0.5 to 0.7, compared to the case of a pure liquid slag. This shows that the contact area between slag and lime is very important and that the presence of the gas phase disturbs lime dissolution. The observed dissolution mechanism was the same for both cases: After reacting with the slag, the lime is peeled off at the edges of the cube with related reduction of the cube side length.

Fig. 4 – Dissolution rate of a lime cube from the Port Talbot Plant at 1600 °C: a) liquid slag b) foamy slag 8

La Metallurgia Italiana - n. 11/12 2018

Produzione d'acciaio con convertitori ad ossigeno Consequences for slag sampling in BOF converter As described in the previous sections, BOF slags have a complex morphology due to the presence of solid phases and gas bubbles. Further contributions to non-homogeneity of the slag are temperature gradients in the slag. It is also possible that the gas dominated parts of a foamy slag would have a lower p(O2) than the bulk liquid parts since the gas mainly contains CO. The determination of slag analysis at the end of the blowing process is a common practice for steelworks, as it delivers information about the state of steel-slag equilibrium and, if correlated with the blowing strategy, charge materials input and steel analysis, can be used for optimizing the process especially in terms of enhanced controllability. However, BOF slag sampling is generality carried out by means of immersion of a steel bar into the slag. After taking out, a layer of slag is sticking on the bar which mainly originates from the fluid phase of slag and thus may only represent the liquid slag and not necessarily the total slag based on the current findings. In this context, Fig. 5 shows the comparison between the simulation results of a recently developed dynamic model by Rout and co-workers [10] and industrial measurements of Cicutti et al.

[11]. It can be seen that the simulated FeO-content of the slag increased during the last 2 minutes of the blow, as implied by the oxygen balance due to a decrease in C removal rate. However, the measured FeO-content of the slag remained more or less constant during this period. Considering the main oxide system CaO-FeO-SiO2, it is possible that the slag reached a double saturation zone (zone 2 or zone 3 in Fig. 1) during this period. In such a situation, a change in slag composition while remaining within the same zone would modify the phases amount in the slag but not the phases composition. A further indication is given by the SiO2 evolution in this period, as the measured value remained constant and did not decrease despite the considerable increase in the total slag mass. Thus, the measured slag evolution, especially during the last 2 minutes, may correspond to the liquid slag composition only, while the total slag contained solid phases (silicates and/or monoxides). It is suggested that during the last 2 minutes of blow, the slag was situated in zone 2 or zone 3 of Fig. 1, with zone 3 being more probable since the final slags in the industrial process are generally lime saturated.

Fig. 5 â&#x20AC;&#x201C; Comparison between the measured slag evolution of Cicutti et al. [11] (dots) and simulation results of Rout et al. [10] (solid and dashed lines) for two post combustion profiles (blue: profile 1 , red: profile 2) : a) FeO b) SiO2 Also, even though it is well-established that the Phosphorus distribution between liquid steel and slag, further noted as Lp, is strongly correlated with the Oxygen potential and thus with the FeO-content of the slag (assuming an FeO-control of the Oxygen potential), the industrial experience showed that the measured Lp-values exhibit a large scatter when represented as a function of the FeO-analysis and slag basicity. This is shown in Fig. 6-a on the example of a large number of heats taken from the Port Talbot Plant in year 2013, which represents the challenges associated with taking representative slag samples.

La Metallurgia Italiana - n. 11/12 2018

Interestingly, the correlation between the measured values of [%P]- and [%Mn]-content of the bath, shown in Fig. 6-b is very pronounced. This can be explained by the fact that the Mn-equilibrium between metal and slag phase was governed by Fe/FexO oxygen potential [12]. As a result, [%Mn]-content of the bath was strongly correlated with (%FeO)-content in slag and subsequently with Lp and [%P]-content of the bath. Due to the homogeneity of the steel bath, correlations based on steel analysis were found to be more reliable than those based on slag analysis.

Oxygen steelmaking

Fig. 6 – Metal-slag-relations based on the analysis of some heats from the TATA Steel’s Plant in Port Talbot (UK) (2200 data from year 2013) a) Lp-FeO-basicity relation at a temperature of 1680°C; b) [%Mn]-[%P]-temperature (blue for low Temp. and red for high Temp.) with [%P] averaged for each Manganese value: the average [%P] had a standard deviation σ lower than 40 ppm for 90 % of data Based on those observations, methods to predict and control the end-Phosphorous content in the steel bath based on monitoring the Manganese content of the bath were developed [13].

Study of P distribution between heterogeneous slag and liquid metal phase As discussed in the previous section, BOF converter slags are expected to be saturated on silicates for a large period of the blow. Thus, the investigation of the effect of the presence of solid phases and particularly, the effect of silicates on Lp is important for a proper assessment of the equilibrium state of Phosphorous in the industrial process. The simulation results based on the thermodynamic BOFDePhos database revealed

that the presence of P2O5 stabilized the C2S phase over the C3S phase. P2O5 dissolved in the solid solution under formation of the 2CaO.SiO2-3CaO.P2O5 solid phase, further denoted as C2S_C3P, by substituting some of the silicates with phosphates [2]. As a result, P2O5 can be dissolved in the C2S_C3P phase up to its stoichiometric solubility limits of 22 wt. %. Fig. 7 shows the simulation results for Lp between a C2S-saturated slag containing 2 wt. % P2O5 and liquid Iron a function of the mass ratio between C2S_C3P and liquid slag, further denoted as W. It can be seen that Lp-values in the range of 10 3 and higher were calculated when W was equal or higher than 0.3. In all cases, the Lp-values were at least by 2 orders of magnitude higher for slags with a high basicity, which denotes the CaO/SiO2 weight ratio in the liquid slag part.

Fig. 7 – Results of the thermodynamic simulations for C2S_C3P saturated slags with 2 % P2O5 at a temperature range of 15501600°C and comparison with the experimental and calculation results of Suito and Inoue [14] at 1560°C, [%P]= 0.5 wt.% and slag/ metal ratio of 1/10 10

La Metallurgia Italiana - n. 11/12 2018

Produzione d'acciaio con convertitori ad ossigeno In a series of experiments, Inoue and Suito [14-16] investigated the mechanism of dephosphorization as well as the Phosphorous distribution in C2S_C3P containing slags. The authors reported that by adding C2S particles to liquid slags, Lp increased drastically compared to the case of a fully liquid slag, due to formation of C2S_C3P phase. Some of their results with respect to Lp measurements and calculations (based on their own developed approach) are presented in Fig.7. The results show that the calculated Lp-values by Suito and Inoue [14] were in the same range as those resulting from the simulations for the high basicity case exceeding 104 when W was higher than 0.1. The experimental behavior was similar to the simulated behavior with Lp increasing with an increase in W. Even though the experimental values were generally below the simulation results with a maximum of 103 reached at a W value of 0.45, they are still considerably higher than in the case of fully liquid slags (case W = 0). The results confirm that the C2S_C3P phase had a strong potential on dissolving P but also suggest that this is associated with kinetic limitations and that a W value exists, above which no further Lp increase occurs. Interestingly, the

authors observed that the Lp of a slag which had an W of 0.2 increased by 70% when stirring was used (see Fig. 7). This is an indication of a possible blocking layer removal phenomena, probably similar to that observed during lime dissolution [17]. In a recent work, Gu et al. [18] observed that the overall mass transfer coefficient for dephosphorization in a liquid slag increased by a 2 order of magnitude in the presence of Carbon due to the stirring effect provided by the formation. It is possible that the stirring effect of CO gas may additionally contribute to Lp enhancement in a heterogeneous slag in a similar manner to the increase observed in the experiments of Suito and Inoue. At the same time, the presence of CO gas may decrease the interface between C2S_C3P and liquid slag and thus negatively contribute to dephosphorization. Thus, additional experimental studies are required to clarify the behavior of this phase under foaming conditions. Finally, the following Lp-equation was developed by regression analysis of the simulation results for C2S_C3P containing slags in the temperature range of 1400-1750°C with a regression coefficient of 0.922: (1)

It should be noted that for temperatures below 1550°C, the formation of solid Fe phases was suppressed in the simulations and thus an extrapolation to the case where liquid Fe would be stable at such temperatures was made. In the industrial process, the presence of C lowers the melting point of the metal and thus, the metal phase remains liquid throughout the blow even when the starting temperature is in the range of 1300°C. In the simulations, the addition of C to the liquid phase would lead to FeO-reduction and thus, to a modification of the initial slag composition. This is also the reason behind the non-availability of experimentally established Lp relations below 1550°C : A reduction of a high FeO-slag with Carbon contained in the metal phase and in some cases, also with Carbon contained in the gas phase was observed [19]. Analysis of industrial findings The industrial analysis of the morphology of BOF slags and the state of P dissolution analyzed by Preßlinger et al. [20] from the Voestalpine plant and DeO et al. [21] from several European and Indian steel plants confirmed that the final slags were C2S-saturated and that a large part of P2O5 was present in the solid phase. However, the maximum dissolved value observed was in the range of 5 wt. %, which is much lower than its solubility limit. In general, the Lp values observed in the industrial process are much lower than those predicted by the thermodynamic simulations and experimental measurements in C2S-saturated slags (presented in Fig. 7). This is possibly due to low P2O5 contents in industrial slags combined with large C2S precipitation amounts, which cause kinetic limitations to the La Metallurgia Italiana - n. 11/12 2018

enrichment process. Even though industrial observations [20, 21] confirmed that the C2S_C3P phase made a considerable contribution to P removal, little attention was given to the incorporation of the role of this phase in P removal or prediction strategies in conventional oxygen steelmaking. Apart from some few works, such as the work of DeO and al. [21], little effort has been made to incorporate the C2S-formation and dissolution in the derivation of industrial based Lp relations or P control strategies. It should be noted further that even though the effect of minor oxides on Phosphorous distribution was subject to many experimental and industrial investigations, it has rarely been connected to their effect on the C2S-phase formation and dissolution. Based on the description of the minor oxides effect on the main system CaO-FeO-SiO2, it is possible that the controversial effect of minor oxides such as MgO, Al2O3 and MnO is associated with their effect on solid phases formation and dissolution. For example, MgO was found to have a positive effect on Lp in laboratory based studies, but some steel plants reported a strong negative effect [22]. It is possible that the positive effect of MgO on Phosphorous distribution in the liquid slag would be counteracted by the decreasing effect on the precipitation amount of the P dissolving C2S_C3P phase (see Fig. 2) which results in an overall negative effect. In addition, MgO may lead to high precipitations of the monoxide phase which does not dissolve Phosphorous. As a result, the slag viscosity increases which may contribute to kinetic limitations in the P enrichment in both liquid and C2S phase. Similarly, Al2O3 addition would cause less C2S formation in fa11

Oxygen steelmaking vor of more liquid slag (see Fig. 2) and subsequently, decrease the P distribution in a C2S-saturated slag. DeO et al. [21] observed that the presence of Al2O3 decreased both the C2S amount and its P content. It was reported at TATA Steelâ&#x20AC;&#x2122;s Plant in IJmuiden (the Netherlands) [23, 24] that Al2O3-contents above 3.5 wt. % had a negative effect on dephosphorization. Recently, Drain et al [25] assessed more than 90 Lp-equations based on industrial data and found out that Al2O3 had a weak negative effect, except at low Oxygen potential ranges. This can be explained by the fact that at low FeO contents in the slag, excessive C2S precipitations would occur and the optimal C2S/ liquid ratio, de-noted as W, would be exceeded. In this case the increasing effect of Al2O3 on the liquid slag amount may have a positive effect on Lp in a C2S-saturated slag. Considered dephosphorization modelling approach

The results of experiments carried out by Ito and Sato [26] and later by Inoue and Suito [15] showed that Lpslag depended mainly on the liquid slag composition and had maximum values

Thermodynamic approach Based on the findings introduced in the previous sections, it can be stated that the description of P distribution between a heterogeneous slag and liquid iron should consider the amount, type and composition of the slag phases. However, most Lp-approaches in the literature were developed for fully liquid slags. At the same time, the direct application of the Lp relation given by (1) would lead to an overestimation of Lp since no kinetic limitations on the P enrichment in C2S_C3P are considered in the thermodynamic simulations. Based on the approach of Inoue and Suito [14, 15], the term Lpslag, describing the P-distribution between the C2S_C3P phase and the liquid slag phase, was introduced:

(2)

at the nose of C2S_saturation region (zone 1 in Fig. 1). The following relation was developed within the presented work to fit their experimental findings: (3)

Finally, the P distribution between a heterogeneous slag containing C2S_C3P and monoxide solid phases and liquid iron can be described by the following equation, where W = C2S/ slagliq is the ratio of the C2S_C3P and liquid slag phase frac-

tions, W<lim is this ratio with a maximum value of Wlim and fmo, denotes the fraction of the monoxide phase (in case of noxide miscibility gap the sum of both phases is considered) in the heterogeneous slag: (4)

The term Lp liq corresponds to the P distribution between the liquid slag and the liquid metal, which was subject to intensive investigations in the literature. Simulation results by of the thermodynamic BOFdePhos database showed that Lpliq was a stong function of temperature, CaO- and FeO-content. The limiting value Wlim was introduced into the equation to limit the P transfer from liquid slag to the solid phase when the C2S/ slagliq ratio exceeds a certain value, which was found to lie in the range of 0.3-0.45 in the experiments of Suito and Inoue [14]. For application of the equation for Lp tot, the total amount of phases and their composition has to be determined accura-

tely and thus, the thermodynamic database was used for this purpose. Finally it should be noted that even though the description of Lp tot does not explicitly include the effect of minor oxides additions, their effect is indirectly incorporated through the consideration of phases amounts and compositions. For example, if MgO is present and the slag is lime saturated, both liquid slag and C2S phase amounts will decrease in favor of increased monoxide phase formation and lower Lptot values compared to the MgO-free case will be calculated. The effect of minor oxides on the P distribution between the liquid slag and the metal phase, Lpliq, was not incorporated so far in the La Metallurgia Italiana - n. 11/12 2018

Produzione d'acciaio con convertitori ad ossigeno thermodynamic database, mainly due to lack of experimental or industrial data covering the effect of minor oxides additions on the main system CaO-FeOx-SiO2 over a wide range of liquid slag compositions and temperatures. Kinetic approach It is a well-established fact that the dephosphorization reaction occurs only as a metal-slag reaction and thus it takes place in the so called emulsion zone in the converter [27], where the ejected metal droplets from the steel bath react with the slag and entrapped gas bubbles. Thus, a 2-zone-modelling approach, similar to that proposed by Schoop et al. [27], was adapted for modelling the dephosphorization reaction, based on the following assumptions: ◦ The droplets ejected from the metal bath to the emulsion zone attain equilibrium with respect to the dephosphorization reaction within their residence time in the emulsion zone. This is based on the fact that high residence time values were reported in a previous work [5]. ◦ The droplets returning to the bath from the emulsion zone

have achieved equilibrium. Thus the extent of P removal in the bath depends further strongly on the circulation rate of droplets between bath and emulsion zone. The circulation rate was calculated based on the blowing number approach suggested by Subagyo et al[28]. However, for the determination of the Oxygen jet impact velocity as a function of lance height, Oxygen flow rate and furnace conditions, a CFD approach was incorporated. Fig. 8 shows the simulation results of the jet axial velocity behavior for a cold and hot ambient temperature and the comparison with corresponding results based on several approaches reported in the literature. It has been found that, in accordance with the results of Alam et al.[29], the approach of Sumi et al. [30] provided the highest accuracy at a hot furnace temperature when the dimensionless axial distance lay below 60 in the BOF converter. At a cold ambient temperature of 20°C however, which is usually the temperature set in most experiments investigating droplets generation and residence time, the equation proposed by Koria et al. [31] followed by that of He and Standish [32] provided the highest accuracies for dimensionless axial distances higher than 15.

Fig. 8 – LES (Large Eddy Simulation) simulation results of jet axial velocity in cold and hot environment a) T∞ =20°C, b) T∞ =1600°C with x: axial coordinate or lance height, de: nozzle exit diameter Strategies for enhanced P control Based on the current findings, the maximum P distribution values in the BOF process are not achieved at lime saturation, but rather at C2S_C3P saturation. The presence of this phase even in small amounts lower than 10 wt. %, would lead to the achievement of ultra-low P values in the metal phase if the enrichment is high. Thus, the focus is on the development of new process control strategies and new technologies to make use of the potential of this phase in removing Phosphorous and overcome the kinetic limitations for its enrichment [33, 34]. An example is presented in Fig. 9. The method consists of blowing powdered lime and, to a lesser extent, also powdered dolomite into the slag (or emulsion zone) through side wall nozzles. It was found that through this technology, the lime consumption can be reduced drastically while high Lp values are achieved as a result of the P dissolution in C2S_C3P. Also, in case powdered dolomite is used, the resulting slags would be high in basicity La Metallurgia Italiana - n. 11/12 2018

and less aggressive to the lining compared to the case where lump dolomite is used as observed in the experiments of Münchberg et al. [35]. By optimizing the lance positions and their blowing angles, the mixing intensity in the slag can be enhanced with minimal carrier gas consumption and thus, the blocking layer formation for both lime dissolution and C2S_ C3P enrichment is avoided. Also, an enhanced dephosphorization at low FeO contents and slag amounts would be possible. The cost-effectiveness of the method lies in the possibility of refining high P hot metal in a single step, decreased lime consumption and refractory wear, increased iron yield as well as in the suppression of slopping and skulls formation. In addition, the new P modelling approach has been implemented into the BOF Design model of SMS used for process design and for supporting the online model [36] as well as into the dynamic online model of BFI [37]. Fig. 10 sketches the implementation scheme with contributions from the thermodynamic 13

Oxygen steelmaking BOFdePhos database regarding amounts and compositions of the involved slag phases, the enhanced Lp model for heterogeneous slags and the transport kinetics based on droplets

generation and circulation between metal bath and metal-slag emulsion zone.

Fig. 9 – Method for refining molten steel for optimal dephosphorisation [33]

Fig. 10 – Implementation scheme of the new dephosphorization modelling approach into the dynamic process models The enhanced dynamic online models can be used for continuous monitoring of the actual heat state in the converter as well as for prediction of its further evolution until end of blow. These information can be assessed by dynamic control functions for adaption of the amount of oxygen to be blown, lime to be charged and heating or cooling materials to be added in order to adjust the targets for temperature, carbon and phosphorus content of steel in a time and cost optimal way [37] taking into account the optimal slag conditions for dephosphorization figured out by the presented research work. CONCLUSION BOF slags are heterogeneous for a large part of the blowing process. Thus, comprehensive approaches for BOF process modelling as well as control strategies should consider the effect of solid phases from a thermodynamic and also from a kinetic 14

perspective. The presence of P-dissolving C2S_C3P phase leads to a drastic increase in Phosphorous distribution. An adequate control of this phase formation and providing suitable conditions for its enrichment on Phosphorous hold the key for a successful dephosphorization. In this context, further work is required to understand the behavior of this phase under conditions relevant to the BOF process, such as high stirring and foaming. Acknowledgment The research leading to the presented results has received funding from the European Union's Research Fund for Coal and Steel (RFCS) research program under grant agreement n° [RFSR-CT-2014-00005].

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Produzione d'acciaio con convertitori ad ossigeno REFERENCES [1]

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Khadhraoui, S; Hack, K; Janßen, T; Odenthal, H-J : Assessment of the state of P2O5-containing oxide systems relevant to steelmaking slags based on a new thermodynamic database developed for the system CaO- FeO-P2O5- SiO2 -MnO- MgO-Al2O3, submitted for publication, steel research int., December 2018.

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Besmann TM, Spear KE. Thermochemical modeling of oxide glasses. Journal of the American Ceramic Society. 2002; 85(12): 2887-94.

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Schlautmann M, Kleimt B, Khadhraoui S, Hack K, Monheim P, Glaser B, Antonic R, Adderley M, Schrama F. Dynamic on-line monitoring and end point control of dephosphorisation in the BOF converter, Proceedings of the 3rd European Steel Technology and Application Days (ESTAD), 2017, Vienna

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Bale CW, Bélisle E, Chartrand P, Decterov SA, Eriksson G, Gheribi A E, Hack K, Jung IH, Kang YB, Melançon J, Pelton AD, Petersen S, Robelin C, Sangster J, Spencer P and Van Ende M A. FactSage Thermochemical Software and Databases 2010 - 2016. Calphad. 2016; vol. 54(2016), 35-53.

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Martinsson J, Glaser B, Sichen D. Study on apparent viscosity and structure of foaming slag. Metallurgical and Materials Transactions B. 2016 Oct 1;47(5):2710-3.

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Rout BK, Brooks G, Rhamdhani MA, Li Z, Schrama FN, Overbosch A. Dynamic Model of Basic Oxygen Steelmaking Process Based on Multizone Reaction Kinetics: Modeling of Decarburization. Metallurgical and Materials Transactions B. 2018;49(3):1022-33.

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Martinsson J, Glaser B, Sichen D. The structure of foaming BOF-converter slag. Ironmaking & Steelmaking. 2017 Dec 8:1-5.

[10] Rout BK, Brooks G, Rhamdhani MA, Li Z, Schrama FN, Sun J. Dynamic Model of Basic Oxygen Steelmaking Process Based on Multi-zone Reaction Kinetics: Model Derivation and Validation. Metallurgical and Materials Transactions B. 2018;49(2):537-57. [11] Cicutti C, Valdez M, Pérez T, Petroni J, Gómez A, Donayo R, Ferro L. Study of slag-metal reactions in an LD-LBE converter. 6th International Conference on Molten Slags, Fluxes and Salts 2000 (Vol. 367). [12] Rout B, Brooks G, Rhamdhani MA, Li Z, Schrama F, van der Knoop W. Dynamic Model of Basic Oxygen Steelmaking Process Based on Multizone Reaction Kinetics: Modeling of Manganese Removal, Metallurgical and Materials Transactions B. 2018; https://doi.org/10.1007/s11663-018-1306-8 [13] Khadhraoui S, Kemminger A, Krause F, inventors; SMS group GmbH, assignee. Method for monitoring and control of P content during refining of Mn-containing melt. Germany 2018. [14] Suito H, Inoue R. Behavior of phosphorous transfer from CaO–FetO–P2O5 (–SiO2) slag to CaO particles. ISIJ international. 2006;46(2):180-7. [15] Inoue R, Suito H. Phosphorous partition between 2CaO· SiO2 particles and CaO–SiO2–FetO slags. ISIJ international. 2006;46(2):174-9. [16] Inoue R, Suito H. Mechanism of dephosphorization with CaO–SiO2–FetO slags containing mesoscopic scale 2CaO· SiO2 particles. ISIJ international. 2006;46(2):188-94. [17] Deng T, Sichen D. Study of lime dissolution under forced convection. Metallurgical and Materials Transactions B. 2012;43(3):57886. [18] Gu K, Dogan N, Coley KS. Dephosphorization kinetics between bloated metal droplets and slag containing FeO: the influence of CO bubbles on the mass transfer of phosphorus in the metal. Metallurgical and Materials Transactions B. 2017 Dec 1;48(6):2984-3001.

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Oxygen steelmaking [19] Ek M, Huber JC, Brosse G, Sichen D. Capacities of some CaO–SiO2–FeO–MnO–MgO slags and slag–solid mixtures in capturing phosphorous. Ironmaking & Steelmaking 40. 2013;40(4):305-11. [20] Preßlinger H, Fluch J, Apfolterer R. Mikroanalytische Untersuchungen über die Einbindung von Phosphor in LD-Schlacken. BHM Berg-und Hüttenmännische Monatshefte. 2012;157(3):136-43. [21] DeO B, Halder J, Snoeijer B, Overbosch A, Boom R. Effect of MgO and Al2O3 variations in oxygen

steelmaking (BOF) slag on slag morphology and phosphorus distribution. Ironmaking & steelmaking. 2005;32(1):54-60.

[22] Chen GJ, He SP. Effect of MgO content in slag on dephosphorisation in converter steelmaking. Ironmaking & Steelmaking. 2015;42(6):433-8. [23] Thornton G, D A. Ironmaking and Steelmaking. 1994;21:247. [24] Van der Knoop W, Van Unen G, Snoeijer A, Boom R. Steel Times 222:141. [25] Drain P, Monaghan B, Zhang G, Longbottom R, Chapman M, J. Chew S. A review of phosphorus partition relations for use in basic oxygen steelmaking. Ironmaking & Steelmaking. 2017;44:721-31. [26] Ito K, Sano N. Phosphorus Distribution between CaO-containing Slag and Carbon-saturated Iron at Hot Metal Temperatures. Tetsu-to-Hagane. 1983;69(15):1747-54. [27] Schoop J, Resch W, Mahn G. Reactions occurring during the oxygen top blown process and calculation of metallurgical control parameters. Ironmaking Steelmaking. 1978;5(2):72-9. [28] Subagyo, Brooks GA, Coley KS, Irons GA. Generation of droplets in slag-metal emulsions through top gas blowing. ISIJ international. 2003;43(7):983-9. [29] Alam M, Naser J, Brooks G. Computational fluid dynamics simulation of supersonic oxygen jet behavior at steelmaking temperature. Metallurgical and Materials Transactions B. 2010;41(3):636-45. [30] Sumi I, Kishimoto Y, Kikuchi Y, Igarashi H. Effect of high-temperature field on supersonic oxygen jet behavior. ISIJ International. 2006;46:1312-7. [31] Koria SC, Lange KW. An experimental study on the behaviour of an underexpanded supersonic gas jet. Archiv für das Eisenhüttenwesen. 1984;55(9):427-32. [32] Standish N, He QL. Droplet generation due to an impinging jet and the effect of bottom blowing in the steelmaking vessel. ISIJ International. 1989;29:455-61. [33] Khadhraoui S, Das S, Odenthal H-J, Krause F, Kemminger A, inventors; SMS group GmbH, assignee. Method for Refining Molten Metal Using a Converter. India; 2018 Feb 16. [34] Khadhraoui S, Defries U, Hofmann A, inventors; SMS group GmbH, assignee. Method for achieving low P-content in steel during refining of molten iron. Germany ; October 2018. [35] Münchberg W, Obst K-H, Mahn G, Nolle D. Mikrogefüge von Schlacken aus dem Sauerstoffaufblaskonverter bei Verwendung von dolomitischem Kalk und kaustischer Magnesia. Stahl und Eisen. 1981;101:587-90. [36] Hofmann, A, Reichel, J, Loginov, S, Das, S. Stable and optimized performance under suboptimal conditions; Proceedings of the 6th European Oxygen Steelmaking Conference (EOSC), 2011, Stockholm [37] Schlautmann M, Kleimt B, Teiter T, Schnabel S, Ponten HJ. Model-based dynamic monitoring and end-point control of converter process, Proceedings of the 6th European Oxygen Steelmaking Conference (EOSC), 2011, Stockholm

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Produzione d'acciaio con convertitori ad ossigeno

Conventional steelmaking route being subject to transformation P. Argenta, M. Marcozzi, M. Dorndorf, P. Duarte, S. Maggiolino, O. Demir

Several aspects in BF ironmaking technology will become the major bottleneck and restrict its further development, even if, currently, the blast furnace (BF) process still represents the predominant technology, especially in China. In this paper Tenova S.p.A. is going to present technologies in the field of direct reduction and electric steelmaking to substitute BF/BOF via DRP and EAF plant solutions in a stepwise transformation process and shows how to achieve comprehensive utilization of each composition of raw materials. The paper will demonstrate results of BF of enhanced performance and productivity with in parallel reduced coke / PCI as well as carbon footprint utilizing DRI. Furthermore the potential usage of hydrogen as reducing agent in ENERGIRON-ZR process to replace carbon carriers, the potentials in utilizing DRI in SAF to produce high purity pig iron (HPPI) and the treatment of VTM and also capabilities of sophisticated automation and measurement packages (iBOF) to enhance performance of BOF process as a very first step in this transformation process.

