MASTER'S THESIS

Thermodynamic Study of Trace Elementsin the Blast Furnace and Basic Oxygen

Furnace

Anton Andersson2014

Master of Science in Engineering TechnologySustainable Process Engineering

Luleå University of TechnologyDepartment of Civil, Environmental and Natural Resources Engineering

Acknowledgement This thesis was carried out at Swerea MEFOS and it was the final task before achieving the Master of Science degree in Sustainable Process Engineering. Moreover it serves as a door opener towards future challenges.

I would like to thank my supervisor Mats Brämming at Swerea MEFOS for his input and assistance throughout the work. Also, my supervisor Associate Professor Caisa Samuelsson and examiner Professor Bo Björkman at Luleå University of Technology for their valuable input. Furthermore, a thank you to the engineers at SSAB Luleå for providing process data and other information, Linda Bergman, Magnus Heintz and Anita Wedholm.

Luleå, August 2014

Anton Andersson

Abstract The iron ore based steel producers in Sweden and Finland operates mainly on pellets produced by LKAB. The introduction of new mining sites is believed to influence the future pellet chemistry. Furthermore, environmental and economic factors act as a driving force towards increasing the material- and energy efficiency by increasing the recirculation of in-plant by-products. All together this amounts to a realized change in feed chemistry. Therefore, it is of interest to study how trace elements behave in the process to be able to follow the material flow of trace elements within the integrated steel plant.

In this thesis it was attempted to describe the distribution of trace elements between metal, slag and gas phase in the blast furnace and basic oxygen furnace (BOF) process using thermodynamic equilibrium calculations. The work was focused on developing an approach to calculate the distribution of zinc in the blast furnace and chromium in the BOF. The same approach was then utilized for lead in the blast furnace and cobalt in the BOF to determine if it was applicable on other trace elements as well.

The blast furnace calculations were divided into three different scenarios representing different parts of the furnace; namely, the hearth, thermal reserve zone and the section above the thermal reserve zone. The results showed that for elements having a cyclical behavior in the furnace, such as zinc and lead, the assumed recirculation rate is directly decisive for the calculated output through the tap hole. And, that more data is needed to confidently estimate a probable recirculation rate that fits the calculations. Furthermore, it was shown that, from a thermodynamic standpoint, no lead or zinc leaves the blast furnace through the top. To describe the output of these elements through the off gas it was argued that a thorough study of the connection between furnace operating parameters and the dust and sludge amount and zinc and lead contents of the dust and sludge is required.

The BOF calculations were executed by adding the oxygen in increments to an open system, allowing the gas to leave between each calculation step. The calculations were carried out for and compared to results of a pilot plant scale converter and an industrial scale converter. From the results it was concluded that the distribution of chromium could be described for the pilot plant scale converter although the comparison of the overall composition of the slag and crude steel was not satisfactory. Furthermore, the distribution of chromium for the industrial scale converter could not be described using the method at hand. It was argued that the failure to describe the outcome resided in the fact that thermodynamic calculations were employed on a process where kinetics is known to play an important part. Cobalt could be described using the method. However, a simple mass balance with the assumption that essentially all cobalt reports to the crude steel phase would give the same results.

Symbols and Nomenclature The symbols used in this report are explained for below.

𝑘𝑔 𝑡𝐻𝑀⁄ = kg of an element per ton of hot metal, used to describe the load in the blast furnace.

𝑡𝐶𝑆 = ton crude steel

(%𝑀𝑒) = Denotes the weight percentage of component Me in the slag phase.

[%𝑀𝑒] = Denotes the weight percentage of component Me in the molten iron or steel.

𝐵𝑎𝑠𝑖𝑐𝑖𝑡𝑦 (𝐵2) = Ratio between mass percentage of CaO and SiO2 in the slag phase.

𝛾𝑥𝑜 = The Henrian activity coefficient of component x.

𝑋𝑦 = The mol fraction of component y.

𝑇 = The temperature given in Kelvin

𝑌𝑖𝑒𝑙𝑑 (%) =𝑡𝑜𝑡𝑎𝑙 𝑘𝑔 𝑜𝑓 𝐸𝑙𝑒𝑚𝑒𝑛𝑡 𝑖𝑛 𝑐𝑟𝑢𝑑𝑒 𝑠𝑡𝑒𝑒𝑙

𝑡𝑜𝑡𝑎𝑙 𝑘𝑔 𝑜𝑓 𝐸𝑙𝑒𝑚𝑒𝑛𝑡 𝑐ℎ𝑎𝑟𝑔𝑒𝑑 𝑡𝑜 𝑡ℎ𝑒 𝐵𝑂𝐹∙ 100

Table of Contents 1 Introduction ............................................................................................................. 1

1.1 Background ....................................................................................................... 1

1.2 Purpose, Aim and Scope ................................................................................... 2

2 Literature Survey ..................................................................................................... 3

2.1 Blast Furnace Ironmaking ................................................................................. 3

2.2 Basic Oxygen Furnace Process ......................................................................... 7

2.3 Modeling of Metallurgical Processes .............................................................. 10

3 Methods and Datasets ............................................................................................ 14

3.1 Thermodynamic Data ...................................................................................... 14

3.2 Blast Furnace Calculations .............................................................................. 15

3.3 Basic Oxygen Furnace Calculations ............................................................... 19

4 Results and Discussion .......................................................................................... 23

4.1 Results of the Blast Furnace Calculations ....................................................... 23

4.2 Basic Oxygen Furnace..................................................................................... 33

5 Discussion on the Thermodynamic Data used ...................................................... 40

6 Concluding Discussion .......................................................................................... 41

6.1 Blast Furnace ................................................................................................... 41

6.2 Basic Oxygen Furnace..................................................................................... 42

7 Conclusions ........................................................................................................... 42

8 Further Studies ....................................................................................................... 43

9 References ............................................................................................................. 44

10 Appendix 1 – Ingoing Values to the Blast Furnace Calculations ....................... 48

10.1 Hearth Equilibrium Calculations ................................................................. 48

10.2 Thermal Reserve Zone Calculations ............................................................ 50

10.3 Above the Thermal Reserve Zone and Below the Throat ........................... 50

11 Appendix 2 – Phase Diagrams ............................................................................ 52

12 Appendix 3 – Results from Calculations on the 2006 Dataset ........................... 54

13 Appendix 4 – Results from Basic Oxygen Furnace Calculations ...................... 55

13.1 Imphos Dataset ............................................................................................. 55

13.2 SSAB Dataset ............................................................................................... 57

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1 Introduction The background, purpose and aim of this Master of Science thesis are described below.

1.1 Background The reason why this thesis is of interest for the industry is briefly described below.

In ore based steelmaking, the blast furnace is the dominating process equipment used for reducing the iron ore to iron. The process operates on coke and iron ore agglomerated as either pellet or sinter. In Sweden and Finland, the iron ore agglomerates used are mainly the LKAB pellets. There are plans for introducing new mining sites which are expected to affect the future chemical composition of the pellet produced.

Furthermore, high primary raw material prices and environmental legislation drive the integrated steel plant to operate at higher material- and energy efficiency. The recirculation of in-plant by-products is a possibility when striving for higher material efficiency and thereby lowering the need of virgin material. The recirculation also reduces the amount of material being landfilled. The by-products subjected for recirculation are internal metal scrap, slags, dusts and sludge from different parts of the process. The slags are utilized to a high extent both as construction material [1] and as slag formers in the process [2]. Most of the iron rich dust and sludge is generally recycled to the sinter in the sinter making process [3]. When operating pellets, some sludge fractions and dusts can be recirculated to the blast furnace through cold bonded agglomeration of briquettes [2].

The dust and sludge from the blast furnace contains high amounts of iron and carbon [3]. The sludge from the basic oxygen furnace (BOF) process contains high amounts of iron [3]. In addition to this, trace elements (with respect to iron and steel) such as zinc, alkalis, lead, antimony, tin and cobalt are present in smaller amounts. Some of these elements such as zinc and alkalis are known to be harmful for the blast furnace process. The knowledge regarding the behavior of the other elements is limited and it is of great interest to investigate how these behave in the process.

At Process Integration at Swerea MEFOS mathematical based models of material flow within the integrated steel plant have been developed. As the demand for an increased recirculation of by-products and a possible change in pellet chemistry may change the chemistry of the feed there is a need of studying the effects of this on the material flow. By thermodynamically studying how trace elements are distributed between the metal, slag and gas phase some insight can be provided to this matter.

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1.2 Purpose, Aim and Scope The purpose and aim of the project is described below. Also, the scope is presented.

The purpose of this project is to study if calculations using thermodynamic equilibrium can be applied to describe the distribution of trace elements between metal, slag and gas phase in the blast furnace and BOF.

The aim is to calculate how zinc is distributed between the hot metal, slag and gas phase in the blast furnace and how chromium is distributed between the crude steel, slag and gas phase in the BOF process. These elements were chosen as there are available thermodynamic data describing the elements and data to use as comparison with the calculations. Furthermore, the project also aims towards using the same way of calculating to investigate if the approach can be applied for the distribution of lead in the blast furnace and cobalt in the BOF.

The thesis is limited to describe zinc in the blast furnace and chromium in the basic oxygen furnace as well as applying the developed calculation approaches to lead in the blast furnace and cobalt in the BOF.

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2 Literature Survey A brief theory regarding iron ore based steelmaking mainly focusing on the blast furnace and basic oxygen furnace process is described below. Also, published research of models describing the blast furnace and basic oxygen furnace are presented.

2.1 Blast Furnace Ironmaking The production of pig iron from iron ore is described briefly below.

An illustration of the iron blast furnace is presented in Figure 1 below. The raw material used in the process is coke, iron ore agglomerates of pellets or sinter and slag formers. [4] Other materials that can be used are cold bonded briquettes of residues, basic oxygen furnace slag [2] and scrap. Alternating layers of coke, iron burden and slag formers are charged at the top. The material descends due to gravity when the coke is continuously burned at the tuyere level and the slag and pig iron is tapped from the hearth.

In the tuyeres hot blast, i.e. preheated air, is introduced together with pulverized coal. The blast adds to the thermal balance of the blast furnace. Also, the reaction between the blast and the pulverized coal and the coke generates carbon monoxide. As the carbon monoxide ascends throughout the furnace the iron ore consisting mainly of hematite, Fe2O3, is reduced. [4]

When the iron oxide is reduced to metallic iron throughout the descent in the furnace the gangue material in the iron ore is reporting to the slag phase together with ash from the combustion of coke and pulverized coal. The iron is softened and melted in the cohesive zone and trickles together with the slag through the coke layers down to the hearth. Due to the density difference between the slag phase and hot metal (HM) the slag floats on the HM. The two phases can thus be separated when tapping the furnace. [4]

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Figure 1: Cross section of the blast furnace. Reactions occurring at different heights are presented. [5]

2.1.1 Reactions and Temperature Profile The ideal temperature profile and reduction scheme in the blast furnace is described in this section.

The reduction of hematite to metallic iron occurs in different stages with different requirements on the reducing gas. To describe this, the carbon monoxide (CO) utilization factor, %ηCO, is introduced as for Equation 1 below [4]:

%𝜂𝐶𝑂 = 100% ∙ % 𝐶𝑂2% 𝐶𝑂+% 𝐶𝑂2

(1)

Table 1 below presents the different utilization factors and CO/CO2 ratios at equilibrium for the reduction of the iron oxides at 900 °C. The reduction reactions given in the table are called the indirect reduction of the iron oxides.