KEYWORDS: DRI/HBI CHARGE INTO BF - DRI+EAF AS BOF/BF SUBSTITUTE H2 AS REDUCING GAS FOR DRP - IBOF PROCESS CONTROL - OXYGEN LANCE DESIGN - DRI TO PRODUCE HPPI AND TREAT VTM, CARBON FOOTPRINT â&#x20AC;&#x201C; HYDROGEN - FLEXIBILITY

INTRODUCTION: CHALLENGES FOR IRON AND STEELMAKERS Iron- and steel makers are facing an ongoing process of transformation. Rapidly changing conditions on political and social level, approval procedures for new installations and their potential impact on public perception, raw material quality and availability, different energy sources on changing price levels, stricter environmental regulations and greenhouse gas emission reduction targets, the upcoming hydrogen era and in general difficult to predict market conditions are topics steelmakers has to take care of. Overcapacities from China and its MES status and the not finally regulated and limited to Europe stricter CO2 emission policy are additional growing threats the European iron- and steel industry has to deal with. Also Europe is facing significant structural overcapacities. But the European steel industry is doing a lot to optimize processes in terms of utilizing BF gas and COG in own power plants, using preheating technologies to enhance energy efficiency squeezing out their plants to secure profitability and economic benefits. That means European steel industry is on a high level in efficiency and on the limit in terms of technical optimization with the existing plant portfolio. The EU wants to cut its CO2 emissions by 40 % the next years and the Iron and Steel industry is one of the largest producers of CO2 emissions. These are facts. From a political perspective, there will be fewer allowances in the next years and they La Metallurgia Italiana - n. 11/12 2018

will be expensive, means the EU sets the permissible emissions firmly and reducing them further annually. In detail: The quantity of allowances should be reduced from year 2021 annually by 2,2%.

Paolo Argenta, Mario Marcozzi Tenova S.p.A. Castellanza (VA), Italy

Markus Dorndorf Tenova Metals Deutschland Essen, Germany

Pablo Duarte, Stefano Maggiolino Tenova HYL San NicolĂĄs de los Garza, Mexico

Orhan Demir

Tenova Pyromet Johannesburg, South Africa

Oxygen steelmaking

Fig. 1 – CO2 certificate price over years But the European steel industry is in a global competition with lower wages and less strict environmental regulations and is additionally facing the emission trading system and related laws and regulations. From that point of view, we are far away from a fair global competition. For the time being the price for CO2 certificates are very low (Figure 1). Studies investigated different possible scenarios with higher prices for energy, fuels and certificates [1]. Three different scenarios were considered: A Baseline Scenario (BS) which dealt with the estimated projection in production and demand, scrap availability and energy resources and CO2 prices is adhered to (based on a study of 2009), an Alternative Scenario 1 (AS1) analyzing the influence of the increase of fuel prices in the Iron & Steel Industry and an Alternative Scenario 2 (AS2) which examines the effect of the variations of the price of CO2 emissions. Target was to explore strategic opportunities for steelmakers in the current climate and how innovation in raw materials, energy distribution and transformation as well as product quality can create opportunities for business. The result can be summarized as follows: No matter which scenario occurs; it is necessary to install new technologies! To reduce the carbon footprint as expected by the European laws and regulations requires an investment in new carbon footprint reducing technologies, and this is without alternative. Potential technologies for decarburization are coke dry quenching, BOF waste heat and gas recovery, continuous casting, scrap pre-heating solutions, oxygen-fuel burners, top gas recovery turbines, CO2 capture and use for fuels/chemical production and direct reduction technologies as well as latest analyzing and measurement technologies. Most of them are already implemented in melt shops and integrated plants. But this will not be enough to meet future needs, to fulfill todays and future environmental regulations and to cope with upcoming threats. What the European steel industry and especially integrated plants need is a paradigm change, a continuous and stepwise transformation process to remain competitive.

TENOVA APPROACH: A STEPWISE TRANSFORMATION PROCESS INTO SUSTAINABLE, ENVIRONMENTALFRIENDLY AND ECONOMIC STEEL PRODUCTION iBOF technology – a step ahead in basic oxygen furnace processing Tenova’s iBOF® “intelligent Basic Oxygen Furnace” is a modular technology developed by Tenova Goodfellow Inc.- in Canada - to reduce GHG emissions while improving yield, productivity and scrap-melting capability cutting operating costs. It includes NextGen® multipoint hybrid laser/extractive off-gas analysis hardware for improved process control & safety, robust mass & energy balance endpoint process models, intelligent slop mitigation technology and automatic tapping control technology for improved safety, reduced slag carryover & minimum ladle reversion. This package represents the first of further steps means optimizing existing plant capacity – and provides features to enhance energy efficiency, process predictability and off gas treatment. Module 1 - C&T Endpoint Detection - uses a combination of Tenova’s newly launched multipoint hybrid NextGen® off-gas analysis technology together with proprietary process models and sensors. Module 1 significantly improves carbon and temperature end-point control and thereby lowers conversion costs (reduced tap alloys, O2, refractory & consumables), increases productivity (fewer reblows) and increases yield (lower FeO). Module 2 - Early Warning Slop Detection: uses advanced sensors and proprietary software to continuously monitor changes in lance vibration. Tenova Goodfellow’s - proprietary software interprets the signals in real-time to obtain a 20-40 second advance warning of the onset of a slop as well as an indication of slop severity the effects of a slop event. Module 3 - Optimized Post Combustion - enhances “in-BOF” post combustion for increased scrap melting by using a combination of proven full spectrum off-gas analysis technology together with dual flow lance technology. Module 4 - Automated Tapping Control - provides control technology for operator assist or for fully automated tapping to improve safety, minimize slag carry-over La Metallurgia Italiana - n. 11/12 2018

Produzione d'acciaio con convertitori ad ossigeno and reduce operating cost. With over 90 installations worldwide, EFSOP® extractive off-gas technology is the world leader in real-time off-gas analysis. The ENERGIRON-ZR process and its features to meet future needs Developed in pilot plant in the 1980’s and successfully started at full industrial operation in 1998, the ENERGIRON Process scheme (Figure 2), now jointly developed by Tenova HYL and Danieli & C. Officine Meccaniche S.p.A., is a major step in reducing the size and improving the efficiency of direct reduction plants. Reducing gases are generated by in-situ reforming of hydrocarbons within the reduction reactor, feeding natural gas as make-up to the reducing gas circuit and injecting oxygen at the inlet of the reactor [2]. The basic ZR scheme permits not only the direct use of natural gas (NG) but same basic

process scheme is used regardless of the reducing gas source, such as reformed gas from an external reformer, syngas from coal gasification, pet coke and similar fossil fuels, hydrogen, and coke-oven gas, among others, depending on availability. Operating conditions of the ZR process are characterized by high temperature (~1080°C) and high pressure (6-8 bar A at top gas). The elevated pressure allows a high productivity of about 10 t/h x m² and low reducing gas velocities of about 2 m/sec, as compared to lower operating pressure processes for which the gas velocities are >5 m/sec. The lower gas velocities reduce dust losses through top gas carry-over, thus lowering the overall iron ore consumption, which is reflected in overall operating costs. A distinct advantage of this process scheme without an integrated reformer is the wider flexibility for DRI carburization while using NG.

Fig. 2 – ENERGIRON ZR Process scheme As indicated in Fig. 2, the ENERGIRON ZR scheme can produce cold DRI (CDRI), hot DRI (HDRI) which can be directly fed to: 1) Hytemp System for transport and direct feeding to an adjacent EAF, 2) to briquetting presses for production of HBI and 3) to a smelting furnace (designed by Tenova) for production of pig iron. The latter is a breakthrough approach for production of pig iron using NG as reducing agent, decreasing to ~50% the carbon footprint as compared to the conventional coal-based technologies (BF, Corex, etc.). One of the key features of this technology is the inherent selective elimination of both by-products of the reduction process: H2O and CO2. As indicated in Figure 3, about 62% of total carbon input to the process and fuel

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is selectively removed via a CO2 amines-based system. This CO2 can be and is being commercialized as valuable by-product for different industries as Carbon Capture and Use (CCU) approach. As current CCU references, some examples for CO2 usage are: Ternium DRI plant (Monterrey, Mexico), Ternium DRI plant (Puebla, Mexico) to Praxair, PTKS DRI plant (Indonesia), PSSB DRI plant (Malaysia), CO2 used in food and beverages industry; JSW Salav (India). HBI/DRI plants of India are providing pure CO2 to Air Liquid for production of dry ice; ENERGIRON DRI plants (Emirates Steel, Abu Dhabi) CO2 (about 25% of total amount) will be compressed and then pumped into oil wells for enhanced oil recovery (EOR) operations.

Oxygen steelmaking

Fig. 3 – Selective CO2 removal in ENERGIRON ZR process Advantages of utilizing High-C DRI in BF operations The use of direct reduced iron (DRI) or hot briquetted iron (HBI) as metallic charge to BF allows a significant reduction of fossil fuels specific consumption. Several steelworks have already used DRI/HBI in the BF during the last decades and have reported the results. In general, each 10% of burden metallization in the mix charge, the coke rate can be decreased to 6% to 7% while the productivity can be increased by 7% to 8%. All the reported results in Figure 3 are based on the use of traditional DRI/HBI which implies carbon levels not higher than 2.0%C. A further decrease of the PCI/coke consumption and increase of the BF productivity can be reached whenever High-C DRI/HBI (≥4.0%C) is used instead of standard DRI/HBI (<2.0%C). In the ENERGIRON DRI, more than 90% of the carbon contained in the High-C DRI is in the form of iron carbide (Fe3C). In this respect, the following additional benefits are expected when using High-C DRI/HBI to the BF: • The secondary reduction of the remaining wustite (FeO) in DRI with the carbon. This reduction reaction generates CO gas which can also reduce the iron ore around the DRI, improving furnace efficiency and decreasing PCI/coke requirements. This effect is limited with the traditional DRI/HBI with lower carbon 20

content. Refer to Figure 4 for the effect of the High-C DRI in the BF. • The additional energy provided by the excess of carbon in the DRI. According to this analysis, the PCI/coke rate can be decreased down to 8% to 9% while the productivity can be increased up to 9% to 10% for each 10% of metallized burden in the feed charge if using High-C DRI/HBI. Considering a DRI/HBI charge for a typical BF-BOF installation of about 35% burden metallization, power as CO2-neutral, and depending on DR Plant location (abroad or on-site), the use of High-C DRI/HBI reduces the CO2 Emissions in 26,3% or 17,6% respectively, depending on DRP location. High-C DRI + OSBF/EAF/SAF – a solution for production of NG-based pig iron The conventional method for production of pig iron is mainly BF, based on the use of coal; coke and PCI. Other route is based on smelters like Corex, also based on coal. The novel approach by Tenova for production of pig iron, using NG as primary reductant, is based on Tenova’s OSBF Technology or Tenova EAF/ SAF combined with the propriety NG based ENERGIRON-ZR La Metallurgia Italiana - n. 11/12 2018

Produzione d'acciaio con convertitori ad ossigeno process, providing a distinct advantage in the production of high purity pig iron (HPPI). Furthermore, the carbon footprint and the overall electrical energy consumption could be reduced by utilizing our hot charge feed system from the reactor into the OSBF or EAF/SAFThe reduction of carbon footprint for production of pig iron by 50% is significant. Hydrogen in ENERGIRON-ZR process As observed in Figure 1 above, the ZR process scheme configuration is the same for any application, regardless of whether using natural gas (CH4), hydrogen (H2), reformed gas from external steam/NG reformer, syngas from coal gasifiers, or COG. This provides a significant flexibility as compared to other DR technology. For the HYL/ENERGIRON DR technology, there are some characteristics which makes this process scheme the most suitable for H2 use; i.e.: • The ZR process scheme is “natively” suitable for any reducing

gas make-up, specifically H2. • H2 make-up directly replaces NG to the process. • High operating pressure to better handle the lightest and more diffusible compound in nature. • Vast experience with high H2 concentration (~70% volume) in HYL/ENERGIRON plants. Since the 1950’s, the HYL/ENERGIRON plants, using reformed gas as source of reducing gas, includes a conventional steam/NG reformer. • Extensive pilot plant campaign in Monterrey during the 1990’s, using ≥ 90% H2 as reducing gas. From the thermodynamic point of view, H2 reduces iron oxide easily than CO, as per change of Gibbs free energy @ 845°C. On the other hand, iron ore reduction with only H2 is a highly endothermic reaction, favored at high temperatures and requiring high H2 concentrations at lower temperatures. Reduction with CO is a highly exothermic reaction, favored at low temperatures taking place at lower CO concentrations.

However, thermodynamic data do not provide information on the rate at which the reduction reactions would take place. This depends on the reactions kinetics which is determined by the prevailing process conditions. These process parameters can be only determined by experimental testing. Kinetically, iron ore reduction with H2 is >4 times faster as compared to CO and, because of the highly endothermic behavior, requires hi-

gher reducing gas temperature.The off line reformer used for DR plants is based standard technology with hundreds of references in DR and hydrogen plants; there are more than 40 HYL/ENERGIRON plants having used this type of NG reformers. Typical operation parameters of the DRP’s with reformer are as follows:

In addition to the vast industrial experience using H2, in the 1990’s, Tenova HYL carried out extensive tests at pilot plant (Figure 3) with ≥ 90% H2; producing H2 from reformed gas from the industrial DR plant after shifting and CO2 removal. These tests provided all necessary information to define: • Process and design parameters mainly related to optimized flow-temperature correlation. • DRI quality in terms of metallization and carbon content,

which is less than 1% above 75% H2. • Optimization of operating pressure, reactor L/D ratio, solids residence time, to consistently achieve the DRI quality, determination of fluidization factor (ƒ) to ensuring proper gas velocities and distribution through the solids bed, among others for the proper gas distribution and design of the scheme for H2 utilization.

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Oxygen steelmaking

Fig. 4 – HYL Pilot Plant tests campaigns with ≥90% H2 Additional benefits of the ENERGIRON ZR technology for the direct use of H2 is related to the high operating pressure (6-8 bar). H2 is the most diffusible gas in nature thus, tight sealing is required to prevent any leak. In the HYL/ENERGIRON DR plants, the high operating pressure is managed through the use of mechanical sealing for pressurization/de pressurization from atmosphere. Mechanical sealing allows higher P due to friable

ores and/or additional fines, preventing any gas leak/air intake. For this case, any additional P is simply handled by the compressor instead of lower gas flow and/or decreasing production rate. On the other hand, the gas velocity of the gas inside any moving bed reduction shaft has the following functionality:

where: ƒ is the fluidization factor. ƒ shall NOT exceed 1,0 Vmf is the velocity of minimum fluidization; i.e. gas velocity at which the solids flow is suspended k is a function of bed porosity, gas density and viscosity, solids density, particle size FG(act) = Specific reducing gas flowrate (actual cond.) Fs = Production rate M = Gas molecular weight T = Reducing Temperature at reactor inlet P = Operating pressure Di = Reactor internal diameter

It can be easily noted that a high operating pressure shaft allows better flexibility for a small diameter (Di), to comply with fluidization, and larger height (L), for a better gas distribution. 22

In terms of energy consumption, the impact of H2, as compared to NG is indicated in Figure 5, crediting the %C in the DRI.

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Produzione d'acciaio con convertitori ad ossigeno

Fig. 5 – Comparative energy consumption figures with H2 use for Direct Reduction in ENERGIRON plants As summary and based on the above facts, ENERGIRON is the only available DR technology fitted for the proven, efficient and reliable use of H2. For this industrial application, to comply with the CO2 emissions targets, H2 shall be generated by high efficiency electrolyzers, like the reversible high-temperature electrolyzer (HTE) type, powered by renewable energy source; i.e., eolic, solar PV. A stepwise replacement of BF/BOF plant capacity by DRP/EAF facilities - CDA The ENERGIRON-ZR process with its characteristics and features such as selective CO2 removal, hydrogen utilization and High-C DRI production provides already the technology to meet future challenges. The SALCOS (Salzgitter Low CO2 Steelmaking) project, a study initiated by Salzgitter AG together with Tenova and Fraunhofer-Gesellschaft (FhG) in 2015, had to analyze the capabilities of already existing technologies to reduce greenhouse gas emissions, to investigate implications on integrated steel works and so to demonstrate the possibility for significant contribution to carbon footprint reduction. The advantage of SALCOS approach and the incorporation of ENERGIRON-ZR/HYL technology is the possibility to go directly for a large industrial-scale pilot plant avoiding the necessity for laboratory scale studies. Especially taking into consideration visible changes of climate, the changed perception of global society as well as the defined climate targets for 2050, this concept provides the possibility to act now. The general approach behind is, in contrast to CO2 usage (CCU) or disposal La Metallurgia Italiana - n. 11/12 2018

(CCS) concepts, the avoidance of carbon carriers and the usage of hydrogen instead. Hydrogen can replace carbon in iron ore reduction processes, leading to the final formation of water (H2O) rather than CO2. Additionally, process heat for steelmaking may be supplied by electrical energy instead by carbon. Based on these simple facts Salzgitter defined the term CDA – “Carbon Direct Avoidance” in order to underline the difference to already existing and not consequent utilized and well thought out concepts. Avoidance of CO2 formation directly in steelmaking processes is more sustainable and also energetically more useful than any further utilization. As one of the major results the SALCOS study could demonstrate the possibility to follow stricter CO2 reduction targets in Europe after 2030 by realizing a stepwise transformation process of integrated iron and steel works towards direct reduction and electrical energy based steelmaking processes. This transformation process, realized in subsequent steps, reduces the environmental impact in terms of CO2 emissions up to around 95% depending on framework conditions. As a first step, an additional gas-based direct reduction plant (DRP) (ENERGIRON ZR process) has to be realized at the integrated site in Salzgitter. The produced high carbon DRI (HC-DRI) from this plant is utilized in existing BF’s to enhance productivity and to reduce coke as well as PCI in parallel. This step already reduces the carbon footprint of steel production of Salzgitter by around 10%, as natural gas used for reduction has a certain amount of hydrogen content. With electrolysis on an industrial scale hydrogen can further replace natural gas and so carbon carriers partly. In case of operating 23

Oxygen steelmaking electrolyzers with power from renewable resources only, the overall CO2 emissions can be reduced up to 18%. Precondition is the availability of electrolyzes capacities on a very large scale (largest capacity ever realized) to provide an appropriate gas mixture (natural gas/hydrogen). In conjunction with the flexibi-

lity of the DR plant / ENERGIRON ZR process utilizing different reducing gases in varying ratios in full operation, this concept represents the most advanced solution for sustainable steelmaking in industry.

Fig. 6 – The maximum CO2 reduction possible by the SALCOS® concept in this ultimate configuration is 95% (source: SZAG). The next step will be the incorporation of a melt shop to produce steel via electric arc furnace route. Distributing HCDRI to EAF (via hot charging / pneumatic transport system – HYTemp) as well as to BF’s or storage bin in different ratios provides highest flexibility in raw materials (Hot DRI, Cold DRI, Scrap) and performance control. This allows further to shut down one of the three BF’s in operation in order to reduce the CO2 emissions significantly up to 25%. Further steps in this transformation process are principally based on the same approach as the steps before, leading to the complete change of steelmaking from the blast furnace/basic oxygen technology to the direct

reduction/electric arc furnace route. With the final configuration the entire integrated steelworks will be transformed and the resulting reduction in CO2 emissions will be in the range of 95% (Figure 6). That means vice versa the remaining carbon footprint of Salzgitter site will be only 5% compared to the present day. Figure 7 visualizes the potential transformation process with / without CO2 off-taking and utilization the of hydrogen in varying quantities as well as its effects on carbon footprint from BF/BOF route to DRP/EAF route in general (approximate values).

Fig. 7 – Stepwise transformation process w/wo CO2 off-taking and utilization of hydrogen in varying quantities 24

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Produzione d'acciaio con convertitori ad ossigeno This stepwise transformation process requires next to considerable investments in new plant equipment adjustments to given regulatory and economic framework to avoid unreasonable OPEX increase, for instance taxes on electrical energy (EEG allocation). These conditions have firstly to be adapted in order to facilitate the realization of the transformation project. To make this transformation process and its effects on greenhouse

gas emissions more tangible, the numbers has been transformed into analogies (Figure 8). With the final configuration of Salzgitter site after transformation the reduction of CO2 emissions will be equivalent to CO2 emission of around 4,6 million cars (around 10,5% emissions of all cars in Germany per year) or is equivalent to 500 million trees (around 6,6 times Berlin city area covered with trees).

Fig. 8 – Equivalents for CO2 reduction of SALCOS project These comparisons visualizing the tremendous effects and the important role steel industry can play in future. The major problem remains and can be reduced to a simple formula: European steel industry has to act locally on CO2 emissions, but to compete worldwide. This implies the willingness of society to invest (public funding) to avoid carbon leakage to non-ETS countries and to keep steelmaking in Europe beneficial and economic. Summary / Final statements Tenova provides sophisticated technologies and features to meet future needs of our customers. Considering present challenges for iron- and steel makers e.g. rapidly changing conditions on political and social level, approval procedures for new installations, raw material quality and availability, different

energy sources on changing price levels, stricter environmental regulations and greenhouse gas emission reduction targets, the upcoming hydrogen era as well as overcapacities in the market, Tenova with its experience, know-how and technologies can support the transformation of making steel in future in a more sustainable, environmental-friendly and economic way. Key attributes we see are: • Highest raw material flexibility and cost control • Ability to fulfill todays and future environmental regulations for all process conditions • Utmost reliability and repeatable process and energy efficiency • Ability for adjustments and modifications to meet future challenges

REFERENCES 1]

Prospective Scenarios on Energy Efficiency and CO2 Emissions in the EU Iron & Steel Industry, European Commission Website, 19.02.2016

Duarte, Pablo; Martínez, Jorge. By-Products and Emissions in an ENERGIRON DR Plant. AISTech 2016. 16-19 May Pittsburgh, PA

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Oxygen steelmaking

Estimation of trace elements in unalloyed scrap R. Ammer, P. Reisinger, M. Egger, R. Tober, G. Salzmann Controlling the concentration of trace elements is crucial for the production of demanding steel grades like electrical or interstitial free steel. The charge mix of different scrap qualities and hot metal should minimize the risk of exceeding the limits of the trace elements chromium, nickel, molybdenum and copper. At the same time, it is crucial to optimize the choice of the input materials towards cost effectiveness. Steel scrap can be highly inhomogeneous, which makes representative sampling extremely difficult. In addition to taking random samples at voestalpine Stahl GmbH in Linz a statistical approach to estimate the content of trace elements is utilized. Using hot metal-, steel- and slag-samples the concentration of the elements in scrap can be calculated. Assuming that these concentrations are distributed normal, the corresponding mean values and standard deviation can be estimated and is used in calculation models. The outcome of this balance can be utilized for multiple purposes. Knowing the underlying distributions of the input materials makes it possible to determine a corresponding distribution of the raw steel. On a short term basis it is possible to detect unintentional mixing of alloyed and unalloyed scrap. Using this information, appropriate measures at the scrapyard can be taken. At the same time, the data over several years shows long term developments of certain scrap grades.

KEYWORDS: SCRAP – RECYCLING - TRACE ELEMENT – BOF – STATISTICAL MODELLING - PRODUCT QUALITY INTRODUCTION: RECYCLING OF STEEL AND TRACE ELEMENTS Steel is known for its ability to be multi-recyclable (1). Subsequently, from an environmental point of view, it is desirable to operate the LD-converter with high scrap rates. The voestalpine Stahl GmbH operates three LD-converters with a tapping weight of 175 t. During normal operation the amount of charged scrap ranges from between 20 to 30 % of the total metallic input.

Rainer Ammer, Peter Reisinger, Martin Egger, Rudolf Tober, Gerhard Salzmann Voestalpine Stahl GmbH, Linz, Austria

Fig. 1 – Steel's life cycle (1) 26

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Produzione d'acciaio con convertitori ad ossigeno As illustrated in figure 1 scrap is “produced” during manufacturing (pre-consumer steel scrap) and the end of the product life (post-consumer steel scrap). Increased use of steel subsequently leads to an increased amount of available scrap. At the same time, many steel grades have high demands regarding the level of impurities, respectively narrow target windows for the concentration of alloying elements. These trace elements

are not added deliberately to the steel and can have a negative effect on the material properties. Among the long list of possible trace elements those which cannot be removed using common steelmaking practice are especially important to control. The following table presents an overview of possible trace elements and its import sources (2).

Tab. 1 – Removability and source of trace elements (2)

Generally steelmaking offers two methods to extract elements which are dissolved in the melt. One pathway is through oxidizing reactions, which lead to a transfer of the element(s) into the slag phase. At steelmaking temperatures some elements can also be transferred into the gas phase and the vapor pressure of an element therefore plays a crucial role. The elements which cannot be removed are those which have a lower affinity towards oxygen compared to iron and those with low vapor pressures. Consequently to control the input of these elements is of great importance. Table 1 also indicates that knowledge of the charge material is a key component towards control of steel quality, since the source of the majority of elements is the scrap. The selection of raw materials depends on a variety of factors, and often these additional criteria conflict with the desire to have a mix with low problematic trace elements. For instance: ◦ The metallic yield has been optimized along the entire production chain. This means that the amount of high quality preconsumer scrap with known chemical analysis is lower nowadays. ◦ The input material should be as cost effective as possible. This means that the amount of low-quality post-consumer scrap is

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generally increasing. ◦ From an economic as well as environmental point of view the use of by-products such as slag or dust products (briquettes, granulate) is favorable. This leads to an increased input of trace elements. ◦ The demand for coated as well as high alloyed steels is high, which leads to an increased level of trace elements within the scrap. (3) The following paper will present an approach to estimate the risk of excessive concentration of trace elements. Because of their high relevance in the steelmaking practice, the focus of this work will be on the elements Cr, Ni, Mo and Cu. Due to the highly inhomogeneous nature of scrap, sampling is difficult, time consuming and not representative. Therefor the mass flow and chemical analysis of all output- and input-streams at the LD-converter was observed. SCRAP TYPES In the present paper three different typs of scrap are analysed. The first scrap type is collected at different process steps within the production facility of voestalpine Stahl GmbH in Linz. The other two grades are aquired on the scrap market.

Oxygen steelmaking Tab. 2 – Analyzed scrap grades scrap types type

origin

specification

type

internal

low alloy steel

pre-consumer

external

E6, E8

pre-consumer

external

post-consumer

Different scrap qualities can be distinguished according to the european steel scrap spezification. Both E6 and E8 scrap are defined as thin new production steel scrap which is uncoated. According to the specification, the sum of the trace elements must be smaller than 0.3 [wt%]. In contrast to the quality E8, the quality E6 must be compressed into scrap packages in order to ease the charge of the scrap. E1 scrap is old used steel scrap with a thickness lower than 6mm. According to the specification, the copper content must be below 0.4 [wt%]. The upper bound for the sum of chromium, nickel and molybdenum is 0.3 [wt%]. (4) MASS BALANCE In order to calculate the concentrations of the elements Cr, Ni, Mo and Cu in alloyed scrap grades a mass balance was carried out. The input materials of the LD-process are hot metal, scrap, slag formers, cooling- and heating agents as well as oxygen. The output consists of a dust loaded off-gas, crude steel and slag. The composition of the different input– and output stre-

ams can be determined by the use of chemical analysis such as optical emission spectroscopy (OES) or X-Ray Fluorescence Spectroscopy (XRF). The frequency of the sample taking routine can differ significantly between the materials. Regarding the input materials the mass is measured by a crane scale or by the weighing system of the addition bins. The determination of the quantities of the output streams is not that straightforward. The mass of the crude steel is weighted at the steel car after tapping. There has to be a correction in order to take the additions during tapping, slag carry-over and residual steel within the converter into account. For the mass balance of chromium, the amount of oxidized chromium must also be taken into account. Therefore knowledge of the slag mass is necessary. The mass of slag can be estimated by a CaO mass balance. For a more accurate estimation, knowledge of other oxides is taken into consideration as well. The exact procedures for the calculation of steel and slag masses are beyond the scope of this paper. Condensed information about the input and output streams is presented in table 2 and table 3.

Tab. 3 – input material input material

quantity

composition

frequency

hot metal

weighted

OES

per heat

scarp

weighted

unknown

per heat

solid slag dust

weighted

XRF

per week

briquettes

weighted

XRF

per month

Tab. 4 – output material output

material

quantity

composition

frequency

dust

average values

XRF

per month

crude steel

weighted mass

OES

per heat

slag

balance

XRF

per heat

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Produzione d'acciaio con convertitori ad ossigeno It can be deduced from the tables 2 and 3 that the quantities and compositions of all relevant material flows are known. The only exception is the composition of the scrap. Calculating a

mass balance (difference between output and input) for all desired elements yields the desired composition. (1)

concentration of the element i

[wt%]

mass of the material flow

[kg]

crude steel

hot metal

slag

SCS

solid converter slag

dust

dust briquettes

SCRAP

scrap

The mass balance [1] can be used for heats which consist only of hot metal and one type of scrap. Although scrap types of lower quality are only used to a limited extend and are often not charged alone, this method is still beneficial in order to gain knowledge about these types. An example of a scrap type which is only used to a limited extent is type 3, while the type 1 and type 2 scraps are used in larger quantities. Mixture of scrap types For the case described above the concentration đ?&#x2018;Ľđ?&#x2018;&#x2013;đ?&#x2018;&#x2020;đ??śđ?&#x2018;&#x2026;đ??´đ?&#x2018;&#x192; consists

The mass respectively concentration of these less used scrap qualities are denoted by đ?&#x2018;&#x161;đ?&#x2018;&#x2013;đ?&#x2018;Ąđ?&#x2018;Śđ?&#x2018;?đ?&#x2018;&#x2019; đ?&#x2018;&#x17D; and đ?&#x2018;Ľđ?&#x2018;&#x2013;đ?&#x2018;Ąđ?&#x2018;Śđ?&#x2018;?đ?&#x2018;&#x2019; đ?&#x2018;&#x17D;.The most important parameter in equation [2] which is required is the concentration of every element đ?&#x2018;&#x2013; within the scrap type đ?&#x2018;&#x2014;. For this calculation the result of the last heat where exclusively type đ?&#x2018;&#x2014; scrap was used is taken. Using equations [1] and [2] the concentrations for a significant number of heats can be calculated. These results are then taken as inputs for a statistical analysis of the underlying elemental distributions within the various scrap types.

of different scrap types. Some scrap types are taken as input parameters for a mixed mass balance. These are types where a significant amount of heats are produced, which have this scrap type as sole solid metal input. Assuming we have a scrap mixture which consists exclusively of scraps which we can use as input and only one unknown scrap type the result of [1] can be used as input for a mixed mass balance. The equation [2] shows the approach for calculating the concentration of less frequently used scrap qualities.

(2)

STATISTICAL APPROACH Smooth kernel distribution Having calculated the content of numerous scrap charges the question arises whether it is possible to calculate the underlying probability density function. A method to get such a function is the use of kernel density estimation (KDE), which is a well described non-parametric approach (5). This means that no prior knowledge of the underlying distribution is necessary. The idea is to create a function which itself is a mixture of continuous distributions. We consider the independent samples X1, â&#x20AC;Ś, Xn and calculate the density function by using equation [3]. (3)

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Oxygen steelmaking The amount of smoothing is controlled by the smoothing bandwith h>0. K is a probability density function called kernel. One

of the most common functions used as K is the so called Gaussian kernel. (4)

In contrast to selecting various kernel functions, which only has a marginal influence on the outcome in most cases, the value of the bandwidth h has to be chosen carefully (5). Fig. 2 illustrates

the influence of different factors on the resulting KDE for the approximation of a discrete distribution.

Fig. 2 â&#x20AC;&#x201C; Influence of the bandwidth on the KDE Methods to determine to smoothing factor automatically do exist. To determine the distributions which are used in the present paper, this factor was chosen manually in order to guarantee a consistent result. Normal distribution The kernel densities are mixtures of probability densities and therefor are probability densities themselves.

Nevertheless for some application it is beneficial to approximate the KDE with a single normal distribution. This makes it easier to illustrate the outcome of the calculations and at the same time the results can readily be implemented in existing models. To approximate the underlying distribution there are quite a few possibilities. Two approaches have been tested. In order to obtain the first fit (fit 1) the minimization problem [5] was solved. (5)

The second approach is to find the best fit for the cumulative

distribution (fit 2), which is described using equation [6].

(6)

For both minimization problems the Nelder Mead global optimization algorithm (6) was used to determine Îź respectively Ď&#x192;. The increment for the discretization was chosen as đ?&#x2018;˘đ?&#x2018;?đ?&#x2018;?đ?&#x2018;&#x2019;đ?&#x2018;&#x;/ đ??żđ?&#x2018;&#x2013;đ?&#x2018;&#x161;đ?&#x2018;&#x2013;đ?&#x2018;Ą/103, which is a compromise between accuracy and

computation speed. The đ?&#x2018;˘đ?&#x2018;?đ?&#x2018;?đ?&#x2018;&#x2019;đ?&#x2018;&#x;đ??żđ?&#x2018;&#x2013;đ?&#x2018;&#x161;đ?&#x2018;&#x2013;đ?&#x2018;Ą is the highest observed concentration. Fig. 3 shows a comparison of the two methods for one of the distributions.

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Produzione d'acciaio con convertitori ad ossigeno

Fig. 3 –Approximation of a KDE Visual inspection of several approximations showed that generally fit 2 maps the underlying distribution better. The tail on the right side of the distribution originates from rather high concentrations. These values can be a consequence of contamination of the scrap (e.g. plated steel sheets) or are consequences of a false classification. These events can’t be mapped by a normal distribution since the probability density function is strictly monotonous.

RESULTS The presented logic was used to calculate the mean value and standard deviation for three scrap types as well as for the hot metal. In the case of the hot metal the concentrations for every heat are measured. The distributions were calculated using the methods presented in the paragraphs above.

Tab. 5 – Mean and standard deviation per input material scrap composition type 2

type 1 element

μ [wt%]

0.0668

σ [wt%]

μ [wt%]

0.0328

0.0605

type 3

σ [wt%]

μ [wt%]

0.0242

0.1739

hot metal σ [wt%]

μ [wt%]

0.1522

0.0275

σ [wt%] 0.0054

0.0177

0.0091

0.0192

0.0081

0.0884

0.0423

0.0053

0.0048

0.0040

0.0044

0.0033

0.0150

0.0121

0.0017

0.0015

0.0195

0.0047

0.0177

0.0052

0.2170

0.1025

0.0044

0.0015

Type 1 und type 2 can be considered as high quality scrap which show relatively low concentrations of trace elements. The concentrations of these two scrap grades are relatively similar. Compared to hot metal, the concentrations are between 2.4 (chromium) and 4.4 (copper) times higher than those of the hot metal. Type 3 is a post-consumer scrap and therefor the load of trace elements is significantly higher. Especially significant, is the difference regarding the copper concentration which is 12 times higher in type 3 compared to type 1 and type 2 scrap. Despite these higher trace element values, the use of type 3 scrap can provide an economic advantage since its price is comparably low. Besides being a useful method to estimate the input of trace elements for both individual scraps and mixture of scraps, it is also helpful in tracking long term trends regarding the composition of different scrap types. It can be seen that generally the amount of copper in unalloyed steel scrap decreased over La Metallurgia Italiana - n. 11/12 2018

0.0017

the past years. The reason might be increased efforts of scrap recycling companies to minimize to contamination with high copper materials like wires or parts of electric engines. At the same time it can be observed that the molybdenum concentration is increasing. In the case of the voestalpine Stahl GmbH the most important source of pre-consumer scrap is the automotive industry. While molybdenum is usually not used as an alloying element for steel-qualities which are used for the outer skin of a vehicle, it is added to high strength steels which are used for structural parts. An example is dual-phase steels where molybdenum delays the formation of ferrite and pearlite, which has an positive effect on the hardenability (7) . The following paragraph shows a method to calculate the estimated distribution for arbitrary input material mixes. The method is especially useful in quantifying the risk of an excessive trace element concentration.

Oxygen steelmaking ESTIMATION OF RISK The advantage of obtaining normal distributions, as presented in the preceding paragraphs, is that they can be readily used for further calculations. For example, it is possible to calculate the probability that the concentration of a specific element will be exceeded for a specified raw material mix. Since the distributions of all input materials are known, the distribution of the crude steel can be estimated. It is a well-known fact from

statistics that the linear combination of normal distributions is again a normal distribution. This holds also if we add weights to the normal distributions. So we would like to calculate đ?&#x2018;&#x2020;=ÎŁđ?&#x2018;?đ?&#x2018;&#x2013;Âˇđ??śđ?&#x2018;&#x2013;đ?&#x2018;&#x203A;đ?&#x2018;&#x2013;=1 where đ??śđ?&#x2018;&#x2013; Ě´ đ?&#x2018; (đ?&#x153;&#x2021;đ?&#x2018;&#x2013;,đ?&#x153;&#x17D;đ?&#x2018;&#x2013;2) and đ?&#x2018;&#x2020; Ě´ đ?&#x2018; (đ?&#x153;&#x2021;đ?&#x2018;&#x2020;,đ?&#x153;&#x17D;đ?&#x2018;&#x2020;2) are normal distributions where Îź is the mean and Ď&#x192; the standard deviation. The parameters đ?&#x153;&#x2021;đ?&#x2018;&#x2020; and đ?&#x153;&#x17D;đ?&#x2018;&#x2020; can be calculated using the following formulae: (7)

Using the information from table 5 the distribution of an arbitrary charge-mix can be calculated. The following figure illustrates the probability of exceeding the limits for nickel which are 0.02 respectively 0.03 [wt%]. While 70 % of the input is hot

metal for the remaining 30 % the amount of type 3 scrap varies between 5 and 20 %. The rest consists of type 1 scrap. Making use of the cumulative distribution function a quick picture of the situation is possible.

Fig. 4 â&#x20AC;&#x201C; Cumulative Distribution of the Ni-content It can be seen that a concentration below 0.02 [wt%] is quite unlikely (below 50 % probability) if the amount of type 3 scrap exceeds 15 %, while adding 5 % of type 3 scrap should be acceptable. If higher limits are tolerable the amount of type 3 scrap can be increased. Fig 4 is also an example of a possible scenario, taking the limit of a certain element as input it would be a straightforward calculation to determine the maximum amount of a certain scrap type in order to keep to risk of exceeding this limit below a certain threshold. CONCLUSIONS The paper presents an approach for the estimation of trace elements in scrap which can be used in LD- or EAF-steelmaking.

The presented approach is statistical in nature, and although singular events like the mix-up of different scrap grades cannot be predicted, this method provides a reliable routine of characterizing scarp mixes and overcomes many of the issues associated with scrap characterization by manual sampling alone. It is also obvious that in order to ensure a cost effective input mix, scraps of varying quality should be selected, however this often increases the risk of exceeding specified upper bounds for various trace elements, particularly for certain steel grades. The advantage of the presented method is that this risk can be quantified. Another area of application is the analysis of long term trends regarding the composition of certain scrap grades.

La Metallurgia Italiana - n. 11/12 2018

Produzione d'acciaio con convertitori ad ossigeno REFERENCES [1]

Worldsteel association. Life cycle assessment in the steel industry [Internet]. worldsteel position paper. 2017 [cited 2018 Aug 8]. Available from: https://www.worldsteel.org/en/dam/jcr:22f5ed93-0311-4f8a-a90c-b9ac424a3d01/ LCA%2520position%2520paper.pdf

[2]

Lachmund H, Bannenberg N. Comparitive assessment of LD converter and electric arc furnace steelmaking in terms of removal of undesirable tramp elements. In: AISTECH conference proceedings. 2008.

[3]

Jungreithmeier A. Verhalten der Spuren- und Begleitelemente in der LD-Stahlerzeugung. 2006.

[4]

EUROFER. European Steel Scrap Spezification [Internet]. 2008 [cited 2018 Aug 16]. Available from: http://www.eurofer.org/ Facts&Figures/ws.res/EurSteelScrapSpec.pdf

[5]

Silverman B. Density estimation for statistics and data analysis. Chapman Hall. 1986;

[6]

Nelder J., Mead R. A Simplex Method for Function Minimization. Comput J. 1965;

[7]

Rana R, Singh S. Automotive Steels: Design, Metallurgy, Processing and Applications. Automotive Steels: Design, Metallurgy, Processing and Applications. 2016.

La Metallurgia Italiana - n. 11/12 2018

È fresco di pubblicazione il nuovo Capitolato acciai elaborato da un gruppo di lavoro che coinvolge rappresentanti del mondo della produzione d’acciaio per utensili, delle fonderie, del trattamento termico, dell’università e dei laboratori di prova e diagnostica. Più in dettaglio, il documento è nato da un’iniziativa del Centro studio Pressocolata dell’Associazione Italiana di Metallurgia (AIM), di cui il gruppo di lavoro fa parte, a seguito dall’accertata necessità di rendere disponibile agli utilizzatori d’acciaio per stampi da pressocolata un documento tecnico che permetta di focalizzare i requisiti che i materiali dovrebbero possedere per assicurare un adeguato livello di prestazioni degli stampi in esercizio. Si tratta, quindi, di un documento, unico in Italia, utile a fonderie e stampisti nella selezione dell’acciaio più adatto per la specifica applicazione, dal momento che riporta le caratteristiche ed i limiti d’accettabilità dei principali acciai destinati alla costruzione di utensili primari e secondari per la pressocolata delle leghe leggere, con un attenzione particolare agli acciai per utensili a caldo a medio tenore di carbonio e legati al cromomolibdeno-vanadio. Inoltre, si applica a barre, blocchi, dischi fucinati e barre laminate d’acciaio per utensili da lavorazioni a caldo, forniti allo stato ricotto e dopo trattamento termico di tempra e rinvenimento, focalizzando l’attenzione esclusivamente sugli aspetti metallurgici di fabbricazione e di trattamento termico. Considerata l’evoluzione tecnologica e scientifica del settore, è previsto che l’attuale versione sia aperta a revisioni ed integrazioni. ISBN: 9788898990177 Disponibile su Amazon (www.amazon.it/dp/B07L33SXKL) o in versione cartacea come indicato sul sito www.aimnet.it Annalisa Pola - Università degli Studi di Brescia

Collana Tecnica AIM Il diagramma di stato Fe-C e le curve TTT

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Gds - Centro MTA Milano, 12 marzo MASTER PROGETTAZIONE STAMPI Corso - Centro P marzo/maggio LEGHE DI NICHEL E SUPER LEGHE GdS - Centro ME Milano, 11 aprile METALLI A SCALA MICRO - NANOMETRICA: TECNICHE DI INDAGINE GdS - Centro MFM Roma, 17 maggio

Corso - Centro R Placcatura e Termospruzzatura Milano, 19-20 giugno Corso - Centro R Milano, 19-20 giugno MICROSCOPIA ELETTRONICA IN SCANSIONE SEM PER METALLURGISTI II ed. MICROSCOPIA ELETTRONICA IN -SCANSIONE SEM PER Corso - Centro MFM METALLURGISTI - II ed. Lecco, 26-27-28 giugno Corso - Centro MFM Lecco, 26-27-28 giugno PRESS & SINTER: EVOLUZIONE DI PRODOTTI, MATERIALI E TECNOLOGIE PRESS & SINTER: EVOLUZIONE DI PRODOTTI, GdS - Centro MP MATERIALI E TECNOLOGIE giugno GdS - Centro MP giugno DEFORMAZIONE DEI PRESSOCOLATI: CAUSE E RIMEDI GdS - Centro P DEFORMAZIONETorino, DEI PRESSOCOLATI: 18 settembreCAUSE E RIMEDI

Milano, 29-30 maggio

GdS - Centro P XIII GIORNATE NAZIONALI CORROSIONE E PROTETorino, 18SULLA settembre ZIONE Convegno – SEGR. XIII GIORNATE NAZIONALI SULLA CORROSIONE E PROTE-

Torino, 5-6 giugno

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Convegno – SEGR. Palermo, 3-4-5 luglio Per ulteriori informazioni rivolgersi alla Segreteria AIM e-mail: info@aimnet.it oppure visitare il sito internet www.aimnet.it La Metallurgia Italiana - n. 11/12 2018

Industry news Dissolution behaviour of ULC steel in carbon saturated hot metal edited by: F. M. Penz, J. Schenk, R. Ammer, K. Pastucha, B. Maunz In addition to hot metal, the Linz-Donawitz oxygen steelmaking process (LD) uses scrap as an iron source. Aside from this fact, scrap acts as a coolant for the exothermic reactions inside the LD converter and will be dissolved in the hot metal. The optimization of the LD process is also focussed on the thermodynamic and kinetic modelling where literature-based dissolution equations are used. In laboratory scale experiments the dissolution behaviour of ULC steel scrap in hot metal with two different carbon concentrations was investigated. For the evaluation of the experiments, a literature model for diffusive melting of scrap in hot metal was examined. Based on the measured ablation rate of cylindrical scrap samples submerged in hot metal, the mass transfer coefficient for the dissolution of ULC steel was determined.

KEYWORDS: BASIC OXYGEN FURNACE – STEELMAKING – SCRAP DISSOLUTION – THERMODYNAMICS –PROCESS MODELLING

Florian Markus Penz K1-MET GmbH, Linz, Austria;

Johannes Schenk K1-MET GmbH, Linz & Chair of Ferrous Metallurgy Montanuniversitaet Leoben, Austria;

Rainer Ammer voestalpine Stahl GmbH, Linz, Austria;

Krzysztof Pastucha Primetals Technologies Austria GmbH, Linz, Austria;

Bernhard Maunz voestalpine Stahl Donawitz GmbH, Donawitz, Austria;

INTRODUCTION Oxygen steelmaking in an LD converter was developed in the early 1950s in Linz and Donawitz (LD) and it gradually became the most dominant method of crude steel production. Large quantities of scrap are used for an ordinary blowing process in an LD converter besides hot metal, which is the main charging material. Scrap is mainly used as a coolant for the process due to heat generation from the oxidation reactions of carbon, silicon, manganese and phosphorus. The dissolution and melting behaviour of scrap influences the whole process cycle in the converter. Many articles concerning the scrap melting process have been published in the past, whereby only a few describe the kinetics of a special steel scrap 36

like ULC in hot metal with changing carbon content. (1 - 12) In this work, an experimental investigation of the melting behaviour of ULC (ultra-low carbon) steel scrap in hot metal with different carbon content was executed. The experiments are based on thermodynamic and kinetic calculations through a MatLab® coded single-zone LD model. Detailed descriptions of the model are published in (13 - 16). In previous calculations it was shown that low alloyed scrap types tend toward a stagnation of dissolution and melting during the oxygen blowing process using literature-based analytical equations (15). Based on these results, laboratory scale experiments were done for validation.

La Metallurgia Italiana - n. 11/12 2018

AttualitĂ industriale DERIVATIONS OF THE OBSERVED DATA A collection of measurement data from LD-converters by various authors on the carbon concentration of the molten steel and the melt temperature during the blowing process is presented in (3). It was found that the change of the carbon content is a function of the melt temperature and its trend is parallel to the liquidus line (austenite-liquid melt) in the Fe Fe3C phase diagram. Most of the data points are in the liquid phase but some are in the two-phase area of austenite and liquid melt. A similar behaviour is also modelled with the LD model used and was partly published in (15). To describe the scrap melting and dissolution behaviour following literature-based considerations were applied.

The literature describes the melting of scrap in different phases where the diffusion process of carbon controls the scrap dissolution as long as the metal phase temperature is below the melting point of the scrap. If the temperature exceeds the melting point of the scrap a model for forced scrap melting is used. The scrap melting point is defined by the liquidus temperature in the phase diagram for the given chemical composition of the scrap. The rate of diffusive melting of scrap is determined by the diffusive mass transport of carbon between the liquid metal and the charged scrap. Equation [1] is a mathematical model, proposed by Zhang L. and F. Oeters (6), where the mass transfer coefficient kmet in [m s-1] is a decisive factor.

(1)

The scrap particles are assumed as spheres in the model used and the radius of the particle is r in unit [m]. Cscrap and CHM are the carbon concentrations in the scrap and hot metal in [wt.%]. Cliq describes the carbon concentration of the scrap on the liquidus line at a given temperature. (6) In this work, the liquidus line was determined for the actual silicon and manganese content of the scrap using the Fe-Fe3C phase diagram generated by the FactSageTM FSstel database. (13, 15, 17) When the temperature exceeds the scrap melting point, forced or convective scrap melting starts, where the mass transfer could be neglected since the heat transfer becomes dominant for the dissolution process. (5, 6, 18) In comparison to diffusive scrap melting, the driving force the temperature difference between scrap and hot metal during forces scrap melting. The equation for the forced scrap melting is cited in other publications. (15, 16) Forced scrap melting is not relevant for this research work, since the melting point of the ULC scrap was above the hot metal temperature in all experiments.

La Metallurgia Italiana - n. 11/12 2018

DESCRIPTION OF EXPERIMENTS Dissolution tests For the experimental procedure, an alumina crucible was charged with 320 g to 345 g of hot metal and positioned in a high temperature vertical tube furnace. The heating rate was 300 K/min to reach the starting temperatures, which are listed in Tab. 1. Before the first and between each further dissolution experiment, a holding time of 30 min was set. To prevent oxidation of the hot metal, the vertical tube furnace was flushed with nitrogen during the heating and melting process. The scrap geometry was cylindrical with a diameter of 12 mm and a length of 30 mm. The whole specimen, including the sample holding, is shown in Fig. 1. The cylinder was submerged into the hot metal with a depth of 20 mm, whereby the axial heat flow to the specimen holding was diminished through a notch. The starting temperature of the specimen was 25 Â°C. The axial movement of the cylinder to the melt was carried out with a vertical pneumatic controlled cylinder. No stirring by rotation of the cylinder or crucible was performed during the dissolution experiment (i.e. static conditions).

Industry news

Fig. 1 - Sample geometry including the sample holding part (units in mm) Before each dissolution experiment, the mass of the cylinder was determined. After a defined dissolution time, the cylinder was extracted from the melt and immediately quenched with water to inhibit further carbon diffusion in the sample and weight change by oxidation with air. Through the mass difference (Î&#x201D;m), the ablation rate of the radius Î&#x201D;r/Î&#x201D;t was evaluated. For this evaluation, the density of the cylinder, defined by equation [2], was used and assumed to be equal at the equilibrium temperature. After the experiments, it was observed that only a melting in radial direction occurred. Local density differences can explain this according to the temperature

gradient in the boundary layer of the hot metal. The denser liquid will move downwards along the cylinder surface and inhibit a melting on the bottom surface of the cylinder, which is why the length of the cylinder remains nearly constant in all experiments. The density of the scrap đ??&#x2020;đ?&#x2019;&#x201D;đ?&#x2019;&#x201E;đ?&#x2019;&#x201C; is defined by equation [2] published by Miettinen in (19) and is dependent on the temperature (in Â°C) and the scrap composition. The scrap is assumed to be ferritic and the density for a multicomponent system (composition đ?&#x2018;Şđ?&#x2019;&#x160;,đ?&#x2019;&#x201D;đ?&#x2019;&#x201E;đ?&#x2019;&#x201C; in wt.-%) is, according to the considered elements i of the scrap, explained with equation [2]. (19) (2)

The compositions of the hot metal and scrap samples used in the experiments are listed in Tab. 1 and Tab. 2, respectively. The dissolution tests were executed with ULC scrap in carbon saturated hot metal (hot metal 1) at three different temperatures (experiment numbers 1 to 3). Experiment number 4 was executed with hot metal 2 (carbon content of 1 wt.-%), which has a composition close to the liquidus line of the ULC

at 1443 Â°C. The starting temperature of the experiment was assumed to be 1550 Â°C according to the calculation of the equilibrium temperature, which will be explained subsequently. According to Zhang and Oetersâ&#x20AC;&#x2122; model, a stagnation of the melting behaviour should occur when the driving concentration difference (%Cscrap - %Cliq) in equation [1] should become zero.

La Metallurgia Italiana - n. 11/12 2018

Attualità industriale Tab. 1 – Hot metal specification

Hot metal specification Hot metal 1

Hot metal 2

Carbon content [wt.-%]

4.58

Silicon content [wt.-%] Manganese content [wt.-%] Phosphorus content [wt.-%] Mass of hot metal [g] Starting temperature [°C] Experiment number [-]

0.37

0.1

0.63

0.161

0.07

0.015

320

345

1305

1370

1450

1550

Tab. 2 – Scrap specification

Scrap specification ULC scrap

S235JR

Carbon content [wt.-%]

0.002

0.1

Silicon content [wt.-%]

<0.001

0.0733

Manganese content [wt.-%]

0.05

0.479

Phosphorus content [wt.-%]

0.003

0.01

Mass of hot metal [g]

26.3

Starting temperature [°C]

Experiment number [-]

1-4

Pre-test

Pre-Tests for Verification of Starting Conditions Before the dissolution experiments, pre-tests were executed to verify the starting conditions for the individual dissolution tests. The starting temperature of the melt was measured with a thermocouple type B, which was directly submerged into the melt. This temperature is shown in Tab. 1 for experiments 1 to 4. In a preceding test series of this research work, the melting and dissolution behaviour of S235JR scrap was investigated. In the course of this investigation, S235JR specimens (see Tab. 2) were used for verification of the equilibrium temperature between the liquid metal and the submerged sample, which results from the heat exchange between hot melt and cold scrap. For this purpose, a thermocouple type S was located in a bore with a diameter of 1.7 mm in the cylinder centre and at a distance of 10 mm from the cylinder tip. The cold cylinder was submerged into the liquid hot metal and the temperature was monitored until the equilibrium between the melt and

La Metallurgia Italiana - n. 11/12 2018

the cylinder was reached. This temperature was always below the starting temperature of the hot metal, according to Tab. 1. The S235JR has a very similar enthalpy to ULC steel as a function of temperature. Fig. 2 shows these functions for ULC steel and S235JR, calculated with the FactSageTM FSstel and FactPS database. Accordingly, it was assumed that the equilibrium temperature measured for S235JR should be the same as for ULC steel. The equilibrium temperature was evaluated with a heat balance, based on equations [3] and [4]. Whereas mScrap and mhot metal are the masses of the scrap and the hot metal in [kg], Q is the heat flux in [W m-2], cp,hot metal is the specific heat capacity in [J kg-1 K-1], Tstart scrap/hot metal and Tequilibrium are the starting temperature of the melt or scrap and the equilibrium temperature in [K]. HScrap(T) is the temperature dependent enthalpy in [J kg-1]. For the determination of Tequilibrium, the heat fluxes of the scrap and hot metal must be equal. (20)

Industry news (3)

(4)

Fig. 2 - Enthalpy of ULC steel and S235JR DISCUSSION AND RESULTS The increase in the scrap core temperature based on the four investigated starting temperatures of the melt is shown in Fig. 3.