Table 1: CO-utilization factors and CO/CO2 ratios at equilibrium for the reduction of iron oxides at 900 °C. [4]

Reaction CO/CO2 %ηCO 3Fe2O3 + CO = 2Fe3O4 + CO2 0 Ca. 100 Fe3O4 + CO = 3FeO + CO2 0.25 80 FeO + CO = Fe + CO2 2.3 30

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From the table it is seen that at 900 °C the required ratio to successfully reduce wüstite, FeO, to metallic iron is 2.3 and the equilibrium is as shown in Reaction 1 below:

𝐹𝑒𝑂 + 3.3 𝐶𝑂 = 𝐹𝑒 + 𝐶𝑂2 + 2.3 𝐶𝑂 (1)

The counter-current outline of the process enables a gas rich in CO to come in contact with the wüstite. The gas is then successively depleted in CO while ascending in the furnace reducing the iron oxides.

An ideal temperature profile is depicted in Figure 2 below. The reactions occurring are also given.

Figure 2: Ideal temperature profile for the blast furnace represented as height above tuyere level. [4]

From the above it is clear that the conditions in the blast furnace with respect to temperature and chemical environment are changing with respect to the position in the vertical direction of the furnace. This is of great importance to consider when estimating the distribution of elements between metal, slag and gas by equilibrium calculations.

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2.1.2 Behavior of Zinc in the Blast Furnace The cyclical behavior of zinc in the blast furnace is described below. Also, thermodynamic studies of zinc in systems of iron, carbon and zinc are accounted for.

2.1.2.1 Zinc Input, Output and Cyclical Behavior Zinc enters the blast furnace in small quantities through the iron ore as oxides, ferrite, silicates or sulfide [6]. It may also enter the process through the coke, recirculated BOF slag [7] and cold bonded briquettes of by-products [8]. The zinc inputs vary between 0.05 and 2.5 kg/tHM [9] noting that some European producers report inputs below 0.16 kg/tHM [10] [11] [12] [13]. The zinc compounds are reduced to metallic zinc vapor by the CO rich gas in the lower regions of the blast furnace where the temperature exceeds 1000 °C. The zinc vapor follows the ascending gas and is reoxidized to zinc oxide in the cooler parts of the furnace. The condensation occurs at temperatures below 520-580 °C [4] and the fine particles are either deposited on the lining and the burden material or exit the blast furnace through the off gas [6]. The deposition of zinc on the lining may disrupt the bricks due to the volume expansion of the transformation from gaseous to solid state. Other problems are scaffold formation which may disturb the burden descent [4]. The zinc deposited on the burden travels down to the lower region where it is reduced and volatilized again, forming a cyclical behavior [6]. The circulation of zinc within the blast furnace is negative with respect to the consumption of reducing agents [7]. The circulating load of zinc is the highest in the temperature region of 800-1200 °C. Samples have shown that zinc concentrations are ten times higher in the shaft than in the charged burden [9].

The zinc output through the off gas is increased by a large difference in the temperature of the top gas and burden material; i.e. a high top gas temperature and a low burden material temperature [9]. Also, a strong central gas flow is favorable for zinc removal at the top. The zinc content in the off gas was studied with laser induced breakdown spectroscopy in [9] showing a periodical change in zinc output related to the charging mechanism of alternating layers of coke and ferrous burden. It was also shown that a decrease in blast rate was correlated to decrease in zinc output through the off gas.

In addition to the off gas, the zinc may leave the blast furnace through the hot metal and slag phase. The zinc removed by tapping is increased with decreasing flame temperature as well as decreasing silicon and manganese content of the hot metal. However, most of the zinc is assumed to evaporate during the tapping, reporting to the cast house dust [9].

Material balance over the blast furnace has shown that 15-27 % of the zinc exits through the flue dust, 45-70 % through the sludge obtained when wet cleaning the off

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gas, 5-10 % through the hot metal, 5 % through the slag. In addition to this 5-10 % is deposited on the refractories inside the furnace [6].

2.1.2.2 Thermodynamic Studies of Zinc The research available on thermodynamic interpretation of the iron-carbon-zinc system is focused mainly on the steelmaking processes due to the use of secondary raw materials such as galvanized steel scrap. However, the carbon contents studied are close to the levels seen in the blast furnace where after the results are applicable in this context as well.

The activity coefficient of zinc has been studied with two different methods; namely thermochemical equilibration of zinc in an iron-carbon alloy achieved by metal-metal equilibrium [14] or gas-metal equilibrium [15] [16]. All studies present a strong positive deviation from the Raoultian behavior. However, the results contradict each other when considering the behavior with changing temperature. The results presented by [14] suggest an increase in zinc activity coefficient with increasing temperature while the results of [15] suggest a decrease in the zinc activity coefficient. Also, the activity coefficient achieved in [14] is considerably larger than that of [15]. A possible explanation suggested by the authors of [14] was the difference in range of XZn in the studies.

2.1.3 Behavior of Lead in the Blast Furnace The available information regarding the behavior of lead in the blast furnace is not as thorough as that of zinc. A short description is presented below.

Lead is brought into the process through the ore [4] or limestone in varying amounts between 10 and 50 g/tHM [9]. The lead compounds are completely reduced in the upper part of the furnace [4]. Since lead is insoluble in iron and has higher density than both the hot metal and slag it flows down and accumulates in the hearth. Although metallic lead has a low vapor pressure, some of it may vaporize and condense in the upper part of the furnace on the burden or lining material [4] [9]. It may also leave with the top gas reporting to the dust or sludge fraction [9]. Thus, lead may show a cyclical behavior as that of zinc.

2.2 Basic Oxygen Furnace Process The production of crude steel in the BOF is described below.

The hot metal from the blast furnace may be pre-treated before entering the basic oxygen furnace to allow for optimal operation when producing high quality steel. The pre-treatment can be designed to remove one or more of the elements silicon, phosphorous and sulfur. [17]

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The main purpose of the BOF process is to oxidize the carbon in the steel. The hot metal produced in the blast furnace has carbon content in the range of 4-5 % [17] [18] which is reduced to the desired level, usually below 0.1 %. Other important aspects are control of sulfur and phosphorous as well as reaching the desired temperature of the melt. [17]

2.2.1 Basic Oxygen Furnace Process Outline Below follows a brief description of the outline of the operation.

Figure 3 depicts the general operational steps of the top-blown BOF process. The distribution between the steel scrap and hot metal is in the range of 75-95 % hot metal and remainder scrap depending on local operational conditions [18]. Additional coolants may be added to utilize the heat evolved during the process. Examples are iron ore, pre-reduced pellets [17] or possibly cold bonded pellets of in-plant by-products [19]. The latter would favor the material efficiency of the integrated steel plant. The composition of the scrap is hard to measure and influence the presence of e.g. elements such as chromium.

Figure 3: General outline of BOF process. [17]

Oxygen is blown through a water cooled lance at supersonic speed into the metal bath under strict control of lance height above the bath level. The oxygen oxidizes iron, silicon, carbon, manganese and phosphorous [17]. Lime and dolomitic lime are added as fluxes. The lance height and flux additions are controlled to reach desired slag formation procedure [18]. The oxidized carbon leaves with the gaseous phase. Oxides of iron, manganese, phosphorous are transferred to the slag phase together with calcium sulfide, CaS [17]. Thus, undesired elements together with iron oxides are

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transferred to the slag phase during the decarburization to reach a steel of certain composition.

The steel is sampled to ensure that the desired composition and temperature of the steel have been reached. The slag and metal are tapped separately. [17] The kinetic energy of the tapping is utilized for mixing of alloys.

2.2.2 Change in Metal- and Slag Composition This section briefly describes the change in metal- and slag composition as well as temperature in the BOF process.

The change in chemical composition during the blowing period is illustrated in Figure 4 below. As the oxidation continues the temperature rises from about 1340 °C to the desired temperature in the range of 1650 °C [18].

Figure 4: Change in chemical composition of the metal phase during the BOF process. [18]

It is clear from Figure 4 that the prerequisites for thermodynamic calculations of trace elements change during the process. There is a difference in both temperature and oxygen activity of the melt during the blow. Also, presence of other elements may alter the activity of the trace elements and thereby the distribution. It can also be noted the difference as compared to the blast furnace which is carried out at reducing conditions and at lower temperatures.

Figure 5 below illustrates the change in slag composition during course of the process. During the oxygen blowing, the slag varies in composition. Some of these effects cannot be accounted for in a thermodynamic calculation. This is true for e.g. the CaO dissolution behavior and FeO formation as these phenomena are not equilibrium controlled.

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Figure 5: Change in slag composition during the oxygen blowing. [18]

2.2.3 The Oxidation of Chromium in Steelmaking This section presents a brief description of the behavior of chromium in steelmaking.

Chromium is dissolved as divalent and trivalent chromium in the slag phase. Increasing temperature, decreasing oxygen potential and decreasing slag basicity results in an increase of the ratio between Cr2+ and Cr3+. Since the BOF process is operated under high slag basicity and high oxygen potential Cr3+ predominates in the slag phase [18].

The chromium distribution ratio, (%Cr)/[%Cr], has been shown to be proportionally increasing with the FeO content of the slag. This is based on data from measurements of slag and metal phases from tapping of the electric arc furnace, laboratory experiments and the outdated open hearth furnace [18].

2.3 Modeling of Metallurgical Processes This section covers research published in the field of modeling of metallurgical processes such as the blast furnace and the basic oxygen furnace as well as examples from other metallurgical industries.

Pyrometallurgical processes are operated at high temperatures with, in general, high reaction rates. On this basis, the assumption of chemical equilibrium can be utilized to simulate the outcome of the operation. The equilibrium calculations can be performed by utilizing software which calculates the quantities of all species in the different phases by a Gibbs energy minimization routine [20] [21].

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To be able to describe processes where the outcome is deviating from equilibrium, mass and heat transfer can be introduced to the model [22] [23] [24] [25]. In these non-equilibrium models the assumption of high reaction rates are assumed to be valid locally. The reactor is therefore divided into several sections. In each section, chemical equilibrium is assumed to be reached and the distribution of the species in the different phases is calculated through a Gibbs energy minimization routine. The material and heat transfer between the sections are defined in accordance with the process conditions.

2.3.1 Modeling of the Blast Furnace The complexity of the blast furnace process is difficult to model even with all the knowledge and experience that have been acquired throughout the years. Some different approaches and simulations are presented below.

As summarized by [24] there are multiple ways to approach the problem of modeling the blast furnace. Mathematical models, models based on data mining, heat and mass balance models and thermodynamic models are some categories that can be distinguished. Each approach has its pros and cons. In this section, focus will be paid towards models with underlying thermodynamic calculations.

A thermodynamic two-step model of the blast furnace was developed [25] to quantitatively describe the behavior of alkalis in the process. The blast furnace was divided into two sections; namely, a hearth reactor and a gas condenser representing the shaft. The mass and heat flow of the model is illustrated in Figure 6 below. Equilibrium was assumed for both reactors and the calculations were proceeded until the alkalis satisfied a defined mass balance criteria between the input and output. The calculations were made for different scenarios of alkali load, basicity and hearth temperature. The results were consistent with plant observations, however, noting that absolute values provided by the model should not be taken as exact. A suggested improvement for the model was to divide the gas condenser into several steps to more accurately describe the alkali behavior.