Fig. 3 - Measurement of the core temperature in the cylinder centre and equilibrium temperature 40

La Metallurgia Italiana - n. 11/12 2018

Attualità industriale As seen in Fig. 3, the equilibrium temperature is reached within 10 seconds. The higher the starting temperature, the steeper the temperature gradient is. The measurement also shows that the equilibrium temperature is 65 °C to 80 °C below the starting temperature of the melt. In Tab. 3 the mea-

sured starting and equilibrium temperatures of the hot metal and the scrap are listed and compared with the calculated equilibrium temperatures from the heat balance obtained from equations [3] and [4] as well as the data from Fig. 3.

Tab. 3 – Starting and equilibrium temperatures of the hot metal and the scrap Starting and equilibrium temperatures Starting temperature Starting temperature Equilibrium temperature Equilibrium temperature calculated[°C] measured [°C] melt[°C] scrap[°C]

Temperature drop measured [°C]

Experiment 1

1305

1239

1230

Experiment 2

1370

1301

1300

Experiment 3

1450

1377

1385

Experiment 4

1550

1452

1470

Based on the well-fitting results of the thermocouple measurements, the determination of heat transfer coefficient kmet according to equation [1] becomes possible. In the following phase diagram (Fig. 4) of the present ULC scrap, the measurement points according to the hot metal composition (magenta) and the temperature development of experiment 4 with a carbon content of 1 wt.-% are shown. The blue line indicates the way of the equilibrium temperature between 10 to 30 s.,

whereby the liquidus line is exceeded during this time. Such behaviour does not occur during the experiment with hot metal with 4.58 wt. % carbon. For each equilibrium temperature, the carbon concentration on the liquidus line (cliq) in the Fe-Fe3C phase diagram was determined. The terms of the carbon differences from equation [1] for each experiment are presented in Tab. 4.

Fig. 4 - Measured equilibrium temperature based on the present ULC scrap composition phase diagram La Metallurgia Italiana - n. 11/12 2018

Industry news Tab. 4 – Carbon differences based on equation [1] Carbon differences of equation 1 Equilibrium temperature [°C]

(%Cscrap - %Cliq) [wt.-%]

(%Cscrap - %CHM) [wt.-%]

Experiment 1

1230

-3.4651

-4.5763

Experiment 2

1300

-2.7467

-4.5763

Experiment 3

1385

-1.7612

-4.5763

Experiment 4

1470

-0.5930

-0.9980

The experimental determination of the ablation rate of the radius (Δr/Δt) for experiments 1 to 3, executed in hot metal

with a carbon content of 4.58 wt.-%, is presented in Fig. 5.

Fig. 5 - Ablation rate of ULC steel scrap in carbon saturated hot metal

At the beginning of the process, a positive ablation rate occurs. This phenomenon is attributed to a shell formation due to solidification of the liquid hot metal on the cold scrap surface. After a certain progressing time, the turning point is reached, the shell formation stops, and the melting of the shell starts. When the ablation rate turns negative, the melting of the mother scrap starts. The melting and dissolution is dependent on the equilibrium temperature, which is reached in this case, according to the measurements in Fig. 3, after 10 s. By using equation [1] the mass transfer coefficient for this

system is calculated. According to the negative sign in the equation, a positive heat transfer is expected if the melting of the mother scrap is in progress. In Fig. 6 the calculated mass transfers of the three starting temperatures are plotted. As mentioned, the mass transfer is negative during the growth and melting of the shell layer at the beginning of the process and becomes positive after 25 s dissolution time at the latest. According to the progress, it could be established that the mass transfer coefficient will increase slightly with the temperature after the equilibrium temperature has been reached.

La Metallurgia Italiana - n. 11/12 2018

Attualità industriale

Fig. 6 - Experimental mass transfer coefficient During the LD-process, the carbon content in the liquid melt decreases and the temperature increases in line with the exothermic reactions of oxidation. Therefore, the melting behaviour of ULC steel scrap was determined in modified hot metal with 1 % C. In Fig. 7 the mass transfer coefficient and

the ablation rate of experiment 4 are presented. According to the heat balance calculation and measurement the specimen temperature increases to 1415 °C after 10 s and 1470 °C after 30 s, which is just above the liquidus line in the quasibinary phase diagram..

Fig. 7 - Mass transfer coefficient and ablation rate of ULC scrap in hot metal containing 1 %C

La Metallurgia Italiana - n. 11/12 2018

Industry news The smaller carbon content in the hot metal results in a strong shell formation at the beginning of the process with a higher negative mass transfer coefficient in comparison to experiments 1 to 3, with higher carbon contents in the hot metal. However, under real process conditions, shell formation will not occur. The scrap will be, due to heat conduction, almost at the equilibrium temperature when those amounts of carbon are reached. This will result in a higher expected negative mass transfer coefficient with an approach to 0. For that reason, a stagnation of the melting behaviour, described by equation [1], where the driving of (%Cscrap - %Cliq) would become zero or negative could be neglected. An explanatory argument is therefore that in this measurement just above the liquidus line, a dissolution of the ULC steel with high mass transfers occurs, which would not be stopped immediately when reaching the liquidus line or ends in an abrupt negative mass transfer coefficient when the temperature falls below the liquidus line. CONCLUSION This publication presents the results of laboratory scale experiments on the dissolution and melting behaviour of ULC steel scrap in hot metal with two different carbon concentrations. For the investigation, four different experiments with various hot metal temperatures were performed. For the evaluation of the hot metal mass transfer coefficient of the diffusive scrap melting process, a literature-based equation was used. To get the true process equilibrium temperature, a pre-test with a thermocouple in the core position of the specimen was executed, resulting in a temperature drop of more than 65 °C. Based on these measurements, the carbon concentrations on

the liquidus line of the current available Fe-Fe3C-Mn phase diagram were evaluated. With the measurements, the ablation rates of the radius were determinable and furthermore, the mass transfer coefficients were defined. What is mentionable is that the mass transfer coefficient is slightly dependent on the temperature. At the beginning of the process a shell formation occurs, which results in a negative mass transfer. If the carbon content in the melt decreases and the temperature increases, an increase in the mass transfer is observable. This results in a high melting rate of ULC scrap just above the liquidus line which would not turn into an abrupt negative mass transfer coefficient if the liquidus line were undercut. In summary, the outcomes of this work clearly indicate that the actual temperature and melt composition have a strong impact on the melting and dissolution behaviour of ULC steel scrap. However, the difference in the dissolution behaviour of hot metal with 4.58 and 1 wt.-% carbon is not fully explainable. More research work must be done to investigate and describe the melting and dissolution of scrap for the conditions in the LD process. ACKNOLEDGMENT The authors gratefully acknowledge the funding support of K1-MET GmbH, metallurgical competence centre. The research programme of the K1-MET competence centre is supported by COMET (Competence Centre for Excellent Technologies), the Austrian programme for competence centres. COMET is funded by the Federal Ministry for Transport, Innovation and Technology, the Federal Ministry for Science, Research and Economy, the provinces of Upper Austria, Tyrol and Styria as well as the Styrian Business Promotion Agency (SFG).

REFERENCES [1]

Turkdogan ET. Fundamentals of steelmaking. The institute of materials. London. 1996.

[2]

Ghosh A, Chatterjee A. Ironmaking and Steelmaking theory and practice. PHI Learning Private Limited. Delhi. 2015.

[3]

Asai S, Muchi I. Effect of scrap melting on the process variables in LD converter caused by the change of operating conditions. Transactions ISIJ. Vol. 11; 1971. p. 107 – 115.

[4]

Gaye H, Wanin M, Gugliermina P, Schittly Ph. Kinetics of scrap dissolution in the converter. Theoretical model and plant experimentation. 68th Steelmaking conference, AIME, Detroit, USA. 1985. p. 91 – 103

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Attualità industriale [5]

Isobe K, Maede H, Ozawa K, Umezawa K, Saito C. Analysis of the scrap melting rate in high carbon molten iron. ISIJ. Vol. 76; No.11, 1990. p 2033 – 2040.

[6]

Zhang L, Oeters F. Schmelzen und Mischen von Legierungsstoffen in Stahlschmelzen. Verlag Stahleisen GmbH. Düsseldorf. 2012.

[7]

Szekely J, Chuang YK, Hlinka JW. The melting and dissolution of low-carbon steels in iron-carbon melts. Metallurgical Transactions; Vol. 3; 1972. p. 2825 – 2833.

[8]

Shukla AK, Deo B, Robertson DGC. Scrap Dissolution in Molten Iron Containing Carbon for the Case of Coupled Heat and Mass Transfer Control. Metallurgical and Materials Transactions B; Vol. 44; 2013. p. 1407 – 1427.

[9]

Den Hartog HW, Kreyger PJ, Snoeijer AB. Dynamic model of the dissolution of scrap in BOF process. C.R.M. Rep. Vol. 37; No. 12, 1973. p. 13 – 22.

[10] Kawakami M, Takatani K, Brabie LC. Heat and Mass Transfer Analysis of Scrap Melting in Steel Bath, Tetsu-to-Hagané. Vol. 85; No. 9, 1999. p. 658 – 665. [11] Kruskopf A, Holappa. Scrap melting model for steel converter founded on interfacial solid/liquid phenomena. Metallurgical Research and Technology. Vol. 115; 2018. p. 201 – 208. [12] Sethi G, Shukla AK, Das PC, Chandra P, Deo B. Theoretical Aspects of Scrap Dissolution in Oxygen Steelmaking Converters. AISTech 2004 Proceedings Volume II, Nashville, USA. 2014. p. 915 – 926. [13] Lytvynyuk Y, Schenk J, Hiebler M, Sormann A. Thermodynamic and Kinetic Model of the Converter Steelmaking Process. Part 1: The Description of the BOF Model. Steel Research int. Vol. 85; No. 4, 2014. p. 537 – 543. [14] Bundschuh P. Thermodynamische und kinetische Modellierung von LD-Konvertern. Dissertation. Montanuniversität Leoben, Austria. 2017. [15] Penz FM, Bundschuh P, Schenk J, Panhofer H, Pastucha K, Paul A. Effect of Scrap Composition on the Thermodynamics of Kinetic Modelling of BOF Converter. 2nd VDEh-ISIJ-JK Symposium, Stockholm, Sweden. 2017. [16] Penz FM, Bundschuh P, Schenk J, Panhofer H, Pastucha K, Paul A. Impact of Carbon, Silicon and Manganese contents on the dissolution and melting behaviour of scrap in a dynamic BOF model. 3rd European steel technology and application days (ESTAD), Vienna, Austria. 2017. [17] Zarl M. Development and evaluation of a BOF pre-processor model. Master Thesis. Montanuniversität

Leoben, Austria. 2017.

[18] Boychenko B, Okhotskiy V, Kharlashin P. The converter Steelmaking. Dnipro-VAL, Dnipropetrovsk. 2006. [19] Miettinen J. Calculation of solidification-related thermophysical properties for steels. Metallurgical and materials transactions B. Vol. 28 B; No. 4, 1997. p. 281 – 297. [20] Seshadri V., Tavares Parreiras R., Da Silva C.A., Da Silva I.A. Transport phenomena: fundamentals and applications in Metallurgical and Materials Engineering. ABM, Sao Paulo. 2011.

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Industry news Effect of converter dart design aspects on its settling behavior edited by: J. van Oord, H.J. Visser A slag dart is used in the steel industry in the tapping of BOS-converters to prevent contamination of the tapped molten steel with carryover slag. The performance of the dart has become increasingly important for higher alloyed steel grades as downgrades are costly. A mathematical model was developed to quantify the major forces on the dart that drive settling into the tap hole. This model was tuned by various measurements in a dedicated water model to quantify the forces on the dart especially during the settling stage. Primary dart geometry parameters were varied. The results show how the forces drive the settling process. Experimental fitting coefficients were derived for pressure and shear driven forces. Also, the effect of tap hole entrance wear was studied. Based on the results, plant trials were carried out using different dart designs. The trial results confirm the model. The mathematical model, the water model experiments and plant trials are discussed.

KEYWORDS: CONVERTER – TAPPING – SLAG DART – ANALYTICAL MODEL – WATER MODEL – VALIDATION – PLANT TRIALS

J. van Oord Tata Steel, Research & Development, IJmuiden, the Netherlands

H.J. Visser Tata Steel, IJmuiden BOS plant No.2, the Netherlands

INTRODUCTION In the final stage of converter tapping, normally when around 80% of the steel has been tapped, a slag dart (see Fig 1.) is dropped into the tilted converter from above the tap hole. Conventional darts consist of a head and an attached tail. The tail is primarily intended as a guide to ensure that the dart settles into the tap hole. The slag dart head acts as a floating valve throttling down the flow at the end of the tap, thereby preventing excessive slag carryover. The aim is to

have dart operation prior to the phase of full bore flow slag carry over. The aim is also to have a short duration of the reduced flow phase. The performance of the slag dart has become increasingly important for higher alloyed steel grades as downgrades are costly. Still, hardly any detailed research has been carried out on the effect of dart design aspects on the settling process. Some investigating numerical calculations were reported by Simonnet et al. (1) although these were not validated and not supported by plant trials.

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AttualitĂ industriale

Fig. 1 - Left: photograph of slag dart. Right: schematic illustration of a slag dart floating in steel (red) and slag (green) at the end of a tap prior to restricting the flow.

Proper dart operation depends on many factors. A poor launch of the dart can result in a mishit, thick and solid slag can cause deflection of the dart but also conditions close to the tap hole entrance in the converter in case of a proper launch can play a role in a failure of operation. The slag depth, the remaining steel level, the converter wear stage, the tap hole wear stage will all vary and have their effect on the settling. Therefore, it is quite common to observe a varying dart performance in operational practice. Apart from the aforementioned practical variations, the dart design itself will also play a role. So to tune the dart design to the specific conditions in the converter, a general understanding of the fluid flow related settling of the dart can help to optimise the dart performance. The dart settling during tapping is a dynamic process. The dart head is expected to float partly in the steel bath and partly in the slag layer as illustrated in Fig 1. The steel bath level lowers while the steel is tapped into the ladle and the dart head lowers towards the hot face of the tap hole. Also the dart tail will gradually enter deeper into the tap hole at the end of the tap. Thus during the final draining phase the fluid flow field close to the dart head and tail changes dramatically, as also described by Simmonet (1). Hence the forces related to fluid flow on the dart change accordingly as the dart settles. Based on this phenomenological description of the fluid flow, it becomes clear that the dart design (head shape, head size, tail dimensions) but also the tap hole shape â&#x20AC;&#x201C; wear stage â&#x20AC;&#x201C; and dimensions are crucial for the settling process. To study the dart settling process in detail, a parametric model to calculate the gravity, buoyancy and fluid flow forces has been developed. This model was fed with empirical data

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derived from experiments in a dedicated water model reported here. Subsequently, measurements were done to quantify the effect of different tap hole wear stages to simulate operational conditions. ANALYTICAL MODEL During the settling stage of the dart, the interaction of the dart with the turbulent flow of the steel draining from the converter results in a complex motion of the dart. The dart tail and head will move lateral in some way and hence the solid wall boundary conditions for the flow field will vary. Therefore, the fluid-body interaction itself is sufficiently complicated to make detailed dynamic numerical simulations tedious and can only be done after some major simplifications. The lowering level of the steel and the influence of the slag phase further adds complexity. Still, careful consideration of the phenomenology of the process reveals that the settling process is mainly driven by the gradual change in the increasing downwards acting fluid flow forces in combination with increasing buoyancy force as the dart head is pulled down further into the steel. Therefore it seems appropriate to describe the settling process by a basic parametric model only describing the three major forces acting on the dart: gravity, buoyancy and the fluid flow force. Taking a detailed look at the forces separately then shows how typical dart design parameters and process parameters will affect the settling process. Note that the darts natural equilibrium floating orientation is (almost) horizontal as the head contains the most mass so the center of gravity is above the center of buoyancy (see (2)) and only prevented when the tail cannot move upright due to the

Industry news confinement of the tap hole geometry. Proper setting of the tail into the taphole at dart release is therefore essential. In the analysis described hereafter, it is assumed the orientation of the dart is vertical as shown on the right in Fig 1. Gravity force A dart is usually made up of two distinct parts, the head and tail, made of different materials. Usually, the head size and head density are selected by plant personnel for the specific converter and tap hole. The selection is usually done by con-

sidering hit rate and the duration of the reduced flow phase. As the head dimensions and or the head density changes the total gravity force, hence the weight changes. The shape and density of the tail are usually considered fixed but are altered here to vary the fluid flow forces. Using a parametric approach in calculating the total dart weight enables to perform studies on specific dart designs. The total gravity force in its basic form, treating head and tail separately, in the parametric model is: (1)

Buoyancy force Assuming the dart partly floats in steel and partly in slag (see

Fig. 1), the buoyancy force on the dart is composed of two distinct parts: (2)

where the volumes are the displaced volumes of slag and steel respectively depending on the submersion of the dart in slag and in steel. A parametric model based on basic dart design parameters then gives a detailed description of the buoyancy force as a function of dart submersion depth. From Fig. 1 it follows that this function should include dart head shape parameters, tail parameter, steel level and slag level. In the absence of slag the buoyancy force follows a curve as shown in Fig. 2. After a gradual linear increase from when the

tip of the tail enters the steel, the curve follows a steep incline when the head starts to submerge into the steel. Note that complexity is added to the curve if a slag layer is included in the calculation. Then, the buoyancy force will commence to act on the dart once the dart tail tip starts enters the slag and increases at a steeper slope once the tail tip reaches the steel. The result is that the dart will reach a higher equilibrium level with slag present.

Fig. 2 - Schematic illustration of the buoyancy function of a typical dart configuration. Bottom figure: dart radius plotted as a function of vertical position. Top: buoyancy curve (expressed in replaced mass) of a typical dart in steel (7000 kg/m3) as a function of its submersion depth. The red dotted line indicates the floating position for a 32 kg dart. 48

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AttualitĂ industriale Fluid flow forces As the converter is emptied during tapping, the dart head will get closer to the hot face of the tap hole and the tail tip will reach deeper and deeper into the tap hole. Initially, the forces will act mainly on the tail but as the remaining steel level lowers, the forces on the head will come into play. It is the aim of this research to quantitatively describe the change in fluid

flow force depending on variations in process conditions and the geometrical design of the dart and tap hole. Starting off, following basic fluid mechanics theory on pipe flow and drag forces (2, 3) the force on the dart tail can be divided in two different contributions: a shear force on the wall of the tail and a dynamic drag force at the bottom of the tail (again: assuming the tail is vertical in the tap hole). (3)

For the shear force this is, in more detail, written as: (4)

where đ?&#x153;?đ?&#x2018;¤đ?&#x2018;&#x17D;đ?&#x2018;&#x2122;đ?&#x2018;&#x2122; is the tail wall shear stress and đ??ˇ and đ??ż are the dart tail diameter and length respectively , đ?&#x2018;Ł a is the characteristic velocity and đ?&#x2018;&#x201C;đ?&#x2018; â&#x201E;&#x17D;đ?&#x2018;&#x2019;đ?&#x2018;&#x17D;đ?&#x2018;&#x; a is the empirical friction factor

of the dart tail. For the dynamic drag on the tip, the relation becomes:

where đ?&#x2018;? is the local pressure at the tip and đ??śđ?&#x2018;Ąđ?&#x2018;&#x17D;đ?&#x2018;&#x2013;đ?&#x2018;&#x2122; is the drag coefficient. Combining Eq. [4] and Eq. [5] gives the following relationship for the fluid flow force on the dart tail in steel

showing a shear and pressure depending components of the force on the dart tail.

(5)

(6)

It appears the force is quadratically dependent on the characteristic velocity in the tap hole and depends on the two empirical coefficients, based on shear and dynamic drag. Applying the equations given here, basically the force balance of gravity, buoyancy and fluid flow force, for any dart tail configuration can be calculated. A parametric model was built to calculate at which equilibrium floating position of the dart the buoyancy force would be equal to the sum of the gravity and the fluid flow forces using basic algebraic solving tools. The model makes use of parameterized buoyancy function that calculates the displaced volumes of steel and slag for a given dart geometry at any position as was visualised in La Metallurgia Italiana - n. 11/12 2018

Fig. 2. The remaining unknowns in the model cannot be found in literature. Based on similarity considerations however, they can be determined empirically using a scaled water model as described in the next section. SET-UP Using the approach of separating the fluid flow force from the other forces, enables us to use a water model to determine the effective drag friction factor on the tail. Thus, a dedicated experimental facility was built to get full quantitative insight in the forces on the dart. 49

Industry news The experimental facility is shown in Fig. 2. The central part of the set-up, is a square transparent box of 500 mm x 500 mm. In the bottom of the main section, water flows into a straight tap hole of 40 mm diameter. The central part is separated from the return lines using two buffers with vertically adjustable side walls. The system, including the buffers, is filled up to a certain level and the water is circulated by means of a pump capable of generating a flow that results in an average

velocity of up to 2.5 m/s in the tap hole. The flow is guided from the buffers beneath the sidewalls and thus distributed homogeneously into the central section. In the circulation line a flow meter is placed to measure the flow rate in the tap hole. Directly above the tap hole, the dart is connected to a load cell that measures the vertical force on the dart. The dart is fixed and thus any dart location during the final phase of the tapping process can be studied in a quasi-static situation.

Fig. 3 - Schematics of the dart forces measurement rig â&#x20AC;&#x201C; side view. The dart is fixed to a load cell. The height of the dart can be adjusted to various positions. The water level can also be adjusted. The flow rate is adjusted by means of varying pump settings and the flow rate through the tap hole is measured by means of a flow meter. Tap hole and dart geometry One dart head size and three different tail diameters were used. The head was made of wood and the interchangeable

tails made of wood with a metal core used for the connection with the head. The geometry parameters of the model dart head and of the different tails are given in Tab. 1.

Tab. 1 â&#x20AC;&#x201C; Model dart geometry in (mm). DART PARAMETERS

Top head diameter

76.3

Bottom head diameter

24.0

Height head

43.3

Diameter tail SD1

9.3

Diameter tail SD2

12.2

Diameter tail SD3

15.2 310

Tail model length

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AttualitĂ industriale In all experiments the water level is set to the tipping point situation: flush with the top of the dart head. This is the situation from where a free floating dart will settle in case the fluid flow force just exceeds the total buoyancy force. If measurements are carried out with the dart (and water level) at different vertical positions, a good insight is obtained in the varying forces during dart settling in operational practice. It is known that in plant operation the hot face of a tap hole wears during its lifetime (80-100 heats). The hot face smoothens, and wear of the inner tap hole results in a slowly

increasing diameter. Thus the flow rate increases over the tap hole lifetime and the total tapping time reduces. Furthermore, the drainage time â&#x20AC;&#x201C; the time the flow is restricted due to a settled dart â&#x20AC;&#x201C; gets shorter. Apparently, the dart settles at a later stage leaving less steel to drain from that moment on. To get insight in this process, the tap hole wear stage is varied by using different tap hole inserts in the model. The different stages are visualized in Fig. 3. Two tap hole inserts (straight and smooth entrance) were used in combination with two different bottom plates resulting in four tap hole wear stages.

Fig. 4 - Schematics of the different tap hole configurations. Measurement method With each measurement the flow rate is increased step by step and the load on the dart is measured for 30 seconds. As the top of the dart head is flush with the water level, the force measured at zero flow results from the buoyancy and gravity force. The increase of the flow results in an increase in load based on the average velocity in the tap hole. The average velocity is considered as characteristic for the fluid flow force.

As the dart does not move vertically during the experiments, the zero flow load result acts as an off-set from the load due to flow. Thus, the remaining total fluid flow force on the dart is measured. A measurement combines the drag and pressure effects, and thus can only be represented by means of a total drag coefficient: (7)

Here the total (measured) drag coefficient CD is used, the velocity in the tap hole is used as the characteristic velocity and the dart tail diameter is used as a characteristic dart design parameter. From Eq. 7 it follows that the load measurement for a certain dart location and tap hole geometry is proporLa Metallurgia Italiana - n. 11/12 2018

tional to v2. The separate drag coefficients for tail and head can be derived by experiments using a tail only and including a dart head at a later stage. The relative contributions of drag and pressure on the tail can be derived from measurements using different tail lengths. 51

Industry news

Fig. 5 - Measurement results for different dart tail diameters. RESULTS General result Eq. 7 is now used to plot the measured loads for different dart tail diameters resulting in the trends shown in Fig. 5. The results are plotted in accordance with Eq. 7 to find the total drag coefficient. The results are shown for the configuration with the level at 100 mm above the hot tap hole entrance and the different diameter tails attached to dart head. It shows that a similar linear increase in drag coefficient is found if the

tail diameter is used as the characteristic length scale and the load is indeed linearly proportional to v2. The results are obtained in line with the standard method for the determination of a drag coefficient of an object in a flow (2,3). The drag coefficient derived from the slope measures 2.3 Âą 0.1. Note that for this configuration the tail has entered into the tap hole for 186 mm and thus the result includes both a dynamic drag as well as a wall shear effect.

Fig. 6 - Drag coefficients for different dart tail depths into the water using the T1 tap hole configuration.

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AttualitĂ industriale Tail only experiments The derived drag coefficient can be evaluated as function of other parameters. Thus, given the theory described above and varying the tail depth enables to separate the dynamic drag and the wall shear contributions to the total drag coefficient. In Fig. 6 the drag coefficient for different tails is shown. Besides a few outliers, an increasing trend is clearly observed for larger tail depths. The coefficient ramps up from roughly 1.8 to 2.8 for the largest measured depth. The results show that given the practical range of tail lengths and tail diameter, the shear effect is less strong than the pressure effect. Once the tail tip penetrates the tap hole, the drag force ramps up quickly and then, after further penetration, gradually increases. By looking at eq. [6] the wall shear effect includes a length

to diameter ratio (đ??żâ &#x201E;đ??ˇ). This ratio therefore should increase quicker for the thinner tail than for the others. This effect is not recognized clearly from the data, but can also not be ruled out. It seems that inaccuracies in the determination of the drag coefficient still allow that a linear trend is valid for use in a practical model. The range of diameters of the tails is too small with respect to the accuracy of the current results - the length to diameter ratios (đ??żâ &#x201E;đ??ˇ) of the thicker and thinner stem only differ roughly 25% with respect to the middle diameter. Uncertainty in the roughness of the tail surface will also lead to variations in the resulting value. Combining the results of the datasets the following values are found while fitting the data with a linear trend:

(8)

(9)

This result can be used to calculate how far the fluid flow force will pull the dart deeper into the tap hole once the tail reaches the tap hole entry. Depending on the mass and the geometry of the dart, the dart will find a natural floating position where the total buoyancy force equals the sum of the gravitational and fluid flow forces. Effect tap hole wear The drag coefficients were determined for different stages of tap hole wear, similar to the aforementioned procedure, i.e. just on the characteristic velocity in the tap hole, and scaled to the dart tail diameter. In Fig. 7 the resulting drag coefficients are plotted as a function of the water level in the model. The water level represents both the remaining level in the model and the distance of the dart head to the tap hole entrance. Clearly for, TS1 and TS2, the two configurations representing

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relatively new tap holes, the downward pulling forces caused by the fluid flow become up to over 5 times larger than the force on the tail only once the dart head gets close to the tap hole entrance. It should be noted that a few data points are missing for the TS1 configuration since it was not possible to lower the dart any deeper given the wear stage configuration (see also Fig. 3 and Fig. 4). This result helps to understand the actual behaviour of the dart over the tap hole lifetime. In the early stages of the tap hole lifetime the fluid flow force on the dart head will pull the dart down from higher levels thus initiating early settling. This results in longer drainage times. The strength of this effect is in operational practice depends on the particular tap hole wear stage and the dart design parameters.

Industry news

Fig. 7 - Drag coefficients as a function of the water level in the model for the SD2 tail configuration.