Figure 6: Illustration of the mass and heat flow of the two-stage model. [25]

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The division of the blast furnace shaft into more sections was realized in the five-unit model presented by [24]. The blast furnace was divided into three equilibrium reactors and two heat exchangers as depicted in Figure 7. The equilibrium reactor, R2, represents the thermal reserve zone of the blast furnace. An interesting assumption made was complete reduction to metallic iron before entering the equilibrium reactor, R1, representing the hearth. In the hearth reactor the final slag and hot metal composition was calculated. The formation of a slag phase was exclusively considered in this reactor, i.e. not in R2. The last reactor, R3, represents the fast heat exchange, in the upper part of the blast furnace, between the burden material and gas phase as seen in Figure 2. The purpose of this reactor was to calculate the final equilibrium composition of the off gas.

The comparison to industrial data was limited by the analysis of the iron ore in the dataset. However, accuracy of the predictions was argued to be good despite slight overestimation of the carbon and silicon content of the hot metal.

Figure 7: The schematics of the five-unit model. R1, R2 and R3 are equilibrium reactors and C1 and H1 are heat exchangers. Explanations of the different streams are found elsewhere. [24]

2.3.2 Modeling of the Basic Oxygen Furnace Process Research published in the field of simulation of the BOF process is generally focused on the decarburization and dephosphorization. Some models are explained for below.

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In BOF steelmaking the carbon and oxygen content of the melt differs from equilibrium [26]. From this standpoint, most of the published research is incorporating measures to account for deviations from equilibrium.

A top-blown basic oxygen furnace was modeled [22] using the non-equilibrium approach. The converter was divided into four steady-state reactors. In each reactor, chemical equilibrium was assumed and calculated by minimizing the Gibbs energy. The reactors represented distinct parts of the converter with correlation to the outline of the process; a hot-spot-, metal-slag-, metal bath reactor and a slag mixer. The model was designed to simulate the outcome for stepwise addition of oxygen to the hot-spot reactor. The model showed good correspondence to published data with regards to the decarburization reaction and silicon content of the liquid metal.

The decarburization and dephosphorization in the BOF process was simulated with a kinetic approach by [27]. The model developed was based on a previous model for hot metal dephosphorization. The reaction kinetics was described by the double-layer theory, explained for in [28], allowing for mass and heat transfer in the model. The model outline is out of the scope of this thesis. However, it is noted that the model accurately simulated the slag FeO content related to the metal carbon content. Also, the simulation results using this model for phosphorous, silicon and manganese were similar to the compared dataset from the industry.

A model describing the BOF composed of a reaction model and a model for material melting and dissolution was developed by [29]. The reaction model was designed to consider both thermodynamics and kinetics. It was based on the assumption that only iron is oxidized by the top blown oxygen. The other elements were then subjected to a coupled oxidation-reduction reaction with the iron oxide. The melting of added scrap was described by two different mechanisms; namely, diffusive scrap melting and forced melting. The latter, where the temperature is above the melting temperature of the material, was also used for the other additions such as ore and FeSi. A validation of the model was presented in [30] comparing simulation results with the outcome from both a 170-ton and 330-ton converter. The calculated metal and slag composition agreed well with the actual. Also, the behavior of the slag and metal during the blow was similar to that presented in section 2.2.2 above.

To model the process phenomena in a top-blown BOF converter, computational fluid dynamics software was coupled with thermodynamic databases [31]. To construct a dynamic model using this approach is out of the scope of this work and such a model is found elsewhere [31].

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2.3.3 Modeling of Other Pyrometallurgical Processes Models including thermodynamic calculations have been proven to be successful in predicting the outcome of other metallurgical processes as well. Some examples are given below.

The process of argon oxygen decarburization (AOD) in stainless steelmaking was simulated by [20] using equilibrium based calculations. The different oxidation stages with altering argon oxygen quotients were simulated individually using the outcome of the previous stage as input in the next stage. For each stage the oxygen and argon gas mix was provided stepwise, allowing equilibrium to be reached after each addition. The model was found to predict the trends of the process well. The accuracy was argued to be lowered due to lack of thermodynamic data, the assumption of constant temperature and from not considering known kinetic factors of the AOD process.

An equilibrium based model of a flash converter was developed and evaluated in [21]. The process was simulated by introducing an initial amount of oxygen followed by stepwise addition of oxygen to calculate the distribution of the simulated elements between the phases as the process proceeded. The model predicted the outcome well, suggesting that the process mainly depends on thermodynamic properties.

Modeling of a Pierce Smith converter was performed in [23]. A dynamic, non-equilibrium, model was developed. The predicted values were compared to plant data and equilibrium calculations. The non-equilibrium approach showed better agreement with plant data than only equilibrium calculations. The model was developed by dividing the converter into horizontally aligned sections connected by heat and mass flows. In each section chemical equilibrium was assumed to be reached.

3 Methods and Datasets The method employed for the calculations are presented below together with the datasets used as comparison and the thermodynamic data utilized.

The calculations were performed with the thermochemical software FactSage 6.1. To calculate the equilibrium composition the module Equilib was used. This function utilizes the Gibbs minimization algorithm and thermochemical functions of ChemSage [32].

3.1 Thermodynamic Data The thermodynamic data used in the calculations are presented below.

The thermodynamic data used in the calculations were taken from the FactSage databases. For solid oxides, solid oxide solutions and slag phases the FToxide database was used. The solution Aslag-liq was used to describe the slag phase in both the BOF

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and blast furnace. For the liquid iron- and steel phase, Fe-LQ from FTmisc was used. Gaseous species were taken from the Fact53 database.

The Henrian activity coefficient for zinc in a liquid iron-carbon-zinc system is given in Equation 2 below. It was taken from the function derived from the experiments conducted in [15].

ln 𝛾𝑍𝑛𝑜 = 3490𝑇− 0.142 + �2136

𝑇+ 2.86� 𝑋𝐶 − �3915

𝑇+ 1.83�𝑋𝐶2(2)

The activity coefficient was recalculated for the carbon content reported for hot metal in the blast furnace operation to represent the form shown in Equation 3 below.

log 𝛾 = 𝐴 + 𝐵𝑇 (3)

Where A and B are constants. The activity coefficient function was merged with the Fe-LQ solution.

3.2 Blast Furnace Calculations The approach used for the blast furnace calculations is presented below together with a short description of the datasets used to compare the calculations with.

In this thesis, no software enabling mass and heat transfer between different reactors was used. From section 2.1 it is clear that the blast furnace cannot be covered in one calculation step where thermodynamic equilibrium is assumed to be valid. To account for the difference in chemical environment and temperature the calculations were performed for a different set of scenarios representing the various parts of the furnace; namely, the hearth, the thermal reserve zone and the part above the thermal reserve zone. The methodologies for these are accounted for in the subheadings below.

The datasets used for comparison were two different zinc balances for blast furnace no. 3 at SSAB Luleå. The first balance was over a long term period provided for the entire operation of 2012; this dataset is referred to as the 2012 dataset. The second was a period over three days during which a full-scale trial with lowered top pressure was performed [33]; this dataset is referred to as the 2006 dataset. Operational data from SSAB Luleå was used as the foundation in assumptions regarding temperatures, pressures and %ηCO.

3.2.1 Hearth Equilibrium Calculations The calculations performed for the lower part of the blast furnace were designed to treat the hearth as an equilibrium reactor. The assumptions regarding the calculations are explained for in detail below.

The hearth equilibrium calculations were performed for both the 2006 and 2012 dataset. The elements considered in the calculations were chosen to represent the

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major constituents of the gas, slag and hot metal phase. These elements are Fe, C, Si, Mn, Ca, Mg, Al, O and N together with the trace element Zn. Elements important in other aspects were left out, e.g. hydrogen, sulfur and phosphorous.

It was assumed that all iron ore is completely reduced to iron before entering the hearth. The material inputs, temperature and pressure used in the calculations are accounted for in appendix 1. The inputs were calculated from the material balances for 2012 and 2006 provided by SSAB. In the data from SSAB, for each year the average composition for the different charged raw materials are given. Also, the input amount is given as an average for each month and for the entire year. For the dataset of 2012 the average composition and input for the entire year was used. For 2006 the average composition for the entire year was used together with the average input for July, i.e. the month specific for the full scale trials. For convenience, all calculations were performed on a per ton hot metal basis.

No solid oxide phases were allowed to exist. This was based on that solid particles in the slag phase are unwanted and avoided in real operation due to the increase in slag viscosity. Furthermore, the slag was not allowed to form before the hearth, i.e. the oxides were added as oxides to the calculations and not as a single liquid slag phase. The limestone charged was assumed to be completely burned during its descent down the furnace.

Solid carbon in the form of graphite was allowed to exist to represent the dead man in the hearth. Graphite was chosen instead of coke since there were no data describing the solubility of coke in liquid iron in the databases used. No distinction between coke carbon and carbon from the pulverized coal injection was made. The refractory was neglected. The hearth temperature was assumed to be the same as the tapping temperature. The pressure was assumed to be the same as the blast pressure and all the blast was allowed to be in equilibrium with the hot metal and slag.

Calculations in [25] showed that the circulating load of alkali increases linearly one to one with increasing alkali load. The same was assumed for zinc in these calculations although the behavior regarding zinc output is different. The circulating load of zinc was assumed to be ten times the input based on the information provided by [9].

3.2.1.1 Illustrating the Effect of Assumptions The effect of assumptions related to blast volume and behavior of zinc was illustrated with calculations based on the 2012 dataset. The reactants, temperature and pressure used in the calculations are presented in appendix 1. In one scenario the zinc was assumed to enter the hearth as condensed zinc oxide. The entire circulating load of zinc was therefore used as input in the calculations regardless of the amount of blast. In the other scenario zinc was assumed to be gaseous. Therefore, a decrease in blast amount resulted in decreased ingoing zinc in the calculations.

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To enable the use of different blast volumes for a fixed amount of ingoing solids the reaction between carbon and oxygen was calculated as a stoichiometric combustion to carbon monoxide. The gas phase used as input was therefore composed of CO(g) and N2(g). The carbon left was also introduced in the calculations to allow dissolution into the hot metal.

3.2.1.2 Trends in Zinc Output A useful application area of thermodynamic calculations is the information that can be received from changed conditions. Therefore, a different set of changes to the 2012 dataset was made to illustrate trends in the behavior of zinc. In [9] there is a summary from Russian blast furnaces which reports circulating loads up to and above 20 kg/tHM. Therefore, the effect of these high loads on the zinc output from the hearth was investigated. The articles referred to in [9] presenting the circulating loads for the Russian blast furnaces could not be accessed to validate how these loads were estimated. However, [33] refers to studies on frozen blast furnaces that have shown circulating loads up to 7.2 kg/tHM. Therefore, it was assumed that the high levels reported by the Russian producers are justified to use in the calculations. Furthermore, scenarios of changes in hearth temperature and pressure were calculated. The input, temperatures and pressures used in the calculations are accounted for in detail in appendix 1.

3.2.2 Thermal Reserve Zone Calculations The method used to calculate how zinc behaves in the thermal reserve zone is presented below.

The blast furnace at SSAB Luleå is capable of operating with a top pressure up to 1.5 atm over pressure. The span of pressures that the outgoing gas may be subjected to is thus between 1 and 2.5 atm. Normal operating conditions show blast pressures of about 3.5 atm and top pressures of approximately 2 atm. Since the furnace is designed to operate with increased top pressure it was assumed, in the calculations, that the pressure in the shaft may vary between 1.5 and 3.5 atm depending on the vertical position in the furnace and the top pressure.

The input for the calculations is presented in appendix 1; it is based on the 2012 dataset. The hematite charged to the blast furnace is assumed to be reduced to wüstite before entering the thermal reserve zone. The carbon left in the hearth equilibrium calculations is assumed to be available for direct reduction of the wüstite. Cementite, Fe3C, was assumed not to form and carbon formation was not allowed. As argued by [24], the slag formation in the blast furnace is dependent on iron ore chemical composition and local equilibrium conditions. To avoid this complexity a liquid slag phase was not allowed forming in this part of the furnace.

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The gas phase was imported from the hearth equilibrium calculations to represent the ascending gas in the process. Slag formers were added as oxides with the exception of calcium. Part of the calcium input was set as limestone while the calcium in the charged BOF slag and briquette was set as lime.

The retention time of the thermal reserve zone was assumed not to favor the kinetics of metal oxide solid solutions or spinel phases with zinc. Thus, the only forms of zinc allowed was Zn(g), Zn(s), Zn(l) and ZnO(s).

The calculations were performed for two different conditions with respect to temperature and pressure. The first condition was designed to represent operation with high top pressure and a probable temperature. The pressure was assumed to be 3 atm, allowing an assumed pressure drop of about 0.5 atm over the cohesive zone. The temperature was set as 900 °C which is within the temperature interval reported for the thermal reserve zone in the literature. The second calculation was designed to find out at what temperature the zinc gas is oxidized and condensed as zinc oxide. A pressure of 2.75 atm was chosen to represent a mean value over the thermal reserve zone.

3.2.3 Calculations for Above the Thermal Reserve Zone and below the Throat The method for the calculations representing the zone above the thermal reserve zone is described below.

Considering the approach presented above in 3.2.2 it is clear that a major limitation is the inability to allow changes in pressure and gas composition in the gas while ascending through the thermal reserve zone. This was enabled by not introducing the solids in these calculations. The calculations were built up as a gas cooling study with changing conditions in both pressure and %ηCO.

The temperature of the gas leaving the thermal reserve zone was set as 900 °C and was cooled down to 500 °C in the calculations. At gas temperatures below this, the interaction with the burden was assumed to be only heat transfer. Therefore, the %ηCO reached its final value of 54.3 % in the calculations, which is the actual value provided by SSAB. The %ηCO of the gas leaving the thermal reserve zone was calculated using the results from the hearth equilibrium calculations as base. The carbon left in that calculation was assumed to directly reduce wüstite to iron forming carbon monoxide. The carbon monoxide of the gas leaving the hearth was then assumed to react with the remainder of the wüstite to form iron and carbon dioxide. The resulting gas composition that was used as input in the calculations is presented in appendix 1. Since no solids are considered, the %ηCO value was changed manually. It was assumed to increase linearly to the its final value during the gas cooling, starting at the value calculated as described above.

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As previously mentioned, the blast pressure is about 3.5 atm in operation and a pressure drop of 0.5 atm was assumed over the cohesive zone. From the ideal temperature profile depicted in Figure 2 it can be seen that the thermal reserve zone is roughly estimated to 11 m high and the section above about 4.5 m high. A linear pressure drop was assumed above the cohesive zone which means that the pressure drops from 3 atm to the top pressure of 2 atm in 15.5 m. From this it was calculated to use a pressure decreasing from 2.3 to 2 atm in these calculations.

The calculations were also conducted for top pressures of 1.5 and 1 atm. The same assumption regarding linear pressure drop was utilized here. The blast pressure was assumed to decrease as much as the top pressure. The calculation scheme for all three cases is presented in appendix 1.

3.2.4 Lead Calculations The approach used for zinc was tried for lead as well. Additional assumptions and alternations to fit the expected behavior of lead are described below.

Due to the behavior of lead as described in section 2.1.3 it was assumed that the load of lead in the hearth is higher than the charged. Comparing the description of zinc and lead in the literature it was assumed that the load was 20 times the input of lead. The calculations for the hearth equilibrium were performed in the same way as for zinc. The input of lead was determined from analyses of trace elements in the pellets, limestone and BOF slag. The lead content of the charged briquette was assumed to be 100 ppm. Analyses of lead in hot metal were provided for different tappings. All analyses showed the same value for lead despite a difference in hot metal composition. Therefore, it was assumed that the lead output from the bottom rather depends on input and load in the hearth instead of hot metal composition. It was therefore decided to perform the lead calculations on the 2012 dataset.

A gas cooling study was performed for the gas phase output from the hearth equilibrium calculations. It was conducted in the same way as for zinc in section 3.2.3 above.

3.3 Basic Oxygen Furnace Calculations The assumptions and procedure employed when calculating the outcome of heats of the BOF is presented below.

3.3.1 Datasets for Comparison The datasets used for comparison are accounted for below.

Two datasets were used to compare with the BOF calculations. The first dataset was the Imphos trials [34], a project founded by the European Research fund for Coal and Steel (RFCS). The second was composed of nine heats from the BOF at SSAB Luleå.

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The Imphos trials were carried out in a six ton converter with metal and slag sampling before, during and after the blow. Chromium content in both hot metal, crude steel and final slag was provided. Out of the 23 available heats eleven were chosen to represent a wide span in slag FeO content. The exclusion of heats was made based upon poor chromium balances, comments in trial data and uncertainties in reported analyses.

The dataset from SSAB Luleå represents the full scale operation as SSAB process over 20 times more crude steel per heat as compared to the Imphos trials. The chromium content of hot metal and crude steel was provided in the dataset. Normally, the slag is not sampled which means that no chromium content of the slag phase was provided. In fact, the slag analyses provided for the heats were all calculated by the software used by SSAB. The heats were chosen to represent a span in calculated slag FeO content.

3.3.2 Calculation Basis The best results were obtained using two different approaches for the datasets where after these are described in separate headlines. Some general information true for both datasets is presented below.

The elements considered in the calculations were Fe, C, Si, Mn, Cr, Ca, Mg and O. The slag analyses found in [34] shows that vanadium reported as V2O5 may constitute up to six weight percent of the slag weight, i.e. a considerable amount. Vanadium was neglected in the calculations due to lack of thermodynamic data and possible effects of this oxide on the slag properties were lost.

In the BOF process, slag formers are added at carefully chosen times into the oxygen blowing to favor the slag formation. In the calculations all material input was introduced before the oxygen blowing started. Adding the slag formers at a certain time into the blowing period was considered to be unnecessary since the complex slag forming process is not thermodynamically controlled.

There are at least two different approaches available when considering the gas phase in the calculations. One is to add all the oxygen at once in one equilibrium calculation. Another is to add the oxygen stepwise and allow the gas phase to leave after each equilibrium calculation, i.e. an open system. As argued by [20], calculating in an open system will lower the partial pressure of carbon monoxide. This will not have any significant effect at higher carbon contents. At lower carbon contents a higher partial pressure of carbon monoxide will result in higher crude steel carbon content. Since the Imphos dataset consists of charges with carbon content down to 0.02 % it was preferred to use the open system approach. Also, from a process point of view it is more realistic to allow the gas phase to leave. Therefore, the open system approach was used.

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During the blow, the temperature steadily increases in the process due to the oxidation of the different elements. Using the open system approach, the software did not allow calculations of a final temperature adiabatically or under heat losses. Therefore, a fix temperature was chosen to apply during the entire blow. The pressure was assumed to be one atmosphere.

The material input for the calculations were calculated from the reported charged weights of hot metal, scrap, lime, dolomite, raw dolomite and iron ore. The composition of the scrap used in the Imphos trials was estimated as a general analysis for scrap of higher quality used at SSAB. Estimations of the composition of different scrap qualities charged to the SSAB process were provided to calculate the amount of ingoing material. These analyses are left out of the report.

Figure 19 and Figure 20 in appendix 2 illustrates that the slag phases may be saturated in e.g. lime and periclase. From the figures it was decided to allow solid CaO, MgO and Ca2SiO4 to exist. It was also assumed that a metal oxide solid solution could exist, based on solid solutions rich in MgO.

Chromium was presented as Cr2O3 in given slag analyses. In the calculation both CrO and Cr2O3 was allowed to form in the slag phase. However, after the thermodynamic calculation, the CrO was recalculated to Cr2O3 to ease the comparison with the actual values.

3.3.2.1 Assumptions and Method Specific for the Imphos Dataset For the most part, no MgO was charged in the Imphos trials. Instead, the input of MgO was calculated from dissolved refractory reporting to the outgoing slag.

The ingoing material together with the slag analysis was used to calculate a slag weight. The slag weight was taken as the average for two calculations; one based on silicon and one on calcium. In the dataset no crude steel weight was provided. This was estimated using the calculated slag weight and total charged iron together with the slag and crude steel analysis in a material balance. This means that no losses of iron to the gas phase was assumed. The values from the slag weight and crude steel weight calculations were then incorporated in a chromium balance calculation to estimate the difference between ingoing and outgoing amount of chromium. If the difference was too high the heat was not considered.

The oxygen input to the calculations was calculated based on the stoichiometric oxygen demand for carbon oxidation and the oxidation of elements to the slag. Comparing this to the actual injected amount of oxygen, the remaining oxygen was assumed to go to post combustion of carbon monoxide. This assumption means that carbon dioxide was not allowed to form in the equilibrium calculations. The oxygen was added in 100 steps in an open system as described earlier.

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The fix temperature used in the calculations was chosen as the end of blow temperature. For the heats with end temperatures well over 1700 °C the average of the two last temperature measurements of the blow were used instead of the end of blow temperature.

3.3.2.2 Assumptions and Method for the SSAB Dataset The approach used for the SSAB dataset differs from the Imphos in the way oxygen is added. The oxygen was added in increments of 48.541 kg O2 until the desired carbon content of the crude steel was reached. The amount of each addition was calculated from the assumption that 340 Nm3/min is added to the process. This was recalculated to kg of O2 per six seconds which results in 48.541 kg O2. Six seconds was chosen to limit the number of calculation steps while keeping the possibility to stop at the desired carbon content.

To avoid problems of no slag formation only CaO and MgO was allowed to be present as solids in the slag phase. The fixed temperature was chosen as the end of blow temperature. Carbon dioxide was not allowed to exist in the gas phase although the oxygen addition in this calculation method differed. This was based on the assumption that post combustion occur from leakage air and at high lance positions for the gas that have already left the open system.

A slag weight was calculated as described above in 3.3.2.1. The dataset provided an estimated steel weight but the chromium content of the slag was not given. A material balance of the heat was used to estimate the chromium slag content. All chromium was assumed to be present as Cr2O3.

The interaction with refractory was neglected in the calculations. Also, iron leaving as iron oxide with the gas phase was not considered.

3.3.2.3 Cobalt Calculations The approach used to calculate the distribution of chromium for the SSAB dataset was employed for cobalt as well. Three trace element analyses of hot metal were provided. The analyses showed cobalt contents of 0.013 %, 0.012 % and 0.012 %. Since the hot metal composition differed between the three analyses it was decided to use the average of the three as input to the BOF. A trace element analysis of scrap was also provided showing cobalt content of 0.016 %. From this it was assumed that all bought and internal scrap had a cobalt content of 0.016 %.

The calculations could not be performed for a specific heat with cobalt analyses for the hot metal, scrap and crude steel. Instead, the ingoing analyses as described above were used for three heats with trace element analysis of crude steel. The heats were chosen to represent a span in reported crude steel carbon content.

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The output of cobalt to the gas cleaning system was estimated from trace element analyses of the fine and coarse sludge. It was assumed that 6 kg/t CS dry coarse sludge and 13 kg/t CS dry fine sludge were formed in the process. Cobalt reporting to the slag phase was estimated from BOF slag analyses and calculated slag weights.

4 Results and Discussion The results are presented and discussed in this section of the report.

4.1 Results of the Blast Furnace Calculations The results for the blast furnace calculations are presented and discussed below. The sections are divided in the same way as the methods describing the calculations are.