PLANT TRIALS Using the parametric model, different dart designs can be evaluated based on operational process parameters. In operational practice in IJmuiden, the tap hole is tapered, and since the model is based on a straight tap hole, the response of a certain dart design is considered with respect to that simplification. It is expected that the dynamic drag and wall shear will increase deeper into the tap hole in line with the increasing velocity. In the following, the use of the model to design darts for a series of plant trials is given. The dart settles once the fluid flow results in a sufficiently large force overcoming the total buoyancy force. If this â&#x20AC;&#x2DC;tipping pointâ&#x20AC;&#x2122; is not reached, the dart will remain floating at a level away from the tap hole, and slag will flow unrestricted into the tap hole prior to end of tap. Using the described model results in conjunction with model calculations of the fluid flow forces in the plant it follows that, given a dart design, the fluid flow on the tail itself may not be sufficient to pull the dart into the tap hole (see Tab. 2 further on). Fig. 7 indicates that settling may still happen as long as the dart head gets sufficiently close to the hot face of the tap hole such that the drag force (coefficient) increases. This shows why reliable settling becomes critical once the hot face of the tap hole wears out, a problem that is indeed observed in plant operation at BOS No.2 in IJmuiden. The reliability of the darts settling is based on its hit rate:

which is tracked each heat based on the operatorsâ&#x20AC;&#x2122; judgement of his observation of the flow at the end of the tap. A clearly restricted flow at the end of tap indicates a proper dart operation. The model indicates that an increased tail diameter results in an increased fluid flow force. This indicates that the dart design can be tuned such that settling still occurs reliable for worn out tap holes. To verify this, a series of plant trials using batches of different dart types was done using the model to define the range of parameters and to evaluate the results. Two different head types were used having slightly different volumes, named 1 and 2. Two different extremes in head density specifications were used, high density, 3.9 kg/liter (H) and low density, 3.3 kg/liter (L), by using different types of refractory. 6 different tail diameters, still considered suitable in practice, were used, indicated as for example a diameter of 51 mm for the 1L51 dart. The parameters of the dart types are listed in Tab. 2. In total at least 20 up to 240 darts were used in the different batches. Also, all darts were used at old tap holes, i.e. at least at over 60% of their usual lifetime.

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AttualitĂ industriale Tab. 2 â&#x20AC;&#x201C; Configurations of the trial darts.

TRIAL DART PARAMETERS Rank

Type

diameter tail (mm)

volume tail (liter)

total volume (liter)

buoyancy (N)

measured weight (N)

fluid flow force (N)

remaining weight (N)

1H63

3.13

11.63

799

411

333

2H55

2.39

11.39

782

412

259

111

1L58

2.65

11.15

766

347

285

133

1H48

1.82

10.32

708

370

202

136

1L51

2.05

10.55

725

338

226

161

1H40

1.26

9.76

670

340

145

184

Using the model, the buoyancy force and the fluid flow force for the tail fully drawn into the tap hole were calculated. Only the volume of the head and the tail are used. It is assumed that this part of the dart governs the buoyancy as it is completely submerged into steel as the dart settles. The stem part above the head is neglected since it is located in slag and its density is close to that of slag. The darts were weighed to calculate the resultant force required to overcome the buoyancy force for dart settling. From the tap hole geometry and total tapping time in plant at a half lifetime tap hole, the characteristic velocity was calculated as 3.8 m/s. This value was used to evaluate the reliability of settling for different dart designs, using eq. (6). A tail length of 0.7 m was used as the drag force is expected to act only on the tail in the narrow part of the worn out tap hole. The head effect is not taken into account in this calculation but is considered play a role in practice. Also the exact tap hole lifetime was not considered. Using these results, the darts were ranked from low to high resultant forces: the excess weights to overcome the buoyancy. Some excess value is present for all dart designs although the values range from 184 to 55 N (translating to 18.8 to 5.6 kg). The lower limit was chosen be a positive value to prevent direct shut-off of the tap hole at dart release. The darts were aimed at investigating the settling behaviour at older tap holes, so variations in the density and diameter were expected

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to have the largest contribution on the performance. The average hit rate results of the trial darts is shown in Fig. 9. In plant, dart operation is subjected to variations that have nothing to do with the basic physics described here, although it still drives the performance. This complicates evaluation of trials using small batches of darts but the results shown in Fig. 9, however, show a clear trend that is in line with the predictions given by the model. It should be mentioned that some clear abnormalities observed in operational practice are removed from the data. Only the batch of the 51 mm tail is not in line with the rank trend but the deviation can still be explained on the small size of the batch and variations in hit rate that generally occur in practice. Note that the trend covers both the effects of the dart density and a wide range of tail diameters, indicating the effect of both parameters are described properly by the model. It should also be mentioned that the observed drainage times were in line with the expectations: markedly longer on the higher ranked darts than on the lower ranked darts. The model is therefore considered to properly cover the dominant physics in practice and can be used to select a dart geometry that suits a certain tap hole configuration. In the end, the final choice of design still depends on decisions driven by a trade-off between process time (a desired short duration of the drainage phase) and a high hit rate.

Industry news

Fig. 8 - Hit-rate of different trial darts as a function of the ranking based on resultant force. CONCLUSIONS AND RECOMMENDATIONS A parametric model of the forces on a dart in a taphole was built. To accurately describe the fluid flow forces for different dart design parameters and process conditions, measurements were done on a dedicated water model. Values for the wall shear friction factor and the dynamic drag coefficient were determined. A series of plant trials confirm the validity of the model. The model can give insight in the behaviour of the dart performance in general or in typical behaviour of a dart on a specific heat. It can also be used in the selection process for

choosing a dart design to obtain the optimum in the trade-off between hit rate and drainage time. ACKNOWLEDGEMENTS Thanks to Chris Smith, Tom Murphy and Jeff Lawrence from Teesside’s Materials Processing Institute for their inspiring cooperation and operating the water model and Fokke Zwaan and Rob Plug from the IJmuiden Centre of Expertise and BOS No.2 Steel Plant production for their coordination and assistance during the plant trials.

REFERENCES [1]

2011, 27th June – 1st July; M. Simonnet, S. Gauthier, JF. Domgin, P Gardin, P. Blumenfeld, Behaviour and improved design of dart system in BOF, Process modelling in BOF and AOD steelmaking, METEC InSteelCon 2011.

[2]

2014; Fluid Mechanics and Hydraulics. Schaum’s Outlines. Ranald V. Giles, Jack B. Evett, PhD, Cheng Liu.

[3]

Batchelor, George (2000). An introduction to fluid dynamics. Cambridge Mathematical Library (2nd ed.).

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premio

edizione 2019

aldo

Dacco

L’AIM è lieta di indire il bando per l’edizione 2019 del prestigioso Premio Aldo Daccò, con l’obbiettivo di stimolare i tecnici del settore e contribuire allo sviluppo e al progresso delle tecniche di fonderia e di solidifiazione con memorie e studi originali. L’Associazione invita tutti gli interessati a concorrere al Premio “Aldo Daccò” 2019, inviando a mezzo email (info@aimnet.it), il testo di memorie inerenti le tematiche fonderia e solidifiazione, unitamente al curriculum vitae dell’autore concorrente, entro il 30 novembre 2019. Saranno presi in considerazione e valutati i lavori riguardanti le varie tematiche di fonderia e di solidifiazione, sia nel campo delle leghe ferrose che in quello delle leghe e dei metalli non ferrosi. Il premio, pari a Euro 3500 lordi, è offrto dalla Fondazione Aldo e Cele Daccò, istituita dalla signora Cele Daccò per onorare la memoria del marito Aldo Daccò, uno dei soci fondatori dell’AIM e suo encomiabile Presidente per molti anni. Le memorie verranno esaminate da una Commissione giudicatrice designata dal Consiglio Direttivo, il cui giudizio sarà insindacabile. Nel giudicare, la Commissione terrà conto, in particolar modo, dell’originalità del lavoro e dell’argomento in relazione alla reale applicabilità dei risultati. Non sono ammesse candidature da chi abbia già ottenuto riconoscimenti, anche per lavori diversi, dalla Fondazione Aldo e Cele Daccò per la ricerca scientifica. Le memorie premiate e quelle considerate meritevoli di segnalazione, potranno essere pubblicate sulla rivista La Metallurgia Italiana. La cerimonia di premiazione avrà luogo in occasione dell’Assemblea dei Soci AIM, che si terrà a Milano nella primavera del 2020. Per informazioni e candidature: Associazione Italiana di Metallurgia Via Filippo Turati 8 · 20121 Milano Tel. 02-76397770 · E-mail: info@aimnet.it

Experts’ corner Reaction near Hotspot in Refining of Molten Iron during Supplying Top-blown Oxygen and Powder Additives edited by: Yu-ichi UCHIDA Faculty of Engineering, Nippon Institute of Technology, JAPAN Hotspot is defined as the region where oxygen gas is supplied onto/into molten iron. An information on the reaction near hotspot with top-blown oxygen is important for better control of oxidation refining of molten iron. The present analysis aims to achieve fundamental understanding of the reaction near hotspot b y reevaluation of the results of the previous laboratory-scale experimental by the present author, in which oxygen gas and powder additives were supplied onto carbon-saturated molten iron at around 1603K. The temperature at the hotspot measured with an optical fiber thermometer was up to about 2700K and relatively higher than those through top lance compared at the same oxygen flow rate, which is considered not to be underestimated compared to another method. The measured temperature was lowered when powder additives such as quicklime and calcium carbonate was supplied to the hotspot. The slag sampled “near” the hotspot had higher basicity and T.Fe, compared to the slag sampled near the crucible wall. Phosphorus distribution ratio and resulting phosphorus content was evaluated from the slag composition after the present author and Miyata et al (calculated). Based on the analysis, slag formed at hotspot with CaO blasting would have originally high potential for phosphorus removal resulting from high FeO content, and would still hold dephosphorizing capacity after moving to and consuming FeO “near” hotspot. KEYWORDS: REFINING – MOLTEN IRON – HOTSPOT – OXYGEN – DEPHOSPHORIZATION – SLAG

INTRODUCTION In oxygen refining of molten iron, reaction taking place in converter during oxygen blowing is rapid and complex. One of the reason is considered to be formation of slag/metal emulsion generated by supersonic oxygen jet. The other noteworthy event is formation of hotspot. Hotspot is defined as the region where oxygen gas is supplied onto/into molten iron. A comprehensive model on decarbulization in converter including such characteristic phenomena was proposed, which comprises of the emulsion zone and the impact zone (corresponding to hotspot)(1-3). With respect to phosphorous refining, recently a syste58

matic work and analysis has been presented based on the data obtained in a 6 ton pilot converter (IMPHOS project) (4-5). In order to develop such analysis on dephosphorization in more detail, further information might be helpful about what happens in hotspot. Another concern on phosphorous refining of molten iron is reconsideration of CaF2 utilization. Among a number of components in steelmaking slag, fluorspar (CaF2) is recognized as a typical fluxing agent for dephosphorization operation (6). Addition of CaF2 is beneficial for reduction of the viscosity and melting temperature of slag. This makes slag more fluid and promotes

equilibrium at the slag-metal interface. However, spent slag containing CaF2 cannot generally be utilized as roadbed or landfill material due to elution of fluorine ions from the slag. Thus, for better environmental control, development of a refining method without using fluorspar is required. The key technique for CaF2-free refining is promotion of the formation of molten slag. One promising candidate is addition of powder in the hotspot with high temperature and high oxygen potential. Several reports have been presented on the effect of supplying auxiliary material to the hotspot on the dephosphorization reaction (7-10). For better control of La Metallurgia Italiana - n. 11/12 2018

Scenari phosphorous refining, the reaction near the hotspot should be more understood. This paper is focused on phosphorous removal from molten iron, aiming to obtain fundamental information on the reaction near the hotspot during oxidation refining with top-blown oxygen. An investigation of the temperature at the hotspot and the slag composition near the hotspot was carried out in a laboratory-scale experiment in which oxygen gas and auxiliary materials were

supplied to the molten iron. TEMPERATURE AT HOTSPOT Overview Hotspot in oxygen refining of molten iron is recognized as an area where oxygen gas is impacted at surface of molten iron. Hotspot is characterized by high temperature and high oxygen potential. A schematic illustration of hotspot under top blowing oxygen in converter is shown in Fig.1. One can calculate the

area in which oxygen reached the surface of molten iron using jet trajectory. Since it is difficult to define the volume of hotspot, there seems no designated concept about them. In order to obtain typical aspect about hotspot with special interest, the measurement of temperature at hotspot has been carried out by many researchers. Heat balance at hotspot would be given as below;

(1)

Here ∆HFeO and ∆HFeO are heat of formation of CO and FeO (MJ/Nm3), ƞC and ƞFe are efficiency of for-mation reaction of CO and FeO at hotspot (-), Fo2 is oxygen flow rate (Nm3/min), Qgas , Qconv

and Qrad are heat transfer by sensible heat of gas, convection and radiation (MJ/min) respectively. One can understand that hotspot temperature would increase with oxygen

flow rate. On the other hand, heat loss by radiation would be enhanced with the increase of hotspot temperature itself.

Fig. 1 - Schematic illustration of hotspot under top blowing oxygen in converter Temperature at hotspot has been reported by many researchers with the measurements thorough top lance or bottom tuyere (11-19). The measured results of hotspot temperature are shown against the oxygen flow rate (in Nm3/min/t) in Fig.2. The condition and methodology for some of the measurements through La Metallurgia Italiana - n. 11/12 2018

top lance are listed in Tab. 1. The data given in circle are obtained thorough top lance, in diamond by FIMT (Fiber In Metallic Tube), and in other symbols through bottom tuyere. Though the metal weight at the experiments are different, one can see the general tendency that the hotspot temperature increases with

the oxygen flow rate. The tem-perature measured through bottom tuyere is relatively higher than those in another experiments compared at the same oxygen flow rate, probably due to the condition that the hotspot at the tip of bottom tuyere is sur-rounded by molten iron and relatively low heat loss by radiation. 59

Experts’ corner Tab. 1 – The condition and methodology for the measurements under top blowing oxygen

Researcher

Metal weight (kg)

(%Cscrap - %Cliq) [wt.-%]

Matsui

200

0.4-0.8

Koch Nakamura Kawakami

40 50 12000

4.1-9.2 1.4-4.0 9.9*

Uchida

0.5-1.5

(%Cscrap - %CHM) [wt.-%] Radiation thermometer + Optical fiber 2 colour pyrometer 2 colour pyrometer Spectrometer FIMT (Fiber In Metallic Tube)

*) calculated from the oxygen blowing condition

The temperature measured through top lance is rather scattered among the experiments. It may be due to dimension of experimental equipment or correction of the data. In the case of FIMT, the measured temperature is relatively higher

than those through top lance, at least, it is not underestimated compared to another method. As referred below, temperature measurement of molten metal by FIMT is done by immersing optical fiber directly into

metal, therefore turbulence from surroundings or emissivity of atmosphere is minimized compared to distant measurement. This would be beneficial for a direct measurement of hotspot.

Fig. 2 - Dependence of hotspot temperature on oxygen flow rate Temperature Measurement with FIMT A schematic illustration of experimental setup carried out by the present author is shown in Fig.3 (19). 10kg of carbonsaturated molten iron was prepared from pure iron in a graphite crucible with a high frequency induction furnace. The 60

initial composition of the molten iron was [Si]=0.25 mass%, [Mn]=0.20mass%, and [P]=0.110 mass%. The initial temperature of the molten iron was around 1603K ±20K. Oxygen gas and various powder auxiliary materials were blown from a top-blowing lance. The temperature at the area where the oxygen gas

was in contact with the molten iron was measured with an optical fiber thermometer through FIMT. FIMT was inserted into the hotspot with the aid of a steel guide. The flow rate of oxygen gas was 5-15L/min. The auxiliary materials used were quicklime, calcium carbonate and iron sand. La Metallurgia Italiana - n. 11/12 2018

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Fig. 3 - Schematic illustration of temperature measurement with FIMT The measured temperature at hotspot is shown in Fig.4. The temperature obtained with oxygen gas alone, without powder supply, increased with the

oxygen flow rate. The temperature at the hotspot measured with powder supply were lower than those without powder supply. With respect to the powder spe-

cies, the temperature was lower with supplying calcium carbonate than quicklime. The temperature also decreased when iron sand was supplied.

Fig. 4 - Hotspot temperature measured with or without powder supply Evaluation of Cooling Effect with Powder Supply Based on the experimental results, it is considered that the temperature at the hotspot is lowered when powder material is supplied to the hotspot. Since the powder materials have different cooling La Metallurgia Italiana - n. 11/12 2018

capacities, the cooling effect of powder materials should be evaluated more quantitatively. In order to estimate an accurate temperature of hotspot from a balance of heat input by oxidation of component in molten iron and heat output by supply

of powder materials, volume of hotspot should be defined. However, it is difficult to reasonably define its volume. Therefore, a relative comparison of the hotspot temperature was made for the different experimental conditions, assuming that the volume of the hotspot is constant 61

Experts’ corner under the same oxygen flow rate. Based on that assumption, the difference between the hotspot temperature with and without powder supply, ∆T, is considered to be proportional to the sum of the output heat, ∆Q, which comprises

the sensible heat required to elevate the powder temperature and the heat of decomposition of the powder compound. The concept of evaluation of the heat balance at the hotspot is shown schematically in Fig.5. The hotspot tempera-

ture without powder supply was given from regression of the experimental data shown in Fig.4 as a base for evaluating ∆T.

Fig. 5 - Concept of evaluating cooling effect of powder supply ∆T and ∆Q at the experiments are shown in Tab.2 (20). For example, in the experiment at the oxygen flow rate of 1.5Nm3/min/t (=15L/min) and the quicklime feeding rate of 2.0kg/min/t (=20g/min), ∆T was 290K. The sensible heat for elevating the temperature of the quicklime to the hotspot temperature was calculated as 3.57MJ/min/t.

∆Q for the other powder materials were also calculated tentatively at 100% decomposition rate (D.R.) of the powder compound. The ratios of ∆T to ∆Q thus obtained (=∆T/∆Q) are shown together in Tab. 2. ∆T/∆Q for supplying calcium carbonate at the 100% decomposition rate is close to that when supplying quicklime. This

agreement would possibly show that the present estimation is not invalid in terms of heat balance. By analogy with the above procedure, ∆T/∆Q =81.2 is adopted for the case of supplying iron sand. Under that estimation, the decomposition rate of iron sand is calculated as about 86%.

Tab. 2 - The condition and methodology for the measurements under top blowing oxygen

Oxygen flow rate (Nm3/min/t) Powder feed rate (kg/min/t) Temperature at hotspot (K) ΔT against no powder feed (K) ΔQ (MJ/min/t) ΔT/ΔQ

1.5

Quicklime 2.0

CaCO3 2.0

Iron sand 1.0

2483

2193

2063

2073

290

420

410

5.43 (D.R.100%) 78 (D.R.100%)

5.93 (D.R.100%) 69 (D.R.100%) 80 (D.R.86%)

3.57 81

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Scenari SLAG COMPOSITION NEAR HOTSPOT During the experiment by the present author referred above (19), sampling of the metal and slag were carried out at intervals of about five minutes. The slag and metal were sampled near the hotspot and near the crucible wall (he-

reinafter referred as bulk), as shown in the inserted illustration in Fig. 3. Fig. 6 shows the relationship between the basicity (=(mass%CaO)/ (mass%SiO2)) and mass%T.Fe of the slag. The data are chosen from the original data within ÂąĎ&#x192; (standard deviation) from the average of either basicity

or (mass%T.Fe). It is clear that the slag near the hotspot displays relatively high basicity and T.Fe, whereas the slag near the crucible (the bulk slag) is lower in both value. Although not shown in the figure, the phosphorus oxide content of the slag near the hotspot are also higher than those of the bulk slag.

Fig. 6 - Relationship between the basicity and (T.Fe) content of the slag DEPHOSHORIZATION BEHAVIOR NEAR HOTSPOT The aforementioned results are further analyzed in terms of the effect of powder

material on the dephosphorization behavior near the hotspot. The phosphorus distribution ratio between the slag and metal, Lp, as shown in the following Eq.

2, was used to evaluate the dephosphorization behavior.

(2)

where (mass%P) and [mass%P] are the phosphorus content in the slag and the molten iron respectively.

The effect of temperature and slag composition on Lp is further considered using the following empirical relation of

Lp reported by Healy (21) and Ogawa et al (22).

(3)

(4)

where (mass%T,Fe) is the total iron content in the slag and [mass%C] are La Metallurgia Italiana - n. 11/12 2018

the carbon content in the molten iron respectively. In Eq. 4 carbon content in

metal is taken into account.

Experts’ corner Hotspot is characterized by a sufficient supply of oxygen and a resulting high oxygen potential. However, in the experiment by the present author (19), the slag was sampled from the surrounding area of the hotspot due to the difficulty of obtaining a quantity of slag sufficient for chemical analysis from the center of the hotspot. Therefore, the following di-

scussion looks into the dephosphorization reaction “near” hotspot. In order to evaluate dephosphorization near hotspot, the temperature at the corresponding zone is required. However, it is difficult to define the temperature “near” hotspot accurately. Therefore, a rough estimation is given, here at this point, assuming that the temperature

“near” the hotspot is the average of the measured temperatures at the hotspot and in the bulk metal. Fig. 7 shows the relationship between the temperature at the hotspot and the assumed temperature near the hotspot thus calculated at a given temperature of bulk metal and a given oxygen flow rate.

Fig. 7 - Estimation of temperature in Reaction Zone

In the case shown in the figure, the hotspot temperature decreased from 2398K to 2130K when quicklime was fed at 10g/min. The temperature of bulk metal being 1603K, then the temperature “near” the hotspot is evaluated to be 1867K, about 263K lower than the hotspot temperature. Using the temperature “near” the hotspot thus obtained, the phosphorus distribution ratios are compared as shown in Figure 6. Lp can be obtained from equation [3] or [4] by using a given temperature and slag composition. [mass%P] in molten

iron equilibrated with slag having Lp thus obtained can be evaluated under a given (mass%P) in slag. Miyata et al discussed the dephosphorization behavior at hotspot in their experiment in 2 ton converter according to the calculated slag composition (10). The slag was assumed as FeO–CaO– P2O5 melt and its composition was calculated (not shown) based on CaO feeding rate, oxygen consumption for Fe oxidation and generation rate of P2O5 from change in [mass%P] at the experiment. Lp and [mass%P] at hot spot was estimated on a basis of Healy’s equation

[3] for temperature ranging from 2273 to 2573 K. The calculated phosphorus content is shown in logarithmic scale against temperature in Fig. 8. The value is sufficiently low compared to phosphorus level in industrial operation. They mentioned that high oxygen potential of the slag melt (more than 60mass%FeO in FeO-CaO binary) could compensate with negative effect of high temperature at hotspot

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Fig. 8 - Calculated [mass%P] against temperature In the experiment by the present author (19), the slag was directly taken inside the crucible, considered to be “near” hotspot. Using the obtained slag composition, Lp is calculated according to both Eq. 3 and 4 so that [mass%P] is evaluated and shown together in Fig. 8. In the calculation under Eq. 4, the carbon content in molten iron is given as 4.6 and 3.0 mass%. The former corresponds to carbon saturated condition in the experiment and the latter to typical carbon content at the final stage of industrial dephosphorization treatment. The calculated [mass%P] through Eq. 3 is strongly dependent on temperature compared to that through Eq. 4. If one apply Eq. 3 for temperature at hotspot as above 2200K, the calculated [mass%P] exceeds 0.1mass% and dephosphorization would not occur in this condition. In the case of Miyata’s calculation, the estimated (mass%T.Fe) of the FeO-CaOP2O5 melt is extremely high and leads to low equilibrium phosphorus content at hotspot temperature. On the other hand, the T.Fe content in the slag sampled “near” hotspot at the experiment by the present author is ranging from 15 La Metallurgia Italiana - n. 11/12 2018

to 26 mass%, which is not so high as Miyata’s estimation for hotspot. Since the slag was sampled “near” hotspot, it is appropriate to apply the temperature at corresponding area to the evaluation of Lp rather than the temperature at substantial hotspot. If the temperature of 1867K estimated in the previous section is applied for both Eq. 3 and 4, the calculated [mass%P] is around 0.001mass%, which would afford to support driving force for dephosphorization in conventional operational level. Therefore, slag formed at hotspot with CaO blasting would have originally high potential for phosphorus removal resulting from high FeO content, and would still hold dephosphorizing capacity after moving to and consuming FeO “near” hotspot. Based on the above discussion, the present author deems that, at least, substantial hotspot and “near” surrounding region would contribute to dephosphorization under CaO blasting condition though the temperature is higher than bulk metal.

oxygen refining of molten iron through understanding of the phenomena at hotspot, the data by the present author in a laboratory-scale experiment in which oxygen gas and several auxiliary materials were supplied to molten iron is re-evaluated. - The temperature measured with FIMT was relatively higher than those through top lance compared at the same oxygen flow rate, and is considered not to be underestimated compared to another method. - The measured temperature of the hotspot was lowered when powder material was supplied to the hotspot. - The slag sampled near the hotspot had higher basicity and T.Fe compared to the bulk slag . - Together with the analysis of the experimental data after Miyata et al, slag formed at hotspot with CaO blasting would have originally high potential for phosphorus removal resulting from high FeO content, and would still hold dephosphorizing capacity after moving to and consuming FeO “near” hotspot.

CONCLUSIONS For the purpose of better control of 65

Experts’ corner REFERENCES [1]

Dogan N, Brooks G A, Rhamdhani M A : ISIJ International, 51 (2011) 7, 1086.

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Dogan N, Brooks G A, Rhamdhani M A : ISIJ International, 51 (2011) 7, 1093.

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Dogan N, Brooks G A, Rhamdhani M A : ISIJ International, 51 (2011) 7, 1102.

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Millman M S, Overbosch A, Kapilashrami A, Malmberg D, Bramming M : Ironmaking Steelmaking, 38 (2011) 7, 499.

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Spooner S, Warnett J M, Bhagat R, Williams M A, Sridhar S : ISIJ International, 56 (2016) 12, 2171.

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Uchida, Y. Kishimoto, Y. Miki and M. Iwase: High Temp. Mater. Proc., 31(2012), 479.

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Murai T, Shimizu H, Matsuno H, Kikuchi Y : CAMP-ISIJ, 14(2001), 969.

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Tanigaki T, Kanbayashi T, Tajiri Y, Tani J, Miyata M, Tamura T : CAMP-ISIJ, 24(2011), 157.

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Miyata M, Tamura T, Higuchi Y : ISIJ International, 57 (2017) 10, 1751

[10] Miyata M, Tamura T, Higuchi Y : ISIJ International, 57 (2017) 10, 1756 [11] Koch K, Fix W, Valentin P : Arch. Eisenhüttenwes., 47 (1976), 659. [12] Koch K, Fix W, Valentin P : Arch. Eisenhüttenwes., 49 (1978), 163. [13] Nakamura M, Tate M : Tetsu-to-Hagané, 63 (1977), 236. [14] Denier G, Grosjean J C, Zanetta H : Ironmaking Steelmaking, 3 (1980), 123. [15] Harada N, Sakuraya T, Nakamura H, Fujii T, Habu Y : Tetsu-to-Hagané, 69 (1983), S1010. [16] Kawakami K : Tetsu-to-Hagané, 74(1988), 831. [17] Chiba K, Ono A, Saeki M, Yamauchi M, Kanamoto M, Ohno T : Ironmaking Steelmaking, 20 (1993), 215 [18] Matsui A, Nabeshima S, Matsuno H, Kishimoto Y : CAMP-ISIJ, 20 (2007) 136. [19] Uchida Y, Sasaki N, Miki Y : The 10th Korea-Japan Workshop on Science and Technology in Ironmaking and Steelmaking, Busan, Korea, 2014 [20] Kubashewski O, Alcock C B : MaterialsThermochemistry, 6thEdition, PergamonPress, New York, (1993) [21] G. W. Healy: J. Iron Steel Inst., 208 (1970), 664. [22] Ogawa Y, Yano M, Kitamura S, Hirata H : Tetsu-to-Hagané, 87(2001), 21.