4.1.1 Hearth Equilibrium Calculations The results for the 2012 dataset are presented in Table 2 through Table 4 below. From the standpoint that this thesis focuses on trace elements it can be argued that only Table 4 is of interest. However, it seems sensible to evaluate an overall performance of the calculations since faults in the prediction of major components may affect the distribution of the trace elements.

Table 2 presents the comparison of calculated hot metal composition and actual composition. It can be seen that the carbon content of the hot metal is overestimated. In the calculations, crystalline graphite was set to represent the dead man instead of amorphous coke which could explain the difference. However, the value is still considered to be satisfactorily close to the data. In the calculation of the activity coefficient function for zinc in hot metal the carbon content of the process data was used. To make sure that the thermodynamic calculations can be utilized without relying on values already achieved in the process, the zinc activity coefficient function was recalculated using the value for the overestimated carbon content. Using the new value in the calculations resulted in a decreased hot metal zinc content of 3 %, from 2.7 to 2.6 ppm.

The calculated values of silicon and manganese were both considered to be acceptable considering they were not the main goal of this thesis. It is unfortunate that the value provided for zinc by the data is below the detection limit. Regardless, it is comforting to see that the calculated value is below this limit as well.

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Table 2: Hot metal analysis from calculations and blast furnace data for the 2012 dataset.

Element Calculated Actual %C 5.11 4.67 %Fe 94.2 94.0 %Mn 0.37 0.34 %Si 0.36 0.43 %Zn 0.0003 <0.0005

Table 3 below presents the slag analysis, B2 basicity and slag rate for both the calculations and the data. Considering the simple approach employed, the calculated contents of the major components are satisfactorily close to the data. The calculated ZnO content is well below the ppm range. The result is realistic due to the high temperature, non-oxidizing conditions and zinc oxide activity coefficient in the slag phase. The blast furnace slag analysis of ZnO is excluded as the value provided in the balance was 82 ppm. Analyses of crushed blast furnace slag from 2009 to 2013 were provided by SSAB Merox. Out of 26 measurements, twelve were below the detection limit of 4 ppm. In addition to this, eleven was above the detection limit and below 20 ppm. No measurement was close to the level of 82 ppm. Furthermore, the slag sampling made 2006 [33], when the zinc rate was almost three times as high as 2012, all gave values below the detection limit of 4 ppm. Therefore, 82 ppm was considered to be unlikely to represent a mean value for the entire year.

Table 3: Slag analysis, B2 basicity and slag rate from calculations and blast furnace data for the 2012 dataset.

Parameter Calculated Actual %CaO 37.5 33.3 %SiO2 35.0 33.7 %MgO 14.8 15.1 %Al2O3 12.7 12.9 %MnO 0.07 0.37 %ZnO 0.0000 / B2 1.1 1.0 Slag rate kg/tHM 164 166

Table 4 below presents the material balance for zinc based on the thermodynamic calculations and the actual balance from the data for the year of 2012. The zinc content of the hot metal was set to half the detection limit provided by the data. The slag output of zinc in the blast furnace data was decided to be left out as argued before. The material balance was closed in the thermodynamic calculations, i.e. the input equals the output. The material balance from the blast furnace data suggests that zinc

25

accumulates in the process over the year. This is of no concern since incomplete balances are common when studying zinc.

The results indicate that the prediction of zinc in hot metal is in the right range, i.e. below the detection limit of the analysis. It is noted that cast house dust is not considered in the blast furnace data material balance.

Table 4: Zn balance from calculations and compared material balance.

Stream Calculated (g Zn/tHM) Actual (g Zn/tHM) Input Total input 100 100 Output Hot Metal 2.7 2.5 Slag 0.0 / Dust & Sludge 97.3 88.7 Total output 100 91.2

4.1.1.1 Dataset from 2006 Only the final zinc balance is provided for this dataset, Table 5. The complete hot metal- and slag composition are found in appendix 3. As explained for earlier the data is from three days, one of which was operated with lowered top pressure. Therefore, there are great variations in the output of zinc through the top. Furthermore, a balance over this short period does not give a good representation of the output through the top due to the behavior of zinc in the furnace. Also, the output of zinc did not exceed the input although the top pressure was lowered considerably. Thus it was assumed that the recirculation rate was not lowered during the three day period. It was therefore decided to exclude the off gas dust and sludge fraction from the balance to be able to use an average over three days instead of one day when comparing the calculated and actual zinc output in the blast furnace bottom.

The calculated zinc output via the hot metal is underestimated but still within a reasonable range of the data. The cast house dust contains vaporized zinc from the tapping. This means that the zinc content in the hot metal and slag phase is not at equilibrium after leaving the furnace. The lower pressure outside the furnace provides a driving force for the volatilization of zinc. Therefore, the prediction of the calculations is worse since the cast house dust should be accounted for in the calculated slag and hot metal output.

The output in the slag as given by the blast furnace data is based on half the detection limit of the analysis. From the arguments regarding ZnO in section 5 below together with the calculations it is concluded that, from a thermodynamic standpoint, the zinc

26

content in the slag is below ppm range. Higher zinc contents as reported in section 2.1.2.1 are an effect of other mechanism. It was tried to account for this by fixing the activity of ZnO in the slag. This approach generates problems regarding the material balance of zinc in FactSage and was therefore discarded.

Table 5: Comparison between zinc outputs from the 2006 dataset.

Stream Calculated (g Zn/tHM) Actual (g Zn/tHM) Input Total input 290 290 Output Hot Metal 8.4 12.2 Slag 0.0 0.4 Cast house dust / 2.9

4.1.1.2 Assumptions in the Hearth Calculations The calculations are directly dependent on the assumed recirculation of zinc in the furnace and not knowing the accuracy of this assumption is devastating for the confidence in the results. Both calculations presented above gave results in the range of the data, although one measurement was below the detection limit. To assure that the assumption made in these calculations is valid, more measurements are needed to compare with. It is likely that the circulating load is not constant but rather changes with zinc rate and operation.

Another important assumption having great effect on the calculation results is the amount of gas allowed to be in equilibrium with the melt and slag as this affect the partial pressure of the zinc gas. The effect of this was studied and the results are presented in Figure 8 below. If the mechanism of zinc dissolution in hot metal and slag is assumed to be from the gas phase then there is little influence. This is an effect of that the total zinc content decreases with decreasing gas amount. However, if a fixed amount of condensed ZnO is assumed to enter the hearth, then this input would be the same regardless of the amount of gas in equilibrium. From the figure it is clear that the latter has a significant effect on the calculation results.

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Figure 8: Comparison of the effect of blast volume on the zinc content in hot metal with two different assumptions regarding the behavior of zinc.

4.1.1.3 Trends in Zinc Output from Blast Furnace Hearth The result of increased circulating load of zinc is illustrated in Figure 9 below. The ppm zinc in hot metal increases linearly which is an effect of not using interaction parameter formalism in the description of zinc. The change in output through the slag is insignificant for the zinc balance.

Figure 9: Effect of change in circulating load on output of zinc in hot metal and slag.

The effect of changed temperature on the zinc content in hot metal and slag is depicted in Figure 10 below. The zinc content in the hot metal decreases with increasing temperature which is expected due to the volatility of this element.

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Figure 10: Effect of change in temperature on output through the tap hole.

Figure 11 below illustrates the effect of changed hearth pressure on the zinc content in hot metal and slag. The increase in dissolved zinc is expected if Le Chatelier’s principle is considered.

Figure 11: Effect of change in hearth pressure on zinc content of hot metal and slag.

From Figure 9 to Figure 11 above it is clear that the combined effect of circulating load, temperature and pressure have large influence on the zinc output through the tap hole. Although there is no data to compare with, the results presented in this section contribute with interesting information whether a change in a parameter pose significant or insignificant changes in zinc output from a thermodynamic standpoint.

The effect of changed B2 basicity on zinc content in the slag phase was calculated but it is left out of this report due to the below ppm level content.

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4.1.2 Thermal Reserve Zone In the calculations for the thermal reserve zone it was seen that, from a thermodynamic standpoint, zinc does not oxidize and precipitate. For a pressure of 2.75 atm zinc oxide starts to form at a temperature of 738 °C which is located above the thermal reserve zone. The results also show that careful consideration of allowed phases in the shaft needs to be taken as the thermodynamic composition may differ from the real.

4.1.3 Above the Thermal Reserve Zone and below the Throat The results from the gas cooling study conducted for the region above the thermal reserve zone are presented in Figure 12 below. The reported temperature interval where zinc condensation begins as described in section 2.1.2.1 does not agree with the results achieved in these calculations. The interval could be intended for other conditions regarding top pressure and %ηCO. In the provided data it is found that the mean top gas temperature of 2012 for blast furnace no. 3 at SSAB Luleå was 133 °C with monthly averages up to 150 °C. This means that thermodynamically, zinc does not leave the blast furnace in the top.

Figure 12: Gas cooling study above the thermal reserve zone. The zinc as gaseous zinc and solid zinc oxide is given as a percentage of the total amount of zinc. The units are presented in brackets to avoid confusion with solidus and gaseous notation.

It is well known that zinc exits the blast furnace via the top gas in both the sludge and dust fraction for furnaces operated with or without top pressure. It was therefore attempted to use the calculations presented above as a base when estimating the outgoing zinc. Considering the ideal temperature profile of the blast furnace as

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depicted by Figure 2 it can be seen that the gas temperature decrease rapidly just below the throat. From the figure it is estimated that the gas has a temperature of 650 °C at a level of 2.5 m below the throat. This temperature is chosen as a cut-off temperature where:

1. The already condensed zinc oxide travels down the furnace with the burden, and

2. The zinc gas condensing when ascending to the top travels with the gas phase out of the furnace.

By doing this, 91.7 % of the circulating load will travel down the furnace, continuing the zinc cycle. 8.3 % or 83 g/tHM would leave through the top. This is not far from the value of 88.7 g/tHM which was presented in Table 4.

The effect of changed top pressure on zinc condensation during the gas cooling is illustrated in Figure 13 below. The figure illustrates how lower top pressure is less favorable for zinc oxidation and condensation. In the full scale trials conducted in [33], a decrease in top over pressure from 95 kPa to 65 kPa resulted in an increased average zinc output of 100 g/tHM in the blast furnace top. In that case it corresponded to a 91 % increase. Choosing 650 °C as a cut-off temperature for top pressures of 2 atm and 1.5 atm in the figure below it was calculated that the lowered top pressure increased the zinc output by 31 %. The major part of the increased output when lowering the top pressure is consequently explained for by other factors than thermodynamics. Lowering the top pressure increases the gas flow through the furnace which results in an increased amount of dust and sludge. A change in either blast rate or top pressure can therefore be interpreted as a change of at which distance from the throat the condensed zinc oxide travels with the gas phase out of the furnace. One way to combine the thermodynamics with furnace parameters is to develop a way to describe the cut-off temperature as a function of e.g. blast rate and top pressure. The data available for comparison is scarce (two operational points to compare with) where after such a function could not be designed in this thesis.

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Figure 13: Effect of changed top pressure on oxidation and condensation of gaseous zinc to zinc oxide.

Although it would be interesting to try to describe the zinc output in the fashion presented above, it is hard not to consider it as wishful thinking. The choice of cut-off temperature has a significant effect on the output, even at small changes. Also, the ideal temperature profile of Figure 2 is very much simplified. The temperature may differ in the radial direction depending on operation, i.e. if the furnace is operated with a chimney with strong central gas flow or possibly some unwanted situation which affects the gas flow.