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Scenari Development and application of high efficiency blowing technology of converters in Shougang edited by: P. Gao, X. Wang, H.Li, Y. Guo, Y. Zhang, X. Zh, H.Yu University of Science and Technology Beijing and Shougang Group Co., Ltd., China University of Science and Technology Beijing, China In order to solve the problem that the actual converter bottom blowing effect was difficult to guarantee due to low bottom blowing intensity and the slag splashing technology under long converter campaign life, high efficiency bottom blowing technology for converter steelmaking was developed in Shougang Co., Ltd. (SGQG). During the development process, a series of key technologies such as optimization of bottom blowing tuyeres arrangement and pool mixing flow field, exposure technology of bottom blowing tuyeres under the conditions of slag splashing, bottom blowing tuyeres long life technology, etc. The optimization of bottom blowing mode was carried out, and the stable control of converter bottom blowing effect was achieved, even at the late age of the campaign, the carbon oxygen product could still be less than 0.0020 under the converter campaign life of about 6500. By applying the above technologies, in the production of ULC and IF steel (converter end point carbon mass fraction was 0.025%~0.05%), the average reduction of converter end oxygen mass fraction was more than 0.025%, TFe mass fraction in end slag was more than 4%, and the erosion of refractory materials was significantly reduced, converter dephosphorization efficiency was significantly improved. KEYWORDS: CONVERTER – COMBINED BLOWING – BOTTOM BLOWING INTENSITY – CARBON OXYGEN PRODUCT – OXYGEN CONTENT OF MOLTEN STEEL – DEPHOSPHORIZATION

INTRODUCTION Top and bottom combined blowing for converter steelmaking was an important technology developed in the 1970s. At present, majority of the middle and large sized converters are using this technology. Compared with the top blowing converters widely used before, the stirring energy of the combinedblowing converter is much more, the reaction between the slag and steel is much closer to equilibrium, and the final molten steel [O] and slag FetO content are lower. Moreover, the higher the bottom-blowing intensity of the combined-blowing converter, the more pronounced of the above-mentioned metallurgical effect [1-2]. At present, the bottom blowing intensity of the combiLa Metallurgia Italiana - n. 11/12 2018

ned-blowing converter is 0.04~0.70 m3/ (t·min). In order to reduce the final molten steel [O] and slag FetO content for producing low-carbon and ultra-low carbon steel products, many high-level steel plants in the world take measures of using high bottom blowing intensity( 0.15 m3/(t·min) or more), but their converter campaign lives is mostly less than 5 000 heats, and many need to change the bottom of the furnace in the campaign life [3]. In order to solve the problem of the short life of the bottom blowing tuyeres and increase steel production after top and bottom combined blowing was introduced in China in 1980s, most steel plants had chosen low bottom blowing intensity (mostly 0.04 ~ 0.06 m3/(t·min)). In

addition, the number of bottom blowing tuyeres of converters in China is also significantly different from that of other countries. The common number is 8 ~ 12 for large and medium-sized converters in China (the number is even 16) compared with mostly four converter bottom blowing tuyeres in some high-level steel plants in the world [4]. Due to the low bottom blowing intensity of converters in China, combined with the large number of bottom blowing tuyeres, the single tuyere flow rate is significantly lower than that of high-level steel plants, and therefore it is easy to be blocked (especially in the later stage of the campaign) [5-6] . In addition, converter steel plants in China had generally adopted the slag 67

Experts’ corner splashing technology since the 1990s, and the converter campaign life has reached 5 500 to 15 000 heats (or even higher), which has played an important role in the rapid increase of the Chinese steel output. However, the “recurring” rise of the converter bottom thickness after the adoption of slag splashing process may easily cause the tuyeres to be partially or completely blocked, which will affect the bottom blowing effect seriously [5-6]. A lot of work was done and technologies such as bottom blowing tuyeres "reversal", "slag mushroom head" and others were developed to maintain the bottom blowing effect after slag splashing was introduced. After the adoption of these technologies, many steel plants have declared that their bottom blowing is "synchronous" with the long converter campaign life based on the gas flow rate displayed on the meter at the late`r age of the campaign, but the carbon oxygen product of molten steel at the blowing end is greatly increased from the previous period, indicating that the actual mixing effect has been greatly reduced [7]. In order to solve the problem that the actual bottom blowing effect was difficult to guarantee due to low bottom blowing intensity and the slag splashing technology under long converter campaign life, Stable Effective Bottom-

blowing Converter (SEBC) technology was developed in Shougang Co., Ltd. (SGQG). During the development process, a series of key technologies such as optimization of bottom blowing tuyeres arrangement and pool mixing flow field, exposure technology of bottom blowing tuyeres under the conditions of slag splashing, bottom blowing tuyeres long life technology, etc. The optimization of bottom blowing mode was carried out, and the stable control of converter bottom blowing effect was achieved, even at the late age of the campaign, the carbon oxygen product could still be less than 0.0020 under the converter campaign life of about 6500. By applying the above technologies, in the production of ULC and IF steel (converter end point carbon mass fraction was 0.025% ~ 0.05%), the average reduction of converter end oxygen mass fraction was more than 0.025%, TFe mass fraction in end slag was more than 4%, and the erosion of refractory materials was significantly reduced, converter dephosphorization efficiency was significantly improved. Steelmaking status of SGQG converters SGQG currently is equiped with five 210 ton converters, four RH refining furnaces, two LF refining furnaces, and four dual-flow slab casters. The annual ou-

tput of steel is about 7.8 million tons. The main products are automobile plate steel, silicon steel, household appliances use steel, high-strength steel, and pipeline steel. The output of cold-rolled steels, such as automobile plate steel, silicon steel, accounts for more than 70% of total output. These steel varieties require both low carbon and low oxygen content at the blowing end, which is a great challenges for converter steelmaking. SGQG 210t converter adopts 12 bottom blowing tuyeres. As shown in Fig.1, the double ring arrangement is used, and the actual bottom blowing intensity is about 0.05 m3/(t·min) while the designed intensity is 0.03 ~ 0.15 m3/ (t·min). In actual production, due to the low bottom blowing intensity and the adoption of slag splashing technology, the bottom tuyeres are easily be covered by the slag layer. Under the condition of campaign life of about 6 500, especially in the later age of the converter life, the bottom blowing tuyeres are difficult to be exposed and the stirring effect is difficult to guarantee. Therefore, the carbon oxygen product increases with the age of the furnace. In the early age of the campaign (furnace age ≤ 2 000 heats), the average carbon oxygen product is about 0.002 0, but in the later age of the campaign (furnace age ≥ 4 000 heats), the average carbon oxygen product exceeds 0.003 0.

Fig. 1 - Schematic diagram of the bottom blowing tuyeres arrangement of SGQG 210 t converter 68

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Scenari Therefore, SGQG started the SEBC program in 2014, and the goal is to achieve a stable control of the bottom blowing effect for the whole campaign life. DEVELOPMENT OF SEBC TECHNOLOGY Optimization of bottom blowing tuyeres arrangement and pool mixing flow field Compared with top blowing, bottom blowing has a much greater impact on the mixing effect of the molten pool[810].At the same time, both the arrangement and quantity of the bottom blowing tuyeres have significant influence on the mixing of the molten steel [11-12]. In order to calculate the influence of the number of the bottom tuyeres on the flow field of the molten steel, Fluent 16.1 was used to do the simulation work. In the numerical simulation cal-

culations, the chemical reactions in the steelmaking process was ignored, the argon bubble size was assumed to be uniform, the bottom blowing tuyeres were equivalent to a round hole with a certain cross section, the flow of molten steel in the converter is treated as an incompressible flow. Then Euler-Euler method was used to solve the continuity equation and momentum equation for argon gas and molten steel separately, and the standard k-ε turbulence model was used for calculation. The calculation results are shown in Fig.2 and Fig.3[13-14]. Fig.2 shows the turbulent kinetic energy (middle) and velocity vector (right) of the longitudinal section at the bold line shown in the left with 12 and 4 bottom blowing tuyeres respectively. As with the same bottom blowing intensity (0.05m3/(t·min)), the bottom blowing flow rate of the single

tuyere when 4 tuyeres arrangement is used is much higher than that of 12 tuyeres. It can be seen from the middle of the figure, the turbulent kinetic energy in the molten pool when 4 tuyeres arrangement is used is significantly greater than that of 12 tuyeres. As shown in the right of the figure, multiple circulation areas is formed in the molten pool when double-circle arrangement of 12 tuyeres is used. The mutual impact between upper and lower circulation of the molten steel leads to dissipation of stirring energy, which weakens the actual mixing effect, causing difficulty to guarantee the mixing effect of internal components and temperature of molten steel. When four tuyeres arrangement is used, a large circulating flow is formed inside the molten pool, especially in the area between the tuyeres and converter body lining.

Fig. 2 - Turbulent kinetic energy and velocity vector of the longitudinal section of bottom blowing Fig.3 shows the speed cloud diagram (left) and vector diagram (right) of cross section 50mm under the surface of steel level. It can be intuitively known from the figure that the velocity of the molten steel with 4 bottom blowing tuyeres at the cross section is significantly larger than that of 12 tuyeres, that’s also because the bottom blowing flow rate La Metallurgia Italiana - n. 11/12 2018

of the single tuyere when 4 tuyeres arrangement is used is much higher than that of 12 tuyeres with the same bottom blowing intensity (0.05m3/(t·min)). Because the confluence of adjacent streams causes energy loss, as shown in Fig.3, a low-velocity zone will be formed between the adjacent streams. At the same time, as the bottom blowing

flow rate of the single tuyere is small when 12 tuyeres arrangement is used, that means a weak stirring ability to the molten steel especially near the liquid surface. Therefore, compared with the four bottom blowing tuyeres, 12 bottom blowing tyueres shows weaker stirring ability to the entire molten pool.

Expertsâ&#x20AC;&#x2122; corner

Fig. 3 - Speed cloud diagram and vector diagram of cross section 50mm under the surface of steel level Based on the above analysis, the original 12 bottom tuyeres of the 210t converter were optimized to 4, which increased the bottom blowing flow rate of the single tuyeres and the stirring energy dissipation caused by the multi-bottomed blower tuyeres was avoided. At the same time, the increase in the flow rate of the single tuyere increases the kinetic energy at the outlet of the tuyere and improves the mixing effect. Exposure control of tuyeres under slag splashing technology Bottom tuyeres exposure ensures the stability and effectiveness of the bottom blowing effect. However, under the con-

a) Slag layer at the bottom of the furnace

dition of slag splashing and low bottom blowing intensity, it is difficult to ensure the exposure of the bottom tuyeres. A non-vacuum induction furnace was used to simulate the effect of the slag splashing process on the exposure of the bottom tuyeres in the laboratory. Fig.4 is SEM photographs of the slag layer formed on the bottom refractory lining and around the bottom blowing tuyere at the end of the experiment under a certain experiment parameter. It can be seen from the figure that the main components of the slag layer on the bottom refractory lining are high melting point materials such as 2CaOÂˇSiO2, , thus having good resistance to erosion and dissolution. At

the same time, the compositions of the slag layer around the tuyere are mainly FeO with relatively low content of CaO, SiO2, which means the melting point of the slag is generally low. Therefore, the separate melting characteristics of the slag and the good liquidity of the lowmelting-point slag can be utilized to control the slag splashing layer on the bottom refractory lining to be a highmelting point material, and the slag layer around the bottom blowing tuyere to be a low-melting point material, which can achieve results of both protecting of the refractory lining of the bottom and ensuring a good bottom blowing effect.

b) Slag layer above the bottom blowing tuyere

Fig. 4 - SEM photograph of the phase of the different components formed after slag splashing process

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Scenari Dynamic maintenance technology of the bottom refractory lining and tuyeres was developed using the results obtained by the laboratory. The combination of slag splashing and furnace gunning are used to ensure that the thickness of the bottom lining and the length of the bottom tuyeres decreases steadily. The main contents include the combination use of light slag splashing, strong slag splashing, slag hanging technology, key maintenance of bottom tuyeres with large particles of mending material. The separate melting characteristics of the slag during the slag splashing process is used, and then by controlling the blowing intensity of bottom blowing

tuyeres during the slag splashing process, the sticky slag is prevented form covering the bottom blowing tuyeres. When the converter is standing still after the slag splashing, the bottom blowing flow rate can be adjusted to a certain content instead of zero. Using the good liquidity of the low-melting-point slag, and controlling the slag with high FeO content and a low melting point to covere the bottom blowing tuyeres, during the blowing process, since the melting point of the slag is low, it can be quickly melted to ensure the bottom blowing effect. Since the melting point of the slag covered on the bottom lining is relatively high, which will achieve the effect

of protecting the lining, and blockage of the bottom tuyeres caused by the converter gunning, especially the area around the bottom tuyeres is avoided. After adopting of the dynamic maintenance technology, the control result is shown in Fig.5. after slag splashing, the bottom lining is covered with thick slag layer with high-melting-point, which will achieve the effect of bottom refractory lining protection. In addition, the slag covered above the tuyere is relatively thin and has a low melting point, since the bottom blowing flow rate of the single tuyere is large, the bottom blowing tuyeres exposure can be guaranteed during the blowing process.

Fig. 5 - Control result of the dynamic maintenance technology Long-life technology of the bottom blowing tuyere Due to the flow rate of the single bottom tuyere increases, the repulsive forces caused by the expansion of the bubbles increase [15-16], which intensifies the erosion of the bottom lining. In the actual production process, once the bottom lining or the bottom tuyeres are seriously eroded, the bottom lining and bottom tuyeres will be covered with the thick slag after slag splashing or gunning, causing the bottom blowing La Metallurgia Italiana - n. 11/12 2018

tuyeres to be blocked, and the bottom blowing effect is even more difficult to guarantee. Therefore, the bottom tuyeres can only be of good use when the thickness of the bottom lining and bottom blowing tuyeres are steadily decreaseing. To this end, long-life technology of the bottom blowing tuyere was developped. The main contents include: 1) Develop a detailed control plan for the thickness of the bottom lining as a function of the age of the converter, and then control

the thickness of the bottom lining in stages accurately, and control the thickness of the bottom lining to continue to decrease steadily and slowly. 2) Establish a reference standard for the determination of the exposure of the tuyeres, and determine the exposure state of the tuyeres according to the reference standard. 3) Combining the dynamic maintenance technology of the bottom tuyeres and furnace lining, adjust the parameters of the slag splashing and gunning dynamically, including slag splashing frequency 71

Expertsâ&#x20AC;&#x2122; corner and strength, MgO content control, gunning parameters, bottom blowing flow rate, and branch pipe distribution ratio adjustment, etc. according to the carbon oxygen product at the end of the blowing of 5 to 10 heats before current heat, the thickness of the bottom lining and the thickness of the bottom tuyeres, and the exposure status of the bottom tuyeres. The ultimate goal is to achieve that the lining life is not reduced due to

the exposure of bottom tuyeres for the full campaign life. The technical application results are shown in Fig.6 and Fig.7. Fig.6 is photos of the bottom tuyeres during a campaign with life of about 6 500, in the whole life, the bottom tuyeres are clearly visible. The change trend of lining thickness with furnace age is shown in Fig.7. It can be seen from the figure that the thicknesses of the bottom lining

and the bottom tuyeres show a uniform and stable decline with the age of the furnace, which guarantees the long-life of the lining under the condition of low blowing intensity and the slag splashing technology. Finally, the goal of the bottom blowing effect be stable in the whole campaign life is realized with the campaign life of about 6 500 heats.

Fig. 6 - Photos of bottom blowing tuyeres at converter campaign ages

Fig. 7 - Thickness change trend of converter bottom lining and bottom blowing tuyere

La Metallurgia Italiana - n. 11/12 2018

Scenari Bottom blowing mode optimization In order to achieve the goal of lower carbon oxygen product at the end of the converter, in the process of converter blowing, a bottom blowing mode with â&#x20AC;&#x153;high-low-highâ&#x20AC;? flow rates is used. The overall control concept is:1)In the early stage of converter blowing, in order to speed up slagging, a larger bottom blowing flow rateshould be used to promote good pre-slag and better dephosphorization. 2)In the middle stage of the blowing, the flow rate of the bottom blowing should be reduced to maintain the lining of the molten pool and the bottom of the furnace. 3) In the later stage of the blowing, improving the bottom blowing flowing rate is effective to improve the dephosphorization effect and reduce the oxygen content of the converter, but the lining erosion will be serious, so the strong bottom blowing time need to be controlled reasonable, and the cooperation of bottom blowing and oxygen lance should be optimized to reduce the molten steel oxygen content and the final slag T.Fe content. In order to ensure the stability of the bottom blowing effect during the entire campaign life, different blowing modes are used during the early, middle and later ages of the converter campaign. A lower bottom blowing flow rate can meet the lower carbon oxygen product in the early age of the campaign; In the middle age of the campaign, more attention should be paid to the maintenance of the bottom tuyeres; In the later age of the campaign, a combination of higher bottom-blowing flow rate and a complementary furnace gunning can ensure the

La Metallurgia Italiana - n. 11/12 2018

stability of the bottom blowing effect. After the optimization of the bottom blowing mode, the long life of the lining of the converter under the condition of slag splashing and low blowing intensity is ensured, and the stability of the bottom blowing effect is also ensured, and the dephosphorization efficiency is improved. APPLICATION OF SEBC TECHNOLOGY The development of SEBC technology in Shougang, high efficiency bottom blowing technology for converter steelmaking is guaranteed under low bottom blowing intensity, the slag splashing technology, and long converter campaign life. The comparison of the converter end-point control before and after the process application is shown in Fig.8. The moving average analysis method in the Microsoft EXCEL software was used in the mapping, that is, according to the order of the abscissa, the values of each 5 corresponding heats on the ordinate are averaged, and then all the average values are connected into a curve. Fig. 11a) is the comparison of the carbon oxygen product at the end of the blowing with the age of the furnace. As can be seen in the figure, before the process was applied, after the converter age of 2,000 heats, the carbon oxygen product at the end of the blowing rose with the converter age. The average carbon oxygen product after 4 000 furnaces was even above 0.0030. After the application of the technologies, the carbon oxygen product at the end of the blowing does not increase with the age of the converter, with a converter campaign life

of more than 6 500 heats, the carbon oxygen product can be stably controlled below 0.002 0 even at the late age of the campaign, achieving the result of stable and effective bottom blowing effect. Fig11b) shows the oxygen content changes with the carbon content at the end of the blowing before and after the technologies application. As can be seen in the figure, the oxygen content at the end of the converter is significantly reduced under the same carbon content at the end of the blowing, especially when low-carbon steel or ultra-low carbon steel is produced (control aim of the carbon content of the converter is 0.025 to 0.05%). The final oxygen content was reduced from 600-850 ppm to 400-550 ppm with an average reduction of over 250 ppm.

Expertsâ&#x20AC;&#x2122; corner

a) Change of carbon oxygen product with furnace age

b) Variation of oxygen content with carbon content

Fig. 8 - Comparison of converter end control before and after the application of technology

Bottom blowing mode optimization In addition, due to the improvement of the bottom blowing effect, T.Fe mass fraction of the final slag of the converter blowing is reduced by an average of 4% or more after the application of the technologies than before. At the same time, the consumption of gunning material was reduced, with an average decrease of 0.57kg/t, a reduction rate of 37.2%; the slag splashing materials including dolomite, light-burned dolomite and magnesium-carbon spheres decreased by an average of 2.9kg/t, with a decrease rate of 65.9%. Due to the stability of the actual bottom blowing effect and the optimization of the bottom blowing mode, the dephosphorization efficiency of the converter is significantly higher than before the application. CONCLUSION The problem that the actual converter bottom blowing effect was difficult to

guarantee due to low bottom blowing intensity and the slag splashing technology under high converter campaign life was solved due to the development and application of high efficiency blowing technology of converters in Shougang. The goal of stable control of the bottom blowing effect of the whole converter campaign life is achieved. 1) Numerical simulation results show that the optimization of the 12 bottom blowing tuyeres into 4 tuyeres can form a large circulation in the molten pool, avoiding the stirring energy dissipation caused by the mutual impact between the upper and lower circulations of the molten steel, which effectively improves the bottom blowing effect; 2) Taking advantage of the separate melting characteristics of the slag and the fluidity of the low-melting point slag, it is possible to control the slag above the bottom lining as high melting point slag to protect the lining, and slag above the bottom tuyeres as low-melting point slag, which can ensure the blowing ef-

fect during the whole campaign life; 3) In order to solve the problem of the increased refractory erosion caused by the increased flow rate of the single tuyere, a long-life technology of tuyeres has been developed, and the thickness of the bottom lining has been controlled steadily decreasing at a slow speed, the parameters of the slag splashing and gunning is dynamically adjusted, which achieves the effect of not only extending the service life of the lining, but also ensuring the bottom blowing effect; 4) After the application of high efficiency blowing technology of converters, for production of low-carbon, ultra-low carbon steel (end point carbon mass fraction of 0.025% ~ 0.050%), the end molten steel oxygen content gets an average reduction of 0.025% or more than before, the final slag TFe mass fraction is reduced by more than 4%, the erosion of the refractory material is significantly reduced, and the dephosphorization efficiency of the converter is significantly improved.

La Metallurgia Italiana - n. 11/12 2018

Scenari REFERENCES [1]

YEQI NU. Bottom Blown Converter Introduction, Elucidation of Mixing Phenomena and Extensive Development, Beijing: Metallurgical Industry Press, 2008:24-25

[2]

EMI T, Whether BOF steelmaking? - A Japanese perspective. Proceedings of the 3rd International Congress on the Science and Technology of Steelmaking, ICS2005, Charlotte, 2005: 269-275

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KISHIMOTO Y, SAITO N. Development and prospect of combined blowing converter in Japan. Tetsu-to-Hagane, 2014, 100 (4): 445-455.

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Wang Xinhua, LI Jinzhu, Liu Fenggang. Technological progress of BOF steelmaking in period of development mode transition. Steelmaking, 33(1), 2017.2:1-12

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Liu L. Technological Progress in Converter Steelmaking. Steelmaking, 2000.

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Yang W Y, Jiang X F, Wang M L, et al. Technical parameters optimization of steelmaking in large vessel. Iron & Steel, 2010, 45(10):27-32.

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Jiang X F. Present Situation and Development of 300t Combined Blowing Converter at Baosteel. Bao Steel Technology, 2002.

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Nakanishi K, Kato Y, Nozaki T, et al. Cold Model Study on the Mixing Rates of Slag and Metal Bath in Q-BOP[J]. Tetsu- to- Hagane, 2009, 66(9):1307-1316.

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YANG Li-bin, ZHENG, et al. Cold Simulation of Oxygen Transfer Rate in BOF. Journal of Iron and Steel Research, International, 2010, 17(9):7-13.

[10] PAUL S, GHOSH D N. Model study of mixing and mass transfer rates of slag-metal in top and bottom blown converters. Metallurgical Transactions B, 1986, 17(3): 461-469. [11] GHOSH D N, SINGH R P. Cold model study of mixing time in hybrid process. Transactions of the Iron and Steel Institute of Japan, 1988, 28 (8): 659-662. [12] SINGH R P, GHOSH D N. Cold model study of mixing and mass transfer in LBE process of steelmaking. ISIJ International, 1990, 30 (11): 955-960. [13] Jiao LI, Hui-Xiang YU, Wang X H, et al. Optimization of arrangement of bottom tuyeres through numerical simulation. Iron & Steel, 2018. [14] Gao P, Haibo LI, Guo Y, et al. Development and application of high efficiency blowing technology of converters in Shougang. Steelmaking, 2018. [15] WEI J H, MA J C, FAN Y Y, et al. Back-attack phenomena of gas jets with submerged horizontally blowing and effects on erosion and wear of refractory lining. ISIJ International, 1999, 39(8): 779-786. [16] YANG Q X, GUSTAVSSON H. Effects of gas jet instability on refractory wear a study by high-speed photography. Scandinavian Journal of Metallurgy, 1992, 21(1):15-26.

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Le Rubriche - Centri di studio Attività dei Comitati Tecnici CT METALLURGIA FISICA E SCIENZA DEI MATERIALI (MFM) (riunione del 12 settembre 2018) Consuntivo di attività svolte - Il convegno “Materiali metallici e processi produttivi innovativi per l’aerospazio”, organizzato in collaborazione con altri Centri di Studio AIM, si è svolto a Napoli nei giorni 19 e 20 luglio. Angella riferisce della buona riuscita della manifestazione, con partecipanti provenienti dal mondo accademico e industriale. Iniziative future - Nel mese di luglio 2019 si ripeterà il corso “Microscopia elettronica in scansione SEM per metallurgisti”. I coordinatori Angella e Bassani valuteranno se utilizzare le stesse sedi del 2017 (Milano e Lecco) o organizzare un corso itinerante coinvolgendo altri enti e Università. Federica Bassani fa presente che il CT CCP (Controllo e Caratterizzazione Prodotti) è interessato a collaborare e chiede ai coordinatori di contattare il presidente di quel centro. - Si discute circa la possibilità di ripetere entro il 2020 il corso “Additive Manufacturing”, dopo il successo di quello tenuto nel 2017. Si prospetta la possibilità di coinvolgere il CT TTM (Trattamenti Termici e Metallografia), qualche membro del quale sta lavorando su questi argomenti, senza però dimenticare i CT Metallurgia delle Polveri e Additive Manufacturing e Metalli Leggeri, che hanno già collaborato nella precedente edizione. - I coordinatori Kaciulis e Mezzi stanno lavorando per la programmazione della GdS “Tecniche di indagini delle superfici”, da organizzare entro fine 2019 in sede da definire. Mezzi pensa di coinvolgere anche altri comitati, come per esempio Controllo e Caratterizzazione Prodotti, per ampliare il bacino dei potenziali partecipanti.

Stato dell’arte e notizie - Il Comitato ha votato per il rinnovo delle cariche eleggendo come Presidente la prof.ssa Emanuela Cerri, come vice presidente l’ing. Paola Bassani e come segretario il dott. Riccardo Donnini. Il presidente uscente Angella si congeda dal suo incarico ringraziando tutti per la fattiva collaborazione avuta e rimanendo a disposizione come membro del CT. La neo-presidente Cerri ha ringraziato per la nomina e si è resa da subito disponibile per esaminare ogni tipo di proposta per nuove manifestazioni; inoltre ha invitato i membri a favorire il coinvolgimento di personale nuovo e giovane.

CENTRO AMBIENTE E SICUREZZA (AS) (riunione del 12 settembre 2018) Consuntivo di attività svolte - Durante la sessione plenaria di apertura del 37° Convegno Nazionale AIM a Bologna si è svolta la cerimonia di assegnazione dei premi “Sicurezza”, e nello stesso pomeriggio sono stati presentati i migliori progetti in una sessione dedicata.