The approach described can also be seen as a case where thermodynamics is used to endorse a method to calculate the output of zinc when, in reality, the assumptions and adjustments made have overtaken all effects of the thermodynamic calculations. It is believed that the calculations presented here are better suited to describe trends and possibly hint as to how much different effects can be explained by a change in thermodynamic prerequisites, e.g. lowered pressure. If, however, the calculation method is to be developed it requires more data on zinc output coupled with blast rate, top pressure, top gas temperature and possibly burden distribution measurements.

4.1.4 Lead Calculations The results from the lead calculations are presented and discussed below.

Table 6 below presents the results from the hearth equilibrium calculations. The output through the hot metal and slag is directly dependent on the assumed lead load in the

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Top P = 2 atm Top P = 1.5 atm Top P = 1 atm

32

hearth. If a higher load was to be assumed the calculations would represent the actual value better. To achieve the correct hot metal analysis for one point of comparison is easy when considering a circulating element. Additional data would be required to decide whether an increased recirculation rate can be employed to describe the output through the hot metal. The calculations suggest that lead does not report to the slag phase. An actual value for the slag lead content was not provided. The calculated dust and sludge amount was obtained by closing the material balance over the furnace. Comparing to the actual amount leaving the blast furnace top, it is clear that the estimated input is off or that the data provided for the sludge and dust is bad.

Table 6: Comparison of calculated and actual lead output. All numbers are given in g/tHM

Parameter Calculated Actual Hot metal 0.1 0.4 Slag 0.0 / Dust and sludge 3.0 12.3

The result of the gas cooling study is presented in Figure 14. As for zinc, no lead leaves the furnace from a thermodynamic standpoint. The figure fails to illustrate how the condensation of lead is independent of the %ηCO-value. This independency is explained by the fact that lead does not form an oxide.

Figure 14: Gas cooling study of lead. The lead is presented as % of total lead being condensed.

2,00

2,05

2,10

2,15

2,20

2,25

2,30

0,0

10,0

20,0

30,0

40,0

50,0

60,0

70,0

80,0

90,0

100,0

500 550 600 650 700 750 800 850 900

Pres

sure

(atm

)

cond

ense

d Pb

, Eta

CO (%

)

Temperature (°C)

%ηCO %of total Pb condensed Pressure

33

As previously explained, sulfur was not incorporated in the calculations. Although lead has high affinity for sulfur it is argued that it would not form lead sulfur compounds at equilibrium since sulfur is found in the solid matrices of the burden. Furthermore, such compounds would not explain the output through dust and sludge due to their properties in comparison to that of lead.

4.2 Basic Oxygen Furnace The results for the BOF calculations are presented below.

4.2.1 Imphos Calculations This section compares the results of the calculations and actual outcome of the Imphos converter trials.

A comparison of the actual and calculated analyses for three heats from the Imphos trials is presented in Table 7 below. Additional heats are presented in appendix 4. As for the blast furnace calculations, it is believed that an overall performance of the calculations should be evaluated. The composition of the calculated metal phase is acceptable considering the simplicity of the calculations. The approach used fails to predict the higher carbon contents. Manganese content is in general overestimated but still, not too exaggerated. The results of the chromium calculations are satisfactory, which is better illustrated in Figure 17 below.

The slag phase is harder to compare as the calculations were performed while excluding some oxides, e.g. those of titanium and vanadium. However, since the oxides incorporated constitute at least 90 % of the actual slag weight in all cases, it was decided to make this plain comparison. For the slag phase the overall comparison is worse than for the metal. The calculated MgO contents that are considerably lower than the actual, e.g. S1830 and S1832, can be explained by the fact that solid MgO was allowed to form in the calculations. The solids were not incorporated in the calculated slag analysis presented below. Also, the analysis provided for the end of blow slag composition is suspected to be off in those cases due to possible bits of refractory in the sampling procedure. This is based on the fact that samples of the slag during the blow showed considerably less MgO. The oxide which deviate the most is FeO which is significantly overestimated. This is true for both heats with accurately predicted carbon contents and those with too low. When employing a different approach in the calculations, adding a certain amount of O2 in steps to reach the desired carbon content, the same scenario with high FeO was achieved for the lower carbon contents. In [26] it is stated that the oxygen potential of the melt controlled by the carbon content is higher than that controlled by the slag FeO content. This explains that reaching the carbon content in an equilibrium calculation will result in a higher slag FeO content.

34

Finally it should be commented that the calculated slag weight is based on the overestimated FeO. But, since some oxides are left out, the weight is in the range of the actual slag weight for most heats.

Table 7: Comparison of calculated and actual analyses for three heats from the Imphos trials.

Parameters S1832 S1830 S1839 Calculated Actual Calculated Actual Calculated Actual Metal

%Fe

99.70 99.68 99.76 99.80 99.78 99.85 %C

0.035 0.070 0.028 0.030 0.020 0.020

%Mn

0.19 0.19 0.13 0.12 0.10 0.08 %Cr

0.012 0.013 0.009 0.010 0.008 0.008

Slag %CaO

49.2 41.7 44.8 36.0 38.2 37.0 %SiO2

12.9 12.4 9.4 9.0 9.8 11.5

%FeO

30.6 16.6 38.1 24.4 43.3 33.6 %MgO

4.5 16.8 4.8 18.3 4.3 5.5

%MnO

2.4 3.7 2.3 3.5 3.7 3.8 %Cr2O3

0.31 0.32 0.33 0.35 0.39 0.43

B2 Basicity

3.8 3.4 4.8 4.0 3.9 3.2 Slag rate kg/tCS 103 115 118 135 154 144

Figure 15 illustrates a comparison between the calculated and actual FeO content for all the heats subjected to the calculations. The figure suggests a systematic overestimation of the FeO content which is best described by a power function.

35

Figure 15: Comparison between actual and calculated (%FeO). A fitted power function with corresponding R2 value is also presented.

One way to describe the behavior of chromium is by comparing the slag FeO content and chromium distribution ratio as presented in Figure 16. The FeO for the calculated values have been recalculated using the function presented in Figure 15 above. The expected increase in the distribution ratio with increasing FeO [17] is seen for both the actual and calculated values. The distribution ratio of the calculated values corresponds well to the actual values.

Figure 16: Chromium distribution ratio as a function of (%FeO). The points are presented in pairs for the specific charges. Where it is uncertain which actual value corresponds to a calculated, arrows are added as aid.

y = 0,0109x2,1344 R² = 0,9623

10,0

15,0

20,0

25,0

30,0

35,0

25,0 27,0 29,0 31,0 33,0 35,0 37,0 39,0 41,0 43,0 45,0

Actu

al (%

FeO

)

Calculated (%FeO)

10,0

15,0

20,0

25,0

30,0

35,0

40,0

10,0 15,0 20,0 25,0 30,0 35,0

(%Cr

)/[%

Cr]

(%FeO)

Calculated Actual

36

Another way to describe the behavior of chromium is to study the yield. Actual and calculated yield for the heats are presented in Figure 17 below. The results are satisfying although showing a tendency of underestimating the yield. This may be explained by the more oxidized slag in the calculations. There are two calculated values deviating considerably from the actual values; namely, S1844 and S1831. The latter is explained by a poor chromium balance. The input to the converter was 1.6 kg whereas the output was 2.1 kg. When recalculating the yield using the output as basis the result is an actual yield of 30.9 %. An explanation why the other heat, S1844, was deviating could not be found.

Figure 17: Actual and calculated chromium yield for the Imphos heats.

From the results presented above it is clear that despite the simple calculation method applied the results are satisfying when considering chromium. However, since the overall results of the calculations, with respect to e.g. carbon content of the steel and general slag analysis was off it lowers the confidence in the results.

4.2.2 SSAB Calculations In this section the results of the calculations for the SSAB heats are presented and discussed.

A comparison between the calculated and actual analyses for three heats of the SSAB dataset is presented in Table 8 below. The remaining heats are presented in appendix 4. As the calculations were designed to reach the desired carbon content it is by default on point. However, the amount of oxygen used in the calculations is considerably lower than that of the actual input. Also, the oxygen content of the melt is lower for the specific carbon content. The manganese content is consistently overestimated

15,0

20,0

25,0

30,0

35,0

40,0

45,0

S1827 S1832 S1844 S1831 S1829 S1846 S1830 S1845 S1833 S1847 S1839

Cr y

ield

(%)

Heat

Calculated Actual

37

whereas the chromium content is varying in comparison to the actual value. The chromium is more thoroughly discussed in conjunction with Figure 18 below.

As for the Imphos trials, the comparison of the slag phase will differ by default since the thermodynamically calculated analysis disregard oxides of e.g. vanadium and titanium. However, the same applies here as for the Imphos calculations i.e. the oxides incorporated constitute at least 89 % of the total slag weight. The calculated slag analysis is provided in two different ways; one with solid MgO and CaO incorporated in the analysis and one where they are not. Focusing on the analysis incorporating the solid phases it is clear that the calculations provide a poor description of the process. The calculated basicity is mostly in the range of the actual. Also, the MgO content differ about one percent unit for all nine heats, which is regarded as acceptable. Apart from this, all other slag related parameters are off.

Contrary to the Imphos trials, the FeO is underestimated in these calculations. Only for three out of four heats with carbon content lower than 0.04 % the FeO is overestimated. The calculated carbon content of the Imphos trials were in general below the actual, except for the heats with very low carbon contents. This may constitute the difference in FeO. The previous explanation provided for the overestimation of FeO can still be considered true, but only for the lower carbon contents. The carbon content, slag FeO content and added oxygen, as seen in Table 8 and appendix 4, provides some insight towards the kinetics of the process. It has been shown [27] [29] that a model employing both thermodynamics and kinetics with the standpoint that the injected oxygen mainly oxidizes iron is very successful in predicting the outcome. The approach used in this thesis fails to capture the oxidation of carbon by iron oxide since most of the oxygen is used directly in the reaction with carbon. Thus, using thermodynamic equilibrium in the case of carbon contents around 0.05 % the correct carbon content is reached before iron oxidation has prolonged long enough, leaving a slag low in FeO and an underestimated oxygen input. For lower carbon contents, as seen in the Imphos trials, it is more thermodynamically favorable to oxidize iron than carbon leaving a slag rich in FeO. However, in both cases the oxidation of carbon by iron oxide is not considered.

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Table 8: Comparison between the thermodynamically calculated analyses and the SSAB analyses. Values in parenthesis are slag analyses not considering the solid CaO and MgO as part of the analysis.

Parameters J9924 H4755 S1157 Calculated Actual Calculated Actual Calculated Actual Added O2 kg 7572.4 8643.1 7669.5 8232.0 7669.5 8898.3

Metal %Fe

99.54 99.69 99.67 97.79 99.62 99.80 %C

0.048 0.048 0.037 0.037 0.039 0.039

%Mn

0.34 0.21 0.23 0.13 0.27 0.13 %Cr

0.027 0.025 0.018 0.025 0.017 0.017

ppm O

415 / 532 568 521 982

Slag %CaO

62.2 (54.5) 47.8 51.8 (47.8) 46.8 50.9 (50.7) 37.8 %SiO2

17.2 (21.5) 12.4 9.3 (12.4) 7.9 14.3 (15.5) 10.9

%FeO

13.5 (16.8) 17.7 23.2 (30.9) 19.4 23.5 (25.4) 31.9 %MgO

4.8 (4.5) 6.6 12.9 (5.2) 12.0 8.4 (5.4) 9.7

%MnO

1.9 (2.4) 4.3 2.3 (3.1) 3.6 2.5 (2.7) 3.0 %Cr2O3

0.26 (0.32) 0.22 0.38 (0.51) 0.25 0.31 (0.33) 0.24

B2 Basicity 3.6 (2.5) 3.9 5.6 (3.9) 6.0 3.6 (3.3) 3.5 Slag rate kg/tCS 66 (53) 89 95 (71) 112 87 (81) 118 %Solids 19.7 24.8 7.5

Opposed to Figure 15 above, there were no correlation between calculated FeO and the FeO reported by SSAB. As previously mentioned, the slag analysis provided by SSAB was calculated by the software at plant site. The comparison is thus between two calculated values which may explain the absence of a correlation.