La Metallurgia Italiana - n. 9 2018

Stato dell’arte e notizie Iniziative future - Per il 2019 viene riproposto il corso itinerante sulla sicurezza, della durata di 3 giorni, una GdS sulla gestione delle comunicazioni con RLS/RSU e una GdS sul tema dei rischi derivanti da attività di logistica interna (viabilità e movimentazione di carichi). Stato dell’arte e notizie - Un nuovo membro del CT, proveniente dall’Università di Bologna, viene accettato e parteciperà alle future riunioni. CT METALLURGIA DELLE POLVERI E TECNOLOGIE ADDITIVE (M) (riunione del 5 ottobre 2018) Consuntivo di attività svolte - Due sessioni del 37° Convegno Nazionale AIM di Bologna erano dedicate all’Additive Manufacturing. I presidenti di quelle sessioni, Casati e Biffi, riferiscono di un ottimo riscontro in termini di partecipazione alle sessioni, oltre ad una qualità medio-alta dei lavori presentati. Si discute della necessità che i relatori delle presentazioni universitarie debbano dare anche informazioni circa la possibile applicazione in ambito industriale, soprattutto per evitare di demotivare i possibili utenti provenienti dall’industria. Manifestazioni in corso di organizzazione - Nel marzo del 2019 si svolgerà la GdS “Trattamento termico di materiali e componenti prodotti per manifattura additiva”, organizzata congiuntamente con il CT Trattamenti Termici e Metallografia” e coordinata dal prof. Pellizzari. Si distribuisce l’elenco dei possibili relatori e argomenti. La sede della manifestazione è fissata presso la ditta BeamIT a Fornovo di Taro. Iniziative future

- La GdS “Press & Sinter: evoluzione di prodotti, materiali e tecnologie” sarà organizzata per il giugno 2019 – coordinatore Federici. Il tema sarà volto a delineare l’evoluzione dei prodotti e le conseguenti implicazioni per materiali e tecnologie utilizzate. - Molinari coordinerà gli sforzi per organizzare la GdS “Analisi critica dell’Additive Manufacturing” per ottobre 2019. Lo scopo è quello di fare il punto sull’Additive Manufacturing, da considerare ormai non più come un fenomeno nuovo. - Il Corso “Additive Metallurgy” sarà co-organizzato con i CT ML e MFM. I coordinatori Casati e Biffi devono verificare se si riuscirà a fare attorno alla metà di luglio 2019 o sarà opportuno rimandarlo al 2020 per evitare sovrapposizioni con altri eventi.

- Il CT accetta due nuovi membri, uno già presente come ospite e l’altro che parteciperà a partire dalla prossima riunione. - Il prof. Bocchini, per molti anni anima del CT e più volte Presidente, ha comunicato le sue dimissioni nell’ambito di una progressiva riduzione dei propri impegni professionali. Il presidente Molinari e tutto il CT esprimono gratitudine e riconoscenza per il contributo fondamentale da lui sempre profuso. CT ACCIAIERIA (A) CT FORGIATURA (F) (riunione congiunta del 19 ottobre 2018)

Consuntivo di attività svolte - La Conferenza EOSC tenutasi a Taranto dal 10 al 12 ottobre ha visto la partecipazione di 160 persone provenienti da 23 paesi, con una buona partecipazione di coreani e giapponesi a fronte di un solo partecipante dalla Cina. La visita agli impianti non è stata effettuata e rinviata sperabilmente al prossimo Convegno Europeo sulla Colata Continua (ECCC 2020), che si terrà a Bari fra due anni Iniziative future - La GdS/Workshop “Difettosità in Colata Continua e Lingotti” è prevista per marzo 2019. Il presidente Mapelli presenta un elenco di possibili argomenti da trattare: se la risposta sarà particolarmente positiva, la manifestazione potrebbe essere estesa su due giornate. - Si discute di un possibile evento sull’applicazione dei prodotti di acciaio piani e lunghi nel settore automotive, con la possibile partecipazione di ANFIA e Assofond. L’argomento di fondo sarà la scelta futura tra auto ibrida, elettrica o fuel cell, certamente con massa dei veicoli ridotta rispetto ai valori odierni. - Nel 2019 si terrà il corso itinerante di metallurgia fuori forno, per il quale si devono verificare le disponibilità delle acciaierie per ospitare le visite. C’è interesse per il mondo degli acciai inossidabili.   - Viene riproposta una possibile GdS sulle materie prime, durante la quale trattare la pulizia del rottame e analizzare anche il possibile contributo di carica di acciaio in altoforno per incrementare la produttività. Durante la giornata si potranno valutare i diversi elementi in lega, la loro importanza e le loro criticità. Si analizzeranno le problematiche legate a silicio, fosforo, zinco; lo sviluppo di nuove leghe come gli acciai ad alto Mn (15%) oppure ad alto Al (10%) che oggi sono applicati principalmente in ambito militare.

Stato dell’arte e notizie - Nell’ambito del CT Acciaieria, la carica di vice presidente è stata assegnata a Giampaolo Foglio. Il comitato ringrazia Francesco Magni, vice presidente uscente e past-president. Sono confermate le cariche di Carlo Mapelli come presidente e di Silvia Barella alla segreteria. - Vengono presentati un nuovo membro del CT

Le Rubriche - Centri di studio Acciaieria e il sostituto di un membro che si ritira dall’attività finora svolta. CT TRATTAMENTI TERMICI E METALLOGRAFIA (TTM) (riunione del 29 novembre 2018) Consuntivo di attività svolte

-La GdS “Trattamenti termici degli acciai per stampi a caldo e a freddo per il settore automotive” si è svolta ad Ivrea l’11 ottobre 2018 con 70 partecipanti. Il presidente Petta conferma il buon successo anche come approfondimento del corso già tenuto nella stessa sede. Il coordinatore Rivolta segnala che i questionari compilati hanno mostrato alti valori di soddisfazione, con qualche richiesta di approfondimento per le future GdS. -La GdS “Ottimizzazione dei trattamenti termochimici e dei processi meccanici nell’industria meccanica” (8 novembre 2018 a Provaglio d’Iseo) si è svolta presso la Gefran con circa 45 partecipanti: Petta ringrazia l’azienda per l’ospitalità e l’organizzazione di ottimo livello, oltre che per l’interessante e apprezzata visita in azienda. Morgano, coordinatore della giornata, traccia un bilancio della giornata, giudicata dai partecipanti tra il buono e l’ottimo nei questionari compilati. La giornata è stata dedicata a Giuseppe Rosso della Silco e a sua moglie Miriam, recentemente scomparsi in un tragico incidente.

Manifestazioni in corso di organizzazione

-Il presidente Petta parla del convegno ECHT 2019 “Heat Treatment & Surface Engineering for automotive”, che si svolgerà a Bardolino del Garda (VR) dal 5 al 7 giugno 2019 con chairman l’ing. Morgano. Data l’importanza strategica in ambito internazionale della buona riuscita di questo convegno, Petta raccomanda a tutti i membri del CT di attivarsi per diffondere la notizia dell’iniziativa e raccogliere memorie e partecipazioni. -La GdS “Trattamento termico di materiali e componenti prodotti per manifattura additiva” è coorganizzata con il CT “Metallurgia delle Polveri e Tecnologie Additive”. In assenza del coordinatore Pellizzari, l’ing. Valentina Vicario, coordinatrice per il CT MP, in collegamento telefonico ragguaglia sullo stato di avanzamento. La manifestazione si terrà presso la Beamit di Fornovo Val di Taro nel marzo 2019: c’è un programma preliminare e sono già stati ipotizzati i relatori. Iniziative future

-Il coordinatore De Sario, oggi assente, ha confermato al presidente Petta la disponibilità di Vimifasteners di Reggiolo (RE) a organizzare la GdS “Bulloneria e trattamenti termici” nella prima metà di maggio 2019. I relatori sono già stati individuati; tra questi, la stessa azienda ospitante presenterà una o due memorie. E’ prevista una visita al sito produttivo. -A ottobre 2019 si terrà una GdS su “Utilizzo

acciai nei sistemi frenanti” con coordinatori Massa e Morgano, che fa il punto e conferma che la sede sarà presso la Brembo al KilometroRosso di Dalmine. -Convegno Nazionale Trattamenti Termici 2020: a seguito di quanto discusso durante la precedente riunione, e degli approfondimenti organizzativi del presidente Petta, il Convegno del 2020 si terrà a Genova. Petta illustra i motivi della scelta e si ripromette di coinvolgere attivamente tutte le aziende operanti nel settore dei trattamenti termici, in modo da rafforzare una associazione di categoria così importante; quindi non solo presentazione di memorie ma anche spazi di rappresentanza. A tale scopo verrà contattato il CT Sviluppo Trattamenti Termici al fine di coordinare e valutare l’organizzazione per le adesioni preliminari al progetto. Stato dell’arte e notizie

-Nel corso della riunione è stata consegnata una targa al prof. Ramous per la lunga militanza nel CT TTM e per le attività formative offerte tramite AIM. -Cinque nuovi candidati hanno manifestato interesse a partecipare alle attività del CT: quattro provengono dal mondo industriale, uno dal settore accademico. Il presidente Petta li presenta e loro illustrano le aziende di appartenenza e le attività da loro svolte. Tutti vengono accettati e dalla prossima riunione parteciperanno come membri effettivi.

La siderurgia italiana piange lo storico presidente di Acciaierie Venete Gianfranco Banzato Si è spento a 82 anni Gianfranco Banzato, padre dell’attuale presidente di Acciaierie Venete e Federacciai Alessandro Banzato. Gianfranco Banzato è stato presidente del Cda di Acciaierie Venete fino al 2013. Era figlio di Marcello Banzato che nel 1957, dopo sette anni di esperienza nella fusione della ghisa, fondò le Acciaierie Fonderie Venete, ribattezzata Acciaierie Venete, nel 1974. A Gianfranco Banzato va il merito di aver intuito la necessità di elevare la qualità della produzione abbandonando le commodities per dedicarsi alle specialties, un processo avviato sin dalla fine degli anni Novanta e che oggi ha fatto dell’azienda padovana una delle principali realtà nella produzione di acciai speciali in Italia. Alla moglie Nadia ai figli Alessandro e Antonella e a tutti i famigliari le nostre condoglianze.

La Metallurgia Italiana - n. 11/12 2018

Atti e notizie CleanTech4 L’Associazione Italiana di Metallurgia ha organizzato la quarta edizione del Convegno europeo sulle tecnologie pulite nell’industria dell’acciaio CLEAN TECH 4. L’evento si è tenuto a Bergamo il 28 e 29 novembre 2018, presso il Centro Congressi Giovanni XXIII, ed è stata un’importante occasione di incontro tra ricercatori e produttori di acciaio, e ha trattato le seguenti tematiche: Tecnologie pulite: riduzione dell’impatto ambientale e delle emissioni per soddisfare e superare gli attuali obiettivi e soddisfare i requisiti delle normative future Efficienza energetica e riduzione delle emissioni di CO2: in che modo l'industria siderurgica può contribuire a un'economia a basse emissioni di carbonio, compresi sviluppi quali ULCOS, AISI CO2, CORSO 50, utilizzo del carbonio ed efficienza energetica Efficienza dei materiali: ottimizzazione dell'uso delle materie prime, implementazione di una gestione integrata dei rifiuti Cicli di vita del prodotto: dimostrazione del ruolo chiave dei prodotti siderurgici in una società sostenibile attraverso l'implementazione dei principi del ciclo di vita nella fase iniziale della progettazione del prodotto. Questo incontro ha avuto un ottimo riscontro ed è stata grande opportunità di scambio di competenze tecniche a livello internazionale tra i numerosi esperti coinvolti nella fabbricazione dell'acciaio e di aggiornamento sulle ultime realizzazioni tecniche in tutte le fasi di produzione e sui benefici ambientali dell'utilizzo dell'acciaio. Chairmen dell’evento sono stati l’ing. Roberto Pancaldi, CEO di Tenova Metals e l’ing. Maurizio Fusato, Direttore di stabilimento di Feralpi Siderurgica di Lonato Brescia.

La Metallurgia Italiana - n. 11/12 2018

Aim news

La Metallurgia Italiana - n. 11/12 2018

Atti e notizie Giornata di Studio “Ottimizzazione dei processi termochimici e meccanici nell’industria meccanica” Provaglio d'Iseo 08/11/2018

Lo scorso 8 novembre si è svolta la Giornata di Studio sull’Ottimizzazione dei processi termochimici nell’industria meccanica, organizzata da Gefran nella bellissima cornice del lago di Iseo. La manifestazione ha visto la partecipazione di numerosi rappresentanti della filiera del manufatto metallico, con memorie che hanno riguardato tematiche dalla materia prima (acciaieria) alla pallinatura finale (controllata). La giornata è stata dedicata alla memoria di Giuseppe Rosso dell’azienda Silco (TO), morto tragicamente in un incidente d’auto insieme alla moglie Miriam, la settimana prima. Giuseppe, oltre ad essere stato uno degli ideatori della manifestazione, ha sempre avuto la proattività di ricercare e proporre momenti di aggregazione tecnica. Lo voglio ricordare con affetto e per la sua disponibilità nella presentazione di memorie recenti, quali i convegni di Mestre e Bologna da poco conclusi e nell’organizzazione di workshop tecnici, sempre accompagnato dalla moglie Miriam. Enrico Morgano CRF - an FCA Company

La Metallurgia Italiana - n. 11/12 2018

Aim news Indice analitico ACCIAIERIA

ACCIAIO INOX FERRITICO

Latest results in EAF optimization of scrap-based melting process: Q-MELT installation in Kroman Celik pag.7/8-41 M. Ansoldi, D. Patrizio, M. Piazza, O. Kuran

Role of the stabilizing elements in the ferritic stainless steel on the microstructural, corrosion properties and durability of the stainless steel weld joints pag.5-11 D. Dyja, R. Marek, S. Zbigniew

Water leak detection in EAf based on Tenova’s offgas technology: recent developments and results in lucchini RS, Lovere, Italy pag.7/8-58 M. Luccini, V. Scipolo, D. Zuliani, L. Poli, D. Masoero

Development of new ferritic stainless steel sheets with excellent heat resistance and formability for automotive exhaust manifolds pag.5-33 J-i. Hamada, N. Kanno, A. Hayashi, N. Hiraide, M. Abe, K. Nishimura, C. Takushima, A. Yakawa, F. Fudanoki

ACCIAIO Mechanical properties evolution on heat treated severe cold rolled UNS S32760 Super Duplex Stainless Steel pag.1-20 C.M. Tromellini, A.F. Ciuffini, A. Gruttadauria, S. Barella, C. Di Cecca, C. Mapelli Development of new tests to assess sulfide stress corrosion cracking of line pipes pag. 6-11 M. Cabrini, S. Lorenzi, T. Pastore, D. Pesenti Bucella, G. Pellegrini, A. Paggi, E. Paravicini Bagliani, P. Darcis Hot-rolled steels for special purposes: high carbon wire-rod for cold-drawing pag.6-38 N. Bolognani, A. Parimbelli Hot ductility behavior of a high alloy steel pag.7/8-15 J.-H. Min, S.-H. Kwon, S.-D. Lee, S.-W. Moon, D-.K. Kim, J.-S. Lee, Y.-U. Heo, C.-H. Yim Caratterizzazione della resistenza al rinvenimento di un acciaio per lavorazioni a caldo AISI H13 modificato pag.9-23 F. S. Gobber, D. Ugues, P. Martelli, M. Rosso

ACCIAIO MICROLEGATO Surpassing steel performance by creating a very fine grained microstructure pag.6-31 M.I. Lembke, L. Oberli, G. Olschewski, R. Dotti ACCIAI INOX Laser welding of plastically deformed lean duplex stainless steel pag. 1-5 I. Calliari, C. Gennari, E. Hurtado Delgado, A. F. Miranda Pérez, B. R. Rodriguez Vargas Martensite quantification, mechanical properties and cold rolling in AISI 301 Austenitic Stainless Steel pag.1-11 P. Piccardo, R. Spotorno, D. Lanteri, F. Canepa, I. Citi Comparative investigation of deep drawing formability in austenitic (AISI 321) and in ferritic (DIN 1.4509) stainless steel sheets pag.1-50 C. de Paula Camargo Pisano, H. J. B. Alves, T. Reis de Oliveira, C. G. Schön

La simulazione virtuale applicata ai trattamenti termici di forgiati pag.9-46 E. Anelli, M. Lucchesi, A. Chugaeva

Comparison Between Microstructures, Deformation Mechanisms and Micromechanical Properties of 316L Stainless Steel Consolidated by Laser Melting pag.3-5 I. Heikkilä , O. Karlsson, D. Lindela, A. Angra, Y. Zhong, J. Olsén

Effetto del carico ciclico sulla diffusione di idrogeno in acciai bassolegati ad alto limite di snervamento pag.10-26 M. Cabrini, S. Lorenzi, T. Pastore, D. Pesenti Bucella

Influence of thermal aging on scc susceptibility of DSS 2304 in the presence of chlorides and thiosulphates pag.3-40 F. Zanotto, V. Grassi, A. Balbo, C. Monticelli, F. Zucchi

ACCIAIO INOX AUSTENITICO

Continuous annealing by resistance heating furnace of stainless steels wire drawn pag.6-48 R. Nemfardi, R. Bedini, G. Zucchelli, D. Bonora, C. Mapelli, S. Barella, A. Gruttadauria, D. Mombelli

Different effects of carbon and nitrogen on precipitation behavior and mechanical properties in austenitic stainless steels pag.1-28 Kyung-Shik Kim, Jee-Hyun Kang, Sung-Joon Kim Strain induced martensite evolution in a rolling contact of SS AISI 304 pag.1-62 M. Werschler, P. Gümpel, K. Werner Strong Potential of Commercialized High Mn Steel Products and Process for Various Applications pag.7/8-69 J. Choi 82

Nitrurazione di acciai inossidabili: processo combinato ionico-gassoso pag.9-30 R. Giovanardi, P. Veronesi, R. Sola, L. Borg, G. Parigi Nitrurazione in scarica ionica a bassa temperatura di acciai inossidabili austenitici pag.10-32 F. Borgioli, E. Galvanetto, T. Bacci

La Metallurgia Italiana - n. 11/12 2018

Atti e notizie AFFINAZIONE Analisi del comportamento a corrosione di campioni di alluminio AA6012 sottoposti ad ECAP e trattamento criogenico pag.2-25 A. Viceré, M. Cabibbo, C. Paoletti, G. Roventi, T. Bellezze ALLUMINIO E LEGHE Influence of Mg and Ti on both eutectic solidification and modifying efficiency in Sr-modified Al-7Si cast alloys pag.2-5 L. Lattanzi, A. Fortini, M. Giovagnoli, M. Merlin Theoretical and Experimental Evaluation of the Effectiveness of Aluminum Melt Treatment by Physical Methods pag.2-16 V.B. Deev , E.S. Prusov , A.I. Kutsenko Analisi del comportamento a corrosione di campioni di alluminio AA6012 sottoposti ad ECAP e trattamento criogenico pag.2-25 A. Viceré, M. Cabibbo, C. Paoletti, G. Roventi, T. Bellezze A comparative cradle-to gate impact assessment: primary and secondary aluminum automotive components case pag.2-46 S. Cecchel, M. Collotta, G. Cornacchia, A. Panvini, G. Tomasoni Metallografia applicata alla failure analysis in ambito automotive: influenza della microstruttura sulle proprietà meccaniche di una lega di alluminio da fonderia pag.5/5 R. Rolli, D. Rossi Leghe di alluminio autoinvecchianti per applicazioni nel settore automobilistico pag.5-19 C. Castella, I. Peter, S. Lombardo, M. Rosso, R. Molina Studio del processo di anodizzazione per la protezione dalla corrosione di schiume metalliche pag.10-12 S. Rossi, M. Fedel, F. Deflorian, M. Bizzotto AMBIENTE E SICUREZZA A comparative cradle-to gate impact assessment: primary and secondary aluminum automotive components case pag.2-46 S. Cecchel, M. Collotta, G. Cornacchia, A. Panvini, G. Tomasoni Water leak detection in EAf based on Tenova’s offgas technology: recent developments and results in lucchini RS, Lovere, Italy pag.7/8-58 M. Luccini, V. Scipolo, D. Zuliani, L. Poli, D. Masoero AUTOMOTIVE A comparative cradle-to gate impact assessment: primary and secondary aluminum automotive components case pag.2-46 S. Cecchel, M. Collotta, G. Cornacchia, A. Panvini, G. Tomasoni Leghe di alluminio autoinvecchianti per applicazioni nel settore automobilistico pag.5-19 C. Castella, I. Peter, S. Lombardo, M. Rosso, R. Molina

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Tribology Test Methods and Simulations of the Effect of Friction on the Formability of Automotive Steel Sheets pag.9-56 S. Jadhav, M. Schoiswohl, B. Buchmayr BIOMATERIALI Microstructural characterization and corrosion behaviour of SLM CoCrMo alloy in simulated body fluid pag.3-45 M. Seyedi, F. Zanotto, E. Liverani, A. Fortunato, C. Monticelli, A. Balbo CARATTERIZZAZIONE MATERIALI Different effects of carbon and nitrogen on precipitation behavior and mechanical properties in austenitic stainless steels pag.1-28 Kyung-Shik Kim, Jee-Hyun Kang, Sung-Joon Kim Comparative investigation of deep drawing formability in austenitic (AISI 321) and in ferritic (DIN 1.4509) stainless steel sheets pag.1-50 C. de Paula Camargo Pisano, H. J. B. Alves, T. Reis de Oliveira, C. G. Schön Surface Oxide State on Metal Powder and its Changes during Additive Manufacturing: an Overview pag.3-34 E. Hryha, R. Shvab, H. Gruber, A. Leicht, L. Nyborg Microstructural characterization and corrosion behaviour of SLM CoCrMo alloy in simulated body fluid pag.3-45 M. Seyedi, F. Zanotto, E. Liverani, A. Fortunato, C. Monticelli, A. Balbo Leghe di alluminio autoinvecchianti per applicazioni nel settore automobilistico pag.5-19 C. Castella, I. Peter, S. Lombardo, M. Rosso, R. Molina Caratterizzazione della resistenza al rinvenimento di un acciaio per lavorazioni a caldo AISI H13 modificato pag.9-23 F. S. Gobber, D. Ugues, P. Martelli, M. Rosso Effetto della velocità di raffreddamento dopo solubilizzazione nei trattamenti termici T6 e T7 sulle proprietà meccaniche della lega EN AC-42000 pag.9-37 D. Cico, E. Gianotti Development of new tests to assess sulfide stress corrosion cracking of line pipes pag. 6-11 M. Cabrini, S. Lorenzi, T. Pastore, D. Pesenti Bucella, G. Pellegrini, A. Paggi, E. Paravicini Bagliani, P. Darcis CARATTERIZZAZIONE DELLE POLVERI Lessons Learnt Through the Development of an Application-Specific Methodology for Metal Powder Characterisation for Additive Manufacturing pag.2-20 M. Meisnar, S. Baker, C. Fowler, L. Pambaguian, T. Ghidini CINETICA DI RIDUZIONE Reaction kinetics of molten iron oxides reduction 83

Aim news using hydrogen pag.7/8-5 M. Naseri Seftejani, J. Schenk CONTROLLO PROCESSI

Muffle tubes: choose the right material for longer service life pag.6-22 A. Spaghetti

Hot-rolled steels for special purposes: high carbon wire-rod for cold-drawing pag.6-38 N. Bolognani, A. Parimbelli

Nitrurazione di acciai inossidabili: processo combinato ionico-gassoso pag.9-30 R. Giovanardi, P. Veronesi, R. Sola, L. Borg, G. Parigi

Latest results in EAF optimization of scrap-based melting process: Q-MELT installation in Kroman Celik pag.7/8-41 M. Ansoldi, D. Patrizio, M. Piazza, O. Kuran

Trattamenti di anodizzazione su titanio commercialmente puro e resistenza a corrosione pag.10-5 D. Prando, S. Beretta, F. Bolzoni, A. Brenna, M.V. Diamanti, M. Pedeferri, M. Ormellese

Water leak detection in EAf based on Tenova’s offgas technology: recent developments and results in lucchini RS, Lovere, Italy pag.7/8-58 M. Luccini, V. Scipolo, D. Zuliani, L. Poli, D. Masoero

Studio del processo di anodizzazione per la protezione dalla corrosione di schiume metalliche pag.10-12 S. Rossi, M. Fedel, F. Deflorian, M. Bizzotto

CORROSIONE New light corrosion resistant steel without chromium pag.1-38 C. Mapelli, A. F. Ciuffini, S. Barella, A. Gruttadauria, D. Mombelli Analisi del comportamento a corrosione di campioni di alluminio AA6012 sottoposti ad ECAP e trattamento criogenico pag.2-25 A. Viceré, M. Cabibbo, C. Paoletti, G. Roventi, T. Bellezze Electrochemical corrosion behaviour of binary magnesium - heavy rare earth alloys pag.2-34 F. Rosalbino, S. De Negri, G. Scavino, A. Saccone Influence of thermal aging on scc susceptibility of DSS 2304 in the presence of chlorides and thiosulphates pag.3-40 F. Zanotto, V. Grassi, A. Balbo, C. Monticelli, F. Zucchi Microstructural characterization and corrosion behaviour of SLM CoCrMo alloy in simulated body fluid pag.3-45 M. Seyedi, F. Zanotto, E. Liverani, A. Fortunato, C. Monticelli, A. Balbo Study of the conservation state of european street furniture in painted cast irons pag.4-5 C. Soffritti, L. Calzolari, A. Balbo, F. Zanotto, C. Monticelli, A. Fortini, G.L. Garagnani Indoor and outdoor atmospheric corrosion monitoring of cultural heritage assets pag.4-34 E. Angelini, C. E. Arroyave Posada, E. Di Francia, S. Grassini, L. Iannucci, L. Lombardo, M. Parvis Role of the stabilizing elements in the ferritic stainless steel on the microstructural, corrosion properties and durability of the stainless steel weld joints pag.5-11 D. Dyja, R. Marek, S. Zbigniew Development of new tests to assess sulfide stress corrosion cracking of line pipes pag. 6-11 M. Cabrini, S. Lorenzi, T. Pastore, D. Pesenti Bucella, G. Pellegrini, A. Paggi, E. Paravicini Bagliani, P. Darcis 84

Effetto della finitura superficiale sul rilascio in soluzione di leghe a base rame utilizzate nell’industria di trafilatura della pasta pag.10-20 G. Ghiara, C. Usai, R. Spotorno, P. Piccardo, A. Adobati Nitrurazione in scarica ionica a bassa temperatura di acciai inossidabili austenitici pag.10-32 F. Borgioli, E. Galvanetto, T. Bacci Trattamenti superficiali ecocompatibili dell’alluminio per l’industria aeronautica pag.10-45 A. Carangelo, A. Acquesta, T. Monetta CREEP Development of new ferritic stainless steel sheets with excellent heat resistance and formability for automotive exhaust manifolds pag.5-33 J-i. Hamada, N. Kanno, A. Hayashi, N. Hiraide, M. Abe, K. Nishimura, C. Takushima, A. Yakawa, F. Fudanoki CREEP Development of new ferritic stainless steel sheets with excellent heat resistance and formability for automotive exhaust manifolds pag.5-33 J-i. Hamada, N. Kanno, A. Hayashi, N. Hiraide, M. Abe, K. Nishimura, C. Takushima, A. Yakawa, F. Fudanoki DEFORMAZIONI PLASTICHE Laser welding of plastically deformed lean duplex stainless steel pag. 1-5 I. Calliari, C. Gennari, E. Hurtado Delgado, A. F. Miranda Pérez, B. R. Rodriguez Vargas Martensite quantification, mechanical properties and cold rolling in AISI 301 Austenitic Stainless Steel pag.1-11 P. Piccardo, R. Spotorno, D. Lanteri, F. Canepa, I. Citi DIFETTI Pulizia in linea del filo metallico su impianti di trafilatura ad alta velocita’: innovazione, ecologia e performance pag. 6-44 G. Orlando, R. Giovanardi, L. Trombi