Figure 18 below illustrates the comparison between calculated and actual chromium yield. The results are not as satisfying as the Imphos calculations. The calculated yield of heats J3399, J9924, X3452 and S1157 agrees well with the actual values while the remaining heats are off. With the exception of S1157 it is noted that heats with carbon content below 0.04 % show an underestimated chromium yield, i.e. H0736, H4755 and X9057. The remaining heats, with carbon content close to 0.05 %, all show a more or less overestimated chromium yields. The difference of H1231 is exaggerated. The heat was charged with 15 ton of scrap rich in chromium with suspected overestimated chromium analysis. The over- and underestimations can be coupled to the FeO content of the slag, which follows the same pattern.

It is noted that the chromium yield is higher for the SSAB heats than for the Imphos trials. If this depends on the lower carbon contents and higher slag FeO contents for

39

the Imphos trials or difference in size is unknown. However, in the calculations this is managed by using the two different approaches.

Figure 18: Comparison between actual and calculated chromium yield for the nine heats of the SSAB dataset.

A comparison involving the chromium distribution ratio was decided to be left out since the dataset was not provided with actual slag analyses. The yield was considered to be a better approximation since the chromium content of the steel was provided.

As a final remark two different methods were used in the calculation of the slag analyses. For the Imphos trials it was decided not to incorporate solid oxides into the analysis as an actual slag sampled was used as reference. For the SSAB calculations the solids were incorporated since the provided slag analysis in the dataset was calculated, probably based on material balance.

4.2.2.1 Cobalt Calculations The results from the cobalt calculations are presented in Table 9 below. The calculations show that, from a thermodynamic standpoint, cobalt reports exclusively to the crude steel. The actual yield is above 100 % which could be expected considering the assumptions made regarding the input of cobalt.

Table 9: Calculated and actual cobalt yield for three heats from the operation at SSAB Luleå. The yield is given in %.

Heat Calculated Actual J1814 99.9 104.0 J1829 99.9 110.9 J1840 100.0 106.5

30,0

40,0

50,0

60,0

70,0

80,0

90,0

S2220 J3399 J9924 H1231 H0736 H4755 X3452 X9057 S1157

Cr y

ield

(%)

Heat

Calculated Actual

40

The output of cobalt to the gas cleaning system as percentage of total input was estimated to 0.9 %. Furthermore, from trace element analyses of BOF slag it was estimated that 0.6 % of the total cobalt input reports to the slag phase. From the above, it is clear that the calculations agree with the actual values, i.e. that essentially all cobalt reports to the crude steel phase. However, knowing this makes the calculations excessive.

5 Discussion on the Thermodynamic Data used This section presents a discussion about the data and databases used in the calculations.

Since the work is based on equilibrium calculations it is natural to question and evaluate the data and databases utilized. As described earlier the solution phase representing the slag was taken from the FToxid oxide database. In [32] the database is explained for presenting that the Al2O3-CaO-FeO-Fe2O3-MgO-SiO2 system has been fully optimized for temperatures and oxygen partial pressures within this work. These oxides cover the major constituents of both the blast furnace and BOF slag. In addition to these components, it is explained that CrO, Cr2O3 and ZnO have been optimized for composition ranges common for both ferrous and non-ferrous applications. When considering the documentation found in [35], describing the used slag phase FToxid-SLAGA, it is clear that ZnO is described for non-ferrous metallurgy. The references used when optimizing ZnO is for slag systems in compositional ranges of zinc and lead-making. As this could impose faults in the results, calculations using the activity coefficient of ZnO for the CaO-SiO2-Al2O3-ZnO system found by [36] were used as comparison. Using the new activity coefficient a higher ZnO content in the slag was achieved; however, still below the ppm range. As argued by [37], the high oxygen potential used in [36] does not represent that of the blast furnace. From the oxidizing conditions of [36] along with the extensive optimization of the components in the FactSage FToxid database it was decided to use the FToxid-SLAGA for ZnO calculations.

The documentation of FToxid-SLAGA [35] presents that no slag system containing oxides of both chromium and iron have been assessed. However, chromium have been have been fully optimized for the Al2O3-CaO-CrO-Cr2O3-SiO2 system and roughly optimized for the CrO-Cr2O3-MgO-SiO2 system. And, as presented above, iron has been assessed as a major constituent in the same systems but in absence of chromium. Therefore, it is assumed that there exists an indirect correlation between iron and chromium and that the results from the calculations are valid in the dilute chromium contents of this work.

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Two different expressions for the activity coefficients of zinc in the iron-carbon-zinc system were found in the available literature. Due to the choice of experimental method of [15] and the temperature interval ranging higher than that of [14], it was decided to use the activity coefficient presented by [15]. The argument against this choice is that the experimental work of [14] was performed in a compositional range closer to that of the present work with regards to zinc. It is realized that the activity coefficient of zinc in hot metal could be the largest source of error in the blast furnace calculations together with the assumption of recirculation rate.

The behavior of chromium in liquid iron is in the FTmisc-FeLQ database expressed by the unified interaction parameter formalism [38]. An associate model for the deoxidation products [39] of amongst others chromium is also used in the database. The use of this database is considered to be the best option when calculating the crude steel composition. Errors in chromium content are more likely to be an effect of e.g. inability to consider kinetic factors.

6 Concluding Discussion This section presents a concluding discussion for both the blast furnace and BOF.

6.1 Blast Furnace From the calculations it is clear that trace elements that leave the furnace through the top as dust and sludge, such as zinc and lead, cannot be described using only thermodynamics. The more sophisticated models presented in [24] and [25] both neglected the solids in the off gas. However, the model in [25] was still able to describe the recirculating behavior of alkalis since they mainly report to the slag phase. It is believed that a thorough study of the dust and sludge with respect to amount and composition for different operational conditions of blast rate, top pressure, top gas temperature and burden distribution is needed in order to complement a model based on thermodynamics. However, the calculations of this thesis may provide some insight towards how much of e.g. the effect of lowered top pressure is explained by thermodynamic prerequisites and by change in furnace parameters.

The calculations also display the importance of other mechanisms in e.g. the estimation of hot metal silicon and ZnO content of the slag. To describe the latter deviation from equilibrium, consideration of e.g. the slag forming may be needed.

The shortage of data only provided one analysis of zinc in hot metal within the detection limit. The calculated value was underestimated as compared to this measurement. From the calculations representing the region above the thermal reserve zone it is concluded that increased top pressure is likely to increase the recirculation rate of zinc. Increasing the recirculation rate in the calculations would bring the results

42

closer to the actual measurements. However, more data to compare with is required to find confidence in such an assumption.

As a final remark, the trace elements studied in this thesis all show a cyclical behavior. Other elements with high boiling points and low vapor pressures may possibly be described using a simple hearth equilibrium calculation, provided that the elements are described by the thermodynamic databases.

6.2 Basic Oxygen Furnace As for the pure thermodynamic model of the AOD presented in [20] the same arguments can be applied in this thesis; namely, expectations on a thermodynamic model should not be high for a process where kinetics is known to play an important part. Also, the fix temperature during the entire blow favors a more effective decarburization compared to an increasing temperature.

In [23] it was shown that a model considering both thermodynamics and kinetics better described the PS-converting in copper making than a model considering only thermodynamics. The models considering both thermodynamics and kinetics [22] [27] [29] in the BOF accurately predicted the outcome of the heat. It was also shown by [21] that thermodynamic calculations similar to those of this thesis are possible to use in other pyrometallurgical processes where kinetic factors can be neglected. From this together with the results of this thesis it is concluded that a pure thermodynamic approach is not suitable to employ when calculating the outcome of the industrial scale BOF.

However, the simple approach did describe the chromium yield of a small converter with proper accuracy while failing to predict e.g. the slag analysis. The results for the larger converter were worse, both for chromium yield and overall results. It is likely that kinetic factors are more pronounced in a converter that is about 20 times larger.

7 Conclusions It was not possible to describe the distribution of zinc and lead in the blast furnace using the approach developed in this thesis.

For the BOF it was possible to estimate the distribution of chromium for a pilot plant sized converter. However, the method could not predict the overall outcome satisfactorily which lowers the confidence in the results. If the used method can be applied on other trace elements could not be distinguished since no such data was available for comparison.

The thermodynamic calculations of this thesis could not describe the chromium distribution of an industrial scale BOF. It requires models incorporating both kinetics

43

and thermodynamics to accurately describe both chromium distribution and overall process outcome.

Cobalt could be described using the thermodynamic equilibrium approach. However, a material balance of cobalt with the assumption that essentially all cobalt reports to the crude steel phase would work just as well.

8 Further Studies If there is an interest to describe the output of zinc and lead through the blast furnace top a thorough study of the amount and composition of dust and sludge depending on furnace parameters should be carried out.

A dynamic model of the blast furnace incorporating thermodynamics coupled with the experimental description of the output through the top could be designed with the goal of successfully estimating the output of volatile trace elements through both the top and tap hole while at the same time estimating the recirculation rates.

The distribution of non-volatile trace elements can be investigated and compared to the hearth equilibrium calculations to establish if the approach is applicable for such elements.

If the trace element distribution of the BOF is to be described in a reliable way, a model incorporating both thermodynamics and kinetics should be developed. Such a model is considered to have advantages over a pure mathematical model since it is based on the fundamental outline of the process as well as the reaction and mass transfer phenomena in the process. Also, relying on this, it is more flexible going outside the boundaries set for a mathematical model. The downside of such a model is cost for additional software and development.

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[37] D. Lindström and D. Sichen, "Study on the Possibility of Using ZnO to Increase the Desulfurization Potential of Blast Furnace Slag and Sulfide Capacities," Steel Research International, vol. 84, no. 1, pp. 48-55, 2013.

[38] I.-H. Jung, "Overview of the applications of thermodynamic databases to steelmaking processes," Calphad, vol. 34, no. 3, pp. 332-362, 2010.

[39] I.-H. Jung, S. A. Decterov and A. D. Pelton, "A Thermodynamic Model for Deoxidation Equilibria in Steel," Metallurgical and Materials Transactions B, vol. 35, no. 3, pp. 493-507, 2004.

[40] Verein Deutscher Eisenhuttenleute, Slag Atlas, Germany: Verlag Stahleisen, 2nd Edition, 1995.

48

10 Appendix 1 – Ingoing Values to the Blast Furnace Calculations The input used in the different calculations is presented in this appendix. All zinc related blast furnace calculations can be remade and altered with the information from this chapter.

10.1 Hearth Equilibrium Calculations The input for the hearth calculations for the 2012 dataset is presented in Table 10 below. The temperature was set to 1480 °C, pressure 3.4 atm and the A and B values in Equation 3 as 1629 and 0.141 respectively.

Table 10: Input for hearth calculations on 2012 dataset.

Species Mass (kg) Fe 947.5 Mn 3.8 C 413.1 CaO 61.7 MgO 24.9 SiO2 65.4 Al2O3 20.9 O2 331.9 N2 937.6 Zn 1.0

The input for the hearth calculations for the 2006 dataset is presented in Table 11 below. The temperature was set to 1478 °C, pressure 3.5 atm and the A and B values in Equation 3 as 1630 and 0.142 respectively.

Table 11: Input for hearth calculations on 2006 dataset.

Species Mass (kg) Fe 950.3 Mn 3.8 C 415.5 CaO 59.0 MgO 27.3 SiO2 64.0 Al2O3 19.7 O2 329.1 N2 866.3 Zn 2.9

49

The calculations conducted to find the general trends were based on the ingoing values presented in Table 10 with the exceptions presented below. For the increased recirculating load, zinc was altered between 0 kg and 20 kg with increments of 1 kg. For the change in hearth temperature, the temperature was changed between 1400 °C and 1600 °C with increments of 10 °C. The calculations of changing pressures were conducted with pressures between 2 atm and 4 atm using increments of 0.25 atm.

The input data for the effect of assumption calculation presented in Figure 8 is given in Table 12 and Table 13 below. In both cases the temperature and pressure was set as 1480 °C and 3.4 atm respectively. The value of A was varied between 10 and 100 with increments 10 in a closed system.

Table 12: Input in calculations of effect of assumptions with changing zinc content.

Species Mass (kg) Fe 947.5 Mn 3.8 C 164.2 CaO 61.7 MgO 24.9 SiO2 65.4 Al2O3 20.9 CO <5.808A> N2 <9.376A> Zn <0.01A>

Table 13: Input in calculations of effect of assumptions with fixed amount of zinc oxide input.

Species Mass (kg) Fe 947.5 Mn 3.8 C 164.2 CaO 61.7 MgO 24.9 SiO2 65.4 Al2O3 20.9 CO <5.808A> N2 <9.376A> ZnO 1.245

50

10.2 Thermal Reserve Zone Calculations The ingoing gas phase and solids are presented in Table 14 below. As described in the method, carbon and the gas phase are output values from the hearth calculations.

Table 14: Input in the thermal reserve zone calculations.

Species Mass (kg) FeO 1189.8 Fe 16.3 CaO 46.2 CaCO3 12.8 MgO 24.0 SiO2 37.8 Al2O3 8.1 C 109.4 CO 588.5 N2 937.4 Zn 0.9973

10.3 Above the Thermal Reserve Zone and Below the Throat The input and calculation scheme for the different top pressures are presented in Table 15 and Table 16 below.

Table 15: Input in the calculations. The A value is explained for in the table below.

Species Mass (kg) CO <843.7-458.1A> CO2 <719.9A> N2 937.4 Zn 0.9973

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Table 16: Calculation scheme for the different top pressures. The parameter A is from Table 15.

1 atm Top P 1.5 atm Top P 2 atm Top P P (atm) P (atm) P (atm) T (°C) A

1.3000 1.8000 2.3000 900 0.455 1.2625 1.7625 2.2625 850 0.523 1.2250 1.7250 2.2250 800 0.591 1.1875 1.6875 2.1875 750 0.659 1.1500 1.6500 2.1500 700 0.727 1.1125 1.6125 2.1125 650 0.795 1.0750 1.5750 2.0750 600 0.863 1.0375 1.5375 2.0375 550 0.931 1.0000 1.5000 2.0000 500 1

52

11 Appendix 2 – Phase Diagrams Figure 19 below presents the ternary phase diagram of the CaO-SiO2-MgO system with four heats marked.

Figure 19: The ternary phase diagram of the CaO-SiO2-MgO system with four heats marked S1847 (1), H1231 (2), J3399 (3) and S1831 (4). Original figure found in [40].

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Figure 20 below presents the ternary phase diagram of the CaO-SiO2-FeO system with three heats marked.

Figure 20: The ternary phase diagram of the CaO-SiO2-FeOx system with three heats marked H1231 (1), S1157 (2) and S1839 (3). Original figure found in [40].

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12 Appendix 3 – Results from Calculations on the 2006 Dataset Table 17 below presents the comparison between the calculated and actual hot metal analysis for the 2006 dataset.

Table 17: Actual and calculated hot metal analysis for the 2006 dataset.

Element Calculated Actual %C 5.12 4.72 %Fe 94.2 94.0 %Mn 0.36 0.33 %Si 0.29 0.40 %Zn 0.0008 0.0012

Table 18 below presents the comparison between the actual and calculated slag analysis for the 2006 dataset.

Table 18: Actual and calculated slag analysis for the 2006 dataset.

Parameter Calculated Actual %CaO 36.1 31.9 %SiO2 35.2 33.6 %MgO 16.6 17.0 %Al2O3 12.0 12.6 %MnO 0.08 0.33 %ZnO 0.0000 <0.0005 B2 1.0 0.9 Slag rate kg/tHM 162 165

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13 Appendix 4 – Results from Basic Oxygen Furnace Calculations The full results from the calculations are presented below.

13.1 Imphos Dataset Table 19 through Table 21 below presents the full results from the Imphos calculations.

Table 19: Comparison of calculated and actual analyses for the three heats S1827, S1828 and S1829 from the Imphos trials.

Parameters S1827 S1828 S1829 Calculated Actual Calculated Actual Calculated Actual Metal

%Fe

99.62 99.62 99.70 99.79 99.72 99.75 %C

0.033 0.09 0.027 0.04 0.033 0.05

%Mn

0.28 0.21 0.19 0.12 0.17 0.14 %Cr

0.016 0.015 0.010 0.007 0.010 0.010

Slag %CaO

49.5 52.3 44.4 42.5 47.1 38.1 %SiO2

13.4 12.9 10.5 10.9 10.7 11.1

%FeO

28.8 14.0 36.3 24.4 34.7 21.9 %MgO

4.6 7.4 5.1 8.9 4.5 16.7

%MnO

3.1 3.3 3.4 3.8 2.6 3.5 %Cr2O3

0.36 0.31 0.33 0.34 0.30 0.30

B2 Basicity

3.7 4.1 4.2 3.9 4.4 3.5 Slag rate kg/tCS 88 95 121 120 112 123

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Table 20: Comparison of calculated and actual analyses for the three heats S1832, S1833 and S1844 from the Imphos trials.

Parameters S1832 S1833 S1844 Calculated Actual Calculated Actual Calculated Actual Metal

%Fe

99.73 99.75 99.71 99.80 99.67 99.72 %C

0.034 0.050 0.033 0.030 0.029 0.030

%Mn

0.17 0.15 0.18 0.13 0.22 0.17 %Cr

0.26 0.33 0.010 0.010 0.015 0.018

Slag %CaO

47.4 39.0 45.5 43.6 46.6 45.8 %SiO2

10.7 10.5 8.0 9.8 13.3 16.4

%FeO

34.7 19.3 38.5 25.7 32.0 17.8 %MgO

4.4 18.9 4.6 8.4 3.9 5.5

%MnO

2.5 3.8 2.9 3.6 3.7 4.1 %Cr2O3

0.26 0.33 0.37 0.36 0.43 0.4

B2 Basicity

4.4 3.7 5.7 4.4 3.5 2.8 Slag rate kg/tCS 102 117 133 129 103 93

Table 21: Comparison of calculated and actual analyses for the three heats S1845, S1846 and S1847 from the Imphos trials.

Parameters S1845 S1846 S1847 Calculated Actual Calculated Actual Calculated Actual Metal

%Fe

99.68 99.76 99.69 99.74 99.73 99.76 %C

0.023 0.02 0.024 0.05 0.032 0.08

%Mn

0.20 0.15 0.18 0.15 0.17 0.11 %Cr

0.014 0.015 0.013 0.014 0.014 0.014

Slag %CaO

44.0 42.4 43.4 41.8 44.7 40.6 %SiO2

10.3 13.1 10.7 11.4 8.6 8.7

%FeO

37.8 25.0 36.5 23.9 38.5 29.2 %MgO

3.5 6.0 5.5 9.2 4.4 8.3

%MnO

3.7 4.1 3.4 3.9 3.1 3.6 %Cr2O3

0.46 0.42 0.42 0.41 0.51 0.49

B2 Basicity

4.3 3.2 4.1 3.7 5.2 4.7 Slag rate kg/tCS 118 106 117 114 111 116

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13.2 SSAB Dataset Table 22 and Table 23 below presents the full results from the calculations representing the SSAB dataset.

Table 22: Comparison between thermodynamically calculated analyses and SSAB analyses. Values in parenthesis are slag analyses not considering the solid CaO and MgO as part of the analysis.

Parameters S2220 J3399 H1231 Calculated Actual Calculated Actual Calculated Actual Added O2 kg 7523.9 8686.0 7281.2 8179.1 7960.7 8972.9

Metal %Fe

99.56 99.73 99.6 99.75 99.55 99.75 %C

0.050 0.050 0.050 0.052 0.049 0.050

%Mn

0.33 0.18 0.29 0.16 0.29 0.14 %Cr

0.025 0.019 0.020 0.019 0.062 0.039

ppm O

394 568 401 533 400 519

Slag %CaO

62.6 (54.8) 50.1 60.9 (56.1) 43.9 61.5 (54.4) 44.6 %SiO2

13.0 (22.2) 11.4 15.0 (23.1) 12.4 13.6 (21.5) 10.4

%FeO

9.4 (16.0) 15.1 9.5 (14.6) 17.7 10.6 (16.7) 18.3 %MgO

13.4 (4.2) 13.2 13.5 (4.4) 14.5 12.3 (4.3) 12.5

%MnO

1.4 (2.4) 2.7 1.1 (1.6) 3.5 1.4 (2.2) 4.0 %Cr2O3

0.18 (0.30) 0.24 0.12 (0.19) 0.11 0.49 (0.78) 0.63

B2 Basicity 4.8 (2.5) 4.4 4.1 (2.4) 3.5 5.4 (2.5) 4.3 Slag rate kg/tCS 90 (52) 111 56 (36) 72 98 (62) 132 %Solids 41.5 35.2 36.8

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Table 23: Comparison between thermodynamically calculated analyses and SSAB analyses. Values in parenthesis are slag analyses not considering the solid CaO and MgO as part of the analysis.

Parameters H0736 X3452 X9057 Calculated Actual Calculated Actual Calculated Actual Added O2 kg 8494.7 8767.4 7620.9 8801.6 8203.4 9289.1

Metal %Fe

99.68 99.73 99.6 99.76 99.67 99.79 %C

0.036 0.036 0.052 0.054 0.035 0.034

%Mn

0.20 0.18 0.28 0.12 0.22 0.13 %Cr

0.025 0.040 0.032 0.030 0.035 0.034

ppm O

546 608 382 413 577 681

Slag %CaO

49.8 (45.9) 47.7 68.6 (56.2) 44.8 46.9 (48.0) 45.3 %SiO2

8.0 (10.0) 8.5 13.1 (23.6) 11.1 12.6 (12.9) 11.9

%FeO

27.7 (34.6) 18.5 7.9 (14.1) 20.1 29.3 (30.0) 20.2 %MgO

11.2 (5.3) 10.3 9.3 (4.1) 9.3 7.9 (5.6) 11.7

%MnO

2.6 (3.3) 4.5 0.9 (1.6) 3.2 2.8 (2.9) 2.6 %Cr2O3

0.66 (0.83) 0.46 0.17 (0.31) 0.18 0.49 (0.50) 0.37

B2 Basicity 6.3 (4.6) 5.6 5.2 (2.4) 4.0 3.7 (3.7) 3.8 Slag rate kg/tCS 110 (88) 108 86 (48) 111 105 (102) 108 %Solids 20.1 30.8 2.5