La Metallurgia Italiana - n. 11/12 2018

Atti e notizie ECOLOGIA

FUSIONE LASER

Pulizia in linea del filo metallico su impianti di trafilatura ad alta velocita’: innovazione, ecologia e performance pag. 6-44 G. Orlando, R. Giovanardi, L. Trombi

ELETTROCHIMICA

Impact of Process Conditions on the Properties of Additively Manufactured Tool Steel H13 processed by LBM pag.3-12 L. Wu, T.Klaas, S. Leuders, F. Brenne, T. Niendorf

Pulizia in linea del filo metallico su impianti di trafilatura ad alta velocita’: innovazione, ecologia e performance pag. 6-44 G. Orlando, R. Giovanardi, L. Trombi Effetto del carico ciclico sulla diffusione di idrogeno in acciai bassolegati ad alto limite di snervamento pag.10-26 M. Cabrini, S. Lorenzi, T. Pastore, D. Pesenti Bucella ENERGIA Caratterizzazione microstrutturale ed elettrica di Cu1.2Mn1.8O4 applicato come rivestimento protettivo in pile a combustibile ad ossidi solidi (SOFC) pag.1-44 R.Spotorno, P.Piccardo, S. Barison, S. Fasolin FATICA TERMICA Development of new ferritic stainless steel sheets with excellent heat resistance and formability for automotive exhaust manifolds pag.5-33 J-i. Hamada, N. Kanno, A. Hayashi, N. Hiraide, M. Abe, K. Nishimura, C. Takushima, A. Yakawa, F. Fudanoki FONDERIA Influence of Mg and Ti on both eutectic solidification and modifying efficiency in Sr-modified Al-7Si cast alloys pag.2-5 L. Lattanzi, A. Fortini, M. Giovagnoli, M. Merlin Metallografia applicata alla failure analysis in ambito automotive: influenza della microstruttura sulle proprietà meccaniche di una lega di alluminio da fonderia pag.5/5 R. Rolli, D. Rossi FORNO ELETTRICO AD ARCO Bypassing Problems Related to Water Cooling: Case Study for Applying ILTEC in a 100-ton EAF pag.7/8-31 M. B. Hanel, A. Filzwieser, R. Degel FRATTURA Hot ductility behavior of a high alloy steel pag.7/8-15 J.-H. Min, S.-H. Kwon, S.-D. Lee, S.-W. Moon, D-.K. Kim, J.-S. Lee, Y.-U. Heo, C.-H. Yim FUSIONE Latest results in EAF optimization of scrap-based melting process: Q-MELT installation in Kroman Celik pag.7/8-41 M. Ansoldi, D. Patrizio, M. Piazza, O. Kuran La Metallurgia Italiana - n. 11/12 2018

In-situ Micro-tensile Testing of Additive Manufactured Maraging Steels in the SEM: Influence of Build Orientation, Thickness and Roughness on the Resulting Mechanical Properties pag.3-27 K. B. Surreddi, C. Oikonomou, P. Karlsson, M. Olsson, L. Pejryd GHISA Study of the conservation state of european street furniture in painted cast irons pag.4-5 C. Soffritti, L. Calzolari, A. Balbo, F. Zanotto, C. Monticelli, A. Fortini, G.L. Garagnani GRANO ULTRAFINE Surpassing steel performance by creating a very fine grained microstructure pag.6-31 M.I. Lembke, L. Oberli, G. Olschewski, R. Dotti IMPIANTI E ATTREZZATURE Latest results in EAF optimization of scrap-based melting process: Q-MELT installation in Kroman Celik pag.7/8-41 M. Ansoldi, D. Patrizio, M. Piazza, O. Kuran IMPIEGHI ALTA TEMPERATURA Caratterizzazione microstrutturale ed elettrica di Cu1.2Mn1.8O4 applicato come rivestimento protettivo in pile a combustibile ad ossidi solidi (SOFC) pag.1-44 R.Spotorno, P.Piccardo, S. Barison, S. Fasolin LAMINAZIONE Martensite quantification, mechanical properties and cold rolling in AISI 301 Austenitic Stainless Steel pag.1-11 P. Piccardo, R. Spotorno, D. Lanteri, F. Canepa, I. Citi Hot-rolled steels for special purposes: high carbon wire-rod for cold-drawing pag.6-38 N. Bolognani, A. Parimbelli MAGNESIO E LEGHE Electrochemical corrosion behaviour of binary magnesium - heavy rare earth alloys pag.2-34 F. Rosalbino, S. De Negri, G. Scavino, A. Saccone METALLI LEGGERI Effetto della velocità di raffreddamento dopo solu85

Aim news bilizzazione nei trattamenti termici T6 e T7 sulle proprietà meccaniche della lega EN AC-42000 pag.9-37 D. Cico, E. Gianotti Studio del processo di anodizzazione per la protezione dalla corrosione di schiume metalliche pag.10-12 S. Rossi, M. Fedel, F. Deflorian, M. Bizzotto

MODELLAZIONE Statistical modeling of heat exchange coefficient evolution during quenching pag.9-5 M. Clenet, M. Bentivegni, G. Druart NITRURAZIONE

Trattamenti superficiali ecocompatibili dell’alluminio per l’industria aeronautica pag.10-45 A. Carangelo, A. Acquesta, T. Monetta

Future trends in gaseous surface hardening of titanium and titanium alloys pag.9-13 T. L. Christiansen, M. S. Jellesen, M. A. J. Somers

METALLOGRAFIA

NORMATIVA INTERNAZIONALE

Martensite quantification, mechanical properties and cold rolling in AISI 301 Austenitic Stainless Steel pag.1-11 P. Piccardo, R. Spotorno, D. Lanteri, F. Canepa, I. Citi

Lessons Learnt Through the Development of an Application-Specific Methodology for Metal Powder Characterisation for Additive Manufacturing pag.2-20 M. Meisnar, S. Baker, C. Fowler, L. Pambaguian, T. Ghidini

Metallografia applicata alla failure analysis in ambito automotive: influenza della microstruttura sulle proprietà meccaniche di una lega di alluminio da fonderia pag.5/5 R. Rolli, D. Rossi

OSSIDAZIONE

METALLURGIA DELLE POLVERI Surface Oxide State on Metal Powder and its Changes during Additive Manufacturing: an Overview pag.3-34 E. Hryha, R. Shvab, H. Gruber, A. Leicht, L. Nyborg MICROSCOPIA ELETTRONICA Indoor and outdoor atmospheric corrosion monitoring of cultural heritage assets pag.4-34 E. Angelini, C. E. Arroyave Posada, E. Di Francia, S. Grassini, L. Iannucci, L. Lombardo, M. Parvis MICROSTRUTTURA E COMPOSIZIONE Different effects of carbon and nitrogen on precipitation behavior and mechanical properties in austenitic stainless steels pag.1-28 Kyung-Shik Kim, Jee-Hyun Kang, Sung-Joon Kim Influence of Mg and Ti on both eutectic solidification and modifying efficiency in Sr-modified Al-7Si cast alloys pag.2-5 L. Lattanzi, A. Fortini, M. Giovagnoli, M. Merlin Theoretical and Experimental Evaluation of the Effectiveness of Aluminum Melt Treatment by Physical Methods pag.2-16 V.B. Deev , E.S. Prusov , A.I. Kutsenko Study of the conservation state of european street furniture in painted cast irons pag.4-5 C. Soffritti, L. Calzolari, A. Balbo, F. Zanotto, C. Monticelli, A. Fortini, G.L. Garagnani Metallografia applicata alla failure analysis in ambito automotive: influenza della microstruttura sulle proprietà meccaniche di una lega di alluminio da fonderia pag.5/5 R. Rolli, D. Rossi 86

New light corrosion resistant steel without chromium pag.1-38 C. Mapelli, A. F. Ciuffini, S. Barella, A. Gruttadauria, D. Mombelli Development of new ferritic stainless steel sheets with excellent heat resistance and formability for automotive exhaust manifolds pag.5-33 J-i. Hamada, N. Kanno, A. Hayashi, N. Hiraide, M. Abe, K. Nishimura, C. Takushima, A. Yakawa, F. Fudanoki Studio del processo di anodizzazione per la protezione dalla corrosione di schiume metalliche pag.10-12 S. Rossi, M. Fedel, F. Deflorian, M. Bizzotto PROCESSI PRODUTTIVI SPECIALI Strong Potential of Commercialized High Mn Steel Products and Process for Various Applications pag.7/8-69 J. Choi PRODUZIONE D’ACCIAIO CON CONVERTITORI AD OSSIGENO A new approach for modelling and control of dephosphorization in BOF converter pag.pag.11/12-5 S. Khadhraoui, H.-J. Odenthal, S. Das, M. Schlautmann, K. Hack, B. Glaser, R. Woolf Conventional steelmaking route being subject to transformation pag.pag.11/12-17 P.Argenta, M. Marcozzi, M. Dorndorf, P. Duarte, S. Maggiolino, O. Demir Estimation of trace elements in unalloyed scrap pag. pag.11/12-26 R. Ammer, P. Reisinger, M. Egger, R. Tober, G. Salzmann Dissolution behaviour of Ulc steel in carbon saturated hot metal pag.pag.11/12-36 F. Markus Penz, J. Schenk, R. Ammer, K. Pastucha, B. Maunz Effect of converter dart design aspects on its settling behaviour pag.pag.11/12-46 J. van Oord, H.J. Visser La Metallurgia Italiana - n. 11/12 2018

Atti e notizie Reaction near hotspot in refining of molten iron during supplying top-blown oxygen and powder additives pag.pag.11/12-58 Yu-ichi UCHIDA Development and application of high efficiency blowing technology of converters in shougang pag. pag.11/12-67 Pan GAO, Xinhua WANG, Haibo LI, Yuming GUO, Yong ZHANG, Xiaodong ZHAO, Huixiang YU PROPRIETA’ MECCANICHE Continuous annealing by resistance heating furnace of stainless steels wire drawn pag.6-48 R. Nemfardi, R. Bedini, G. Zucchelli, D. Bonora, C. Mapelli, S. Barella, A. Gruttadauria, D. Mombelli RAME E LEGHE Indoor and outdoor atmospheric corrosion monitoring of cultural heritage assets pag.4-34 E. Angelini, C. E. Arroyave Posada, E. Di Francia, S. Grassini, L. Iannucci, L. Lombardo, M. Parvis Effetto della finitura superficiale sul rilascio in soluzione di leghe a base rame utilizzate nell’industria di trafilatura della pasta pag.10-20 G. Ghiara, C. Usai, R. Spotorno, P. Piccardo, A. Adobati

M. B. Hanel, A. Filzwieser, R. Degel SIMULAZIONE NUMERICA La simulazione virtuale applicata ai trattamenti termici di forgiati pag.9-46 E. Anelli, M. Lucchesi, A. Chugaeva STORIA DELLA METALLURGIA Massi-frantoio rinvenuti in scavo presso la miniera di magnetite di Pietra Bianca 2 (Biella) pag.4-17 A. Gattiglia, M. Rossi, P. de Vingo TECNOLOGIE Massi-frantoio rinvenuti in scavo presso la miniera di magnetite di Pietra Bianca 2 (Biella) pag.4-17 A. Gattiglia, M. Rossi, P. de Vingo TECNOLOGIE ADDITIVE Lessons Learnt Through the Development of an Application-Specific Methodology for Metal Powder Characterisation for Additive Manufacturing pag.2-20 M. Meisnar, S. Baker, C. Fowler, L. Pambaguian, T. Ghidini

RICOTTURA

Continuous annealing by resistance heating furnace of stainless steels wire drawn pag.6-48 R. Nemfardi, R. Bedini, G. Zucchelli, D. Bonora, C. Mapelli, S. Barella, A. Gruttadauria, D. Mombelli

Impact of Process Conditions on the Properties of Additively Manufactured Tool Steel H13 processed by LBM pag.3-12 L. Wu, T.Klaas, S. Leuders, F. Brenne, T. Niendorf

RIVESTIMENTI

Caratterizzazione microstrutturale ed elettrica di Cu1.2Mn1.8O4 applicato come rivestimento protettivo in pile a combustibile ad ossidi solidi (SOFC) pag.1-44 R.Spotorno, P.Piccardo, S. Barison, S. Fasolin SALDATURA

Surface Oxide State on Metal Powder and its Changes during Additive Manufacturing: an Overview pag.3-34 E. Hryha, R. Shvab, H. Gruber, A. Leicht, L. Nyborg

Laser welding of plastically deformed lean duplex stainless steel pag. 1-5 I. Calliari, C. Gennari, E. Hurtado Delgado, A. F. Miranda Pérez, B. R. Rodriguez Vargas

Properties Of Ti-6Al-4V Components Produced By Digital Metal® Binder Jetting Technology pag.3-52 M. Persson, R. Carlström, K. Gustavsson, S. Nilsson, C. Palmqvist, B. Brash

Role of the stabilizing elements in the ferritic stainless steel on the microstructural, corrosion properties and durability of the stainless steel weld joints pag.5-11 D. Dyja, R. Marek, S. Zbigniew

TITANIO E LEGHE

SOLIDIFICAZIONE

Properties Of Ti-6Al-4V Components Produced By Digital Metal® Binder Jetting Technology pag.3-52 M. Persson, R. Carlström, K. Gustavsson, S. Nilsson, C. Palmqvist, B. Brash

The interfacial convection in fluxes in the continuous casting process pag.7/8-23 P. R. Scheller, Yong Lin, Q. Shu

Future trends in gaseous surface hardening of titanium and titanium alloys pag.9-13 T. L. Christiansen, M. S. Jellesen, M. A. J. Somers

SICUREZZA Bypassing Problems Related to Water Cooling: Case Study for Applying ILTEC in a 100-ton EAF pag.7/8-31

Trattamenti di anodizzazione su titanio commercialmente puro e resistenza a corrosione pag.10-5 D. Prando, S. Beretta, F. Bolzoni, A. Brenna, M.V. Diamanti, M. Pedeferri, M. Ormellese

La Metallurgia Italiana - n. 11/12 2018

Aim news TRAFILATURA Comparative investigation of deep drawing formability in austenitic (AISI 321) and in ferritic (DIN 1.4509) stainless steel sheets pag.1-50 C. de Paula Camargo Pisano, H. J. B. Alves, T. Reis de Oliveira, C. G. Schön Defects, their source and detection in wire drawn products pag.6-5 C. Mapelli, S. Barella, A. Gruttadauria, D. Mombelli, A. F. Ciuffini, M. Cusolito Pulizia in linea del filo metallico su impianti di trafilatura ad alta velocita’: innovazione, ecologia e performance pag. 6-44 G. Orlando, R. Giovanardi, L. Trombi

Influence of thermal aging on scc susceptibility of DSS 2304 in the presence of chlorides and thiosulphates pag.3-40 F. Zanotto, V. Grassi, A. Balbo, C. Monticelli, F. Zucchi Hot-rolled steels for special purposes: high carbon wire-rod for cold-drawing pag.6-38 N. Bolognani, A. Parimbelli Caratterizzazione della resistenza al rinvenimento di un acciaio per lavorazioni a caldo AISI H13 modificato pag.9-23 F. S. Gobber, D. Ugues, P. Martelli, M. Rosso Nitrurazione di acciai inossidabili: processo combinato ionico-gassoso pag.9-30 R. Giovanardi, P. Veronesi, R. Sola, L. Borg, G. Parigi

Tribology Test Methods and Simulations of the Effect of Friction on the Formability of Automotive Steel Sheets pag.9-56 S. Jadhav, M. Schoiswohl, B. Buchmayr

Effetto della velocità di raffreddamento dopo solubilizzazione nei trattamenti termici T6 e T7 sulle proprietà meccaniche della lega EN AC-42000 pag.9-37 D. Cico, E. Gianotti

TRATTAMENTI SUPERFICIALI

La simulazione virtuale applicata ai trattamenti termici di forgiati pag.9-46 E. Anelli, M. Lucchesi, A. Chugaeva

Nitrurazione di acciai inossidabili: processo combinato ionico-gassoso pag.9-30 R. Giovanardi, P. Veronesi, R. Sola, L. Borg, G. Parigi Trattamenti di anodizzazione su titanio commercialmente puro e resistenza a corrosione pag.10-5 D. Prando, S. Beretta, F. Bolzoni, A. Brenna, M.V. Diamanti, M. Pedeferri, M. Ormellese Effetto della finitura superficiale sul rilascio in soluzione di leghe a base rame utilizzate nell’industria di trafilatura della pasta pag.10-20 G. Ghiara, C. Usai, R. Spotorno, P. Piccardo, A. Adobati Nitrurazione in scarica ionica a bassa temperatura di acciai inossidabili austenitici pag.10-32 F. Borgioli, E. Galvanetto, T. Bacci Trattamenti superficiali ecocompatibili dell’alluminio per l’industria aeronautica pag.10-45 A. Carangelo, A. Acquesta, T. Monetta TRATTAMENTI TERMICI Mechanical properties evolution on heat treated severe cold rolled UNS S32760 Super Duplex Stainless Steel pag.1-20 C.M. Tromellini, A.F. Ciuffini, A. Gruttadauria, S. Barella, C. Di Cecca, C. Mapelli

TRIBOLOGIA Strain induced martensite evolution in a rolling contact of SS AISI 304 pag.1-62 M. Werschler, P. Gümpel, K. Werner Tribology Test Methods and Simulations of the Effect of Friction on the Formability of Automotive Steel Sheets pag.9-56 S. Jadhav, M. Schoiswohl, B. Buchmayr VALUTAZIONE MATERIALI Electrochemical corrosion behaviour of binary magnesium - heavy rare earth alloys pag.2-34 F. Rosalbino, S. De Negri, G. Scavino, A. Saccone Effetto del carico ciclico sulla diffusione di idrogeno in acciai bassolegati ad alto limite di snervamento pag.10-26 M. Cabrini, S. Lorenzi, T. Pastore, D. Pesenti Bucella Trattamenti superficiali ecocompatibili dell’alluminio per l’industria aeronautica pag.10-45 A. Carangelo, A. Acquesta, T. Monetta

La Metallurgia Italiana - n. 11/12 2018

Atti e notizie Indice autori Abe, M. pag.5-33 Acquesta, A. pag.10-45 Adobati, A. pag.10-20 Alves, H. J. B. pag.1-50 Ammer, R. pag.11/12-26 Anelli, E. pag.9-46 Angelini, E. pag.4-34 Angra, A. pag.3-5 Ansoldi, M. pag.7/8-41 Argenta, P. pag.11/12-17 Arroyave Posada, C. E. pag.4-34 Bacci, T. pag.10-32 Baker, S. pag.2-20 Balbo, A. pag.3-40; pag.4-5 Barella, S. pag.1-38; pag.1-20; pag.6-5; pag.6-48 Barison, S. pag.1-44 Bedini, R. pag.6-48 Bellezze, T. pag.2-25 Bentivegni, M. pag.9-5 Beretta, S. pag.10-5 Bizzotto, M. pag.10-12 Bolognani, N. pag.6-38 Bolzoni, F. pag.10-5 Bonora, D. pag.6-48 Borg, L. pag.9-30 Borgioli, F. pag.10-32 Brash, B. pag.3-52 Brenna, A. pag.10-5 Brenne, F. pag.3-12 Buchmayr, B. pag.9-56 C. de Paula Camargo Pisano pag.1-50 Cabibbo, M. pag.2-25 Cabrini, M. pag. 6-11 Calliari, I. pag. 1-5 Calzolari, L. pag.4-5 Canepa, F. pag.1-11 Carangelo, A. pag.10-45 Carlström, R. pag.3-52 Castella, C. pag.5-19 Cecchel, S. pag.2-46 Choi, J. pag.7/8-69 Christiansen, T. L. pag.9-13 Chugaeva, A. pag.9-46 Cico, D. pag.9-37 Citi, I. pag.1-11 Ciuffini, A. F. pag.1-38 Clenet, M. pag.9-5 Collotta, M. pag.2-46 Cornacchia, G. pag.2-46 Cusolito, M. pag.6-5 Darcis, P. pag. 6-11 Das, S. pag.11/12-25 De Negri, S. pag.2-34 Deev, V.B. pag.2-16 Deflorian, F. pag.10-12 Degel, R. pag.7/8-31 Demir, O. pag.11/12-17 Di Cecca, C. pag.1-20 Di Francia, E. pag.4-34 La Metallurgia Italiana - n. 11/12 2018

Diamanti, M.V. pag.10-5 Dorndorf, M. pag.11/12-17 Dotti, R. pag.6-31 Druart, G. pag.9-5 Duarte, P. pag.11/12-17 Dyja, D. pag.5-11 Egger, M. pag.11/12-26 Fasolin, S. pag.1-44 Fedel, M. pag.10-12 Filzwieser, A. pag.7/8-31 Fortini, A. pag.2-5; pag.4-5 Fortunato, A. pag.3-45 Fowler, C. pag.2-20 Fudanoki, F. pag.5-33 Galvanetto, E. pag.10-32 Gao, P. pag.11/12-67 Garagnani, G.L. pag.4-5 Gattiglia, A. pag.4-17 Gennari, C. pag. 1-5 Ghiara, G. pag.10-20 Ghidini, T. pag.2-20 Gianotti, E. pag.9-37 Giovagnoli, M. pag.2-5 Giovanardi, R. pag. 6-44 Glaser, B. pag.11/12-5 Gobber, F. S. pag.9-23 Grassi, V. pag.3-40 Grassini, S. pag.4-34 Gruber, H. pag.3-34 Gruttadauria, A. pag.1-20; pag.1-38; pag.6-48; pag.6-5 Gümpel, P. pag.1-62 Guo, Y. pag.11/12-67 Gustavsson, K. pag.3-52 Hack, K. pag.11/12-5 Hamada, J-I. pag.5-33 Hanel, M. B. pag.7/8-31 Hayashi, A. pag.5-33 Heikkilä, I. pag.3-5 Heo, Y.-U. pag.7/8-15 Hiraide, N. pag.5-33 Hryha, E. pag.3-34 Hurtado Delgado, E. pag. 1-5 Iannucci, L. pag.4-34 Jadhav, S. pag.9-56 Jee-Hyun, K. pag.1-28 Jellesen, M. S. pag.9-13 Kanno, N. pag.5-33 Karlsson, O. pag.3-5 Karlsson, P. pag.3-27 Khadhraoui, S. pag.11/12-5 Kim, D-.K. pag.7/8-15 Klaas, T. pag.3-12 Kuran, O. pag.7/8-41 Kutsenko, A.I. pag.2-16 Kwon, S.-H. pag.7/8-15 Kyung-Shik, K. pag.1-28 Lanteri, D. pag.1-11 Lattanzi, L. pag.2-5 Lee, J.-S. pag.7/8-15 89

Aim news Lee, S.-D. pag.7/8-15 Leicht, A. pag.3-34 Lembke, M.I. pag.6-31 Leuders, S. pag.3-12 Li, H. pag.11/12-67 Lindela, D. pag.3-5 Liverani, E. pag.3-45 Lombardo, L. pag.4-34 Lombardo, S. pag.5-19 Lorenzi, S. pag. 6-11; pag.10-26 Lucchesi, M. pag.9-46 Luccini, M. pag.7/8-58 Maggiolino, S. pag.11/12-17 Mapelli, C. pag.1-20; pag.1-38; pag.6-5; pag.6-48 Marcozzi, M. pag.11/12-36 Marek, R. pag.5-11 Markus Penz, F. pag.11/12-36 Martelli, P. pag.9-23 Masoero, D. pag.7/8-58 Maunz, B. pag.11/12-36 Meisnar, M. pag.2-20 Merlin, M. pag.2-5 Min, J.-H. pag.7/8-15 Miranda Pérez, A. F. pag. 1-5 Molina, R. pag.5-19 Mombelli, D. pag.1-38; pag.6-5; pag.6-48 Monetta, T. pag.10-45 Monticelli, C. pag.3-40; pag.4-5 Moon, S.-W. pag.7/8-15 Naseri Seftejani, M. pag.7/8-5 Nemfardi, R. pag.6-48 Niendorf, T. pag.3-12 Nilsson, S. pag.3-52 Nishimura, K. pag.5-33 Nyborg, L. pag.3-34 Oberli, L. pag.6-31 Odenthal, H.-J. pag.11/12-5 Oikonomou, C. pag.3-27 Olschewski, G. pag.6-31 Olsén, J. pag.3-5 Olsson, M. pag.3-27 Orlando, G. pag. 6-44 Ormellese, M. pag.10-5 P. de Vingo pag.4-17 Paggi, A. pag. 6-11 Palmqvist, C. pag.3-52 Pambaguian, L. pag.2-20 Panvini, A. pag.2-46 Paoletti, C. pag.2-25 Paravicini Bagliani, E. pag. 6-11 Parigi, G. pag.9-30 Parimbelli, A. pag.6-38 Parvis, M. pag.4-34 Pastore, T. pag.10-26 Pastucha, K. pag.11/12-36 Patrizio, D. pag.7/8-41 Pedeferri, M. pag.10-5 Pejryd, L. pag.3-27 Pellegrini, G. pag. 6-11 Persson, M. pag.3-52 Pesenti Bucella, D. pag.10-26 Peter, I. pag.5-19 Piazza, M. pag.7/8-41 90

Piccardo, P. pag.1-11; pag.1-44; pag.10-20 Poli, L. pag.7/8-58 Prando, D. pag.10-5 Prusov, E.S. pag.2-16 Reis de Oliveira, T. pag.1-50 Reisinger, P. pag.11/12-26 Rodriguez Vargas, B. R. pag. 1-5 Rolli, R. pag.5/5 Rosalbino, F. pag.2-34 Rossi, D. pag.5/5 Rossi, M. pag.4-17 Rossi, S. pag.10-12 Rosso, M. pag.9-23 Roventi, G. pag.2-25 Saccone, A. pag.2-34 Salzmann, G. pag.11/12-26 Scavino, G. pag.2-34 Scheller, P. R. pag.7/8-23 Schenk, J. pag.7/8-5 Schlautmann, M. pag.11/12-5 Schoiswohl, M. pag.9-56 Schön, C. G. pag.1-50 Scipolo, V. pag.7/8-58 Seyedi, M. pag.3-45 Shu, Q. pag.7/8-23 Shvab, R. pag.3-34 Soffritti, C. pag.4-5 Sola, R. pag.9-30 Somers, M. A. J. pag.9-13 Spaghetti, A. pag.6-22 Spotorno, R. pag.1-11; pag.10-20 Sung-Joon, K. pag.1-28 Surreddi, K. B. pag.3-27 Takushima, C. pag.5-33 Tober, R. pag.11/12-26 Tomasoni, G. pag.2-46 Trombi, L. pag.6-44 Tromellini, C.M. pag.1-20 Uchida, Y.- I. pag.11/12-58 Ugues, D. pag.9-23 Usai, C. pag.10-20 Van Oord, J. pag.11/12-46 Veronesi, P. pag.9-30 Viceré, A. pag.2-25 Visser, H.J. pag.11/12-46 Wang, X. pag.11/12-67 Werner, K. pag.1-62 Werschler, M. pag.1-62 Woolf, R. pag.11/12-5 Wu, L. pag.3-12 Yakawa, A. pag.5-33 Yim, C.-H. pag.7/8-15 Yong, L. pag.7/8-23 Yu, H. pag.11/12-67 Zanotto, F. pag.3-40; pag.4-5 Zbigniew, S. pag.5-11 Zhang, Y. pag.11/12-67 Zhao, X. pag.11/12-67 Zhong, Y. pag.3-5 Zucchelli, G. pag.6-48 Zucchi, F. pag.3-40 Zuliani, D. pag.7/8-58

La Metallurgia Italiana - n. 11/12 2018

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2019

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Heat Treatment & Surface Engineering for Automotive

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