Inclusion formation afs

The term "clean steel" is commonly used to describe steels that have low levels of the solute elements sulfur,phosphorus, nitrogen, oxygen and hydrogen; controlled levels of the residual elements copper, lead, zinc,nickel, chromium, bismuth, tin, antimony and magnesium; and, a low frequency of product defects that canbe related to the presence of oxides created during the act of steelmaking, ladle metallurgy, casting androlling. This last definition causes extreme problems to the steel manufacturer as the definition of "clean" isnot absolute, is based upon the product formed from the casting and the in-service use or life of the product.In addition, the definition "clean" is comparative as each customer of a steel product has the ability to buysteel from around the world and compare the performance level of a given product based upon the supplier.In this system, as steel is a commodity, the best steel producer defines the level of quality that is expected bya customer and, as steel producers are continually striving to produce "cleaner" steels, the cleanlinessstandard desired by the customer is continuously changing as a function of time and technologicalimprovements. The term "clean steel" is therefore continually variable depending upon the application andthe competition between steel suppliers.

Due to the variable nature of the term "clean steel" it is more accurate to talk about "high purity steels" assteels with low levels of solutes and "low residual steels" as steels with low levels of impurities that originatefrom scrap remelting and "clean steels" as those steels with a low frequency of product defects that can berelated to the presence of oxides. For example, there are "high purity, low residual clean steels" such as ultradeep drawing steel sheets for automobiles which require ultra-low carbon contents (< 30 ppm), low nitrogencontents (< 30 ppm) and the absence of oxide inclusions with diameters greater than 100 microns; and "lowresidual clean steels", such as those used for drawn and ironed cans, which are a standard low carbon steel(1006), without particular high purity requirements, but are ultra clean with the requirement that oxidediameters must be less than 20 microns. In addition, in forging and bearing grades, there are "clean steels"that require strictly controlled inclusion size distributions.

As measurement is a key in the determination of permissible inclusion content thin products which undergosignificant drawing are very susceptible to variations in inclusion size distribution, as cracks or wire breakscan be easily counted. For example, in drawn and ironed cans, it is very common to count the number ofcracked flanges per million formed cans and generally a performance of less than 10 ppm would be lookedupon as an excellent performance. Another application which is amenable to measurement is bearing life. It iswell known that total inclusion content (as measured by total oxygen content) has traditionally correlatedwith bearing life and decreased total oxygen contents (below 10 ppm) have lead to significant increases inbearing life and thus the drive to very low inclusion contents in bearing steels. In addition to total oxygencontent, the total length of stringer inclusions after forging correlates well with bearing life and, at low totaloxygen levels, efforts to reduce inclusion clustering leads to very long fatigue life for bearings.

There is one constant in the world of high purity, low residual and clean steels and that is the continual driveto reduce solute and residual contents in all steels and to control the frequency and size distribution of theinclusions that are found in all steels. Thus this chapter will focus upon an understanding of the fundamentalsof the production of high purity, low residual clean steels, emphasizing the limits in current technology andthe potential for the production of high purity, low residual steels with a low frequency of inclusions withaverage diameters less than 5 microns.

Clean steels are steels with a low frequency of inclusions of average diameter less than 5 microns. The majorproblems in clean steel manufacture are incomplete separation of clustered solid inclusions (> 5 microns indiameter), the presence of sporadic larger liquid inclusions due to emulsification of covering slags and the

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presence of solid materials that originate from the refractories used to contain steels. The equipment used toproduce clean steel varies greatly between different steel plants; however, current clean steelmaking andcasting practices are based upon the following principles:

Oxygen, which is dissolved in liquid steel at the steelmaking and melting stage, must betransformed into a solid or a gas and removed before casting.

The external oxygen sources which are responsible for the reoxidation of liquid steel, must beeliminated at every step in the process.

The physical entrapment of the liquid fluxes used during steel refining and casting must beeliminated.

Refractories in contact with liquid steel must be chemically stable and resistant to corrosion anderosion.

These practical principles of clean steel manufacture are based upon an understanding of the importance ofmaintaining chemical equilibrium between the elements dissolved in liquid steel and the slag and refractorysystems which are in contact with the liquid steel, and, of controlling fluid flow to avoid conditions at liquidslag-steel interfaces which could result in the physical entrapment of the covering slag.

Clean steel manufacture is dependent upon an understanding of the fundamental steps necessary to produce aclean steel:

generation of the inclusion;transport of the inclusion to an interface;separation of the inclusion to the interface; and,removal of the inclusion from the interface.

Success or failure in the production of clean steel is dependent upon a complete understanding and executionof these four issues.

1. Inclusion Generation

The formation of inclusions during steelmaking is inevitable as oxygen is more soluble in liquid iron than insolid iron. In addition, at 1600 C and gas phase oxygen partial pressures greater than 6 x 10-9 atmospheres,liquid iron will spontaneously oxidize to liquid iron oxide. The saturation limit for oxygen in liquid iron incontact with iron oxide at 1600 C is 0.23 wt%, a value which decreases with temperature according to theequilibrium between iron, oxygen and iron oxide[1-2] The soluble oxygen content of pure liquid iron incontact with a gas phase can be reduced if the gas phase oxygen partial pressure is less than 6 x 10-9 atm: apartial pressure which can be achieved with carbon monoxide/carbon dioxide mixtures with less than 1%carbon dioxide, for example. Practically, it is necessary to shield liquid iron at all times to prevent theformation of iron oxide and generally this is accomplished by use of a liquid slag, with a low diffusivity ofoxygen, as a physical barrier between the ambient atmosphere and the liquid steel.

The solubility limit for oxygen in solid iron is low and the partial pressure of oxygen in equilibrium with solid


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iron and iron oxide at 1000 C is 1 x10-15 atm.[3]. Thus, if a pure liquid iron sample in equilibrium with FeO issolidified and cooled, iron oxide would precipitate interdendritically as spherical liquid iron oxide inclusions.

The key to making a clean steel is to determine the mechanism by which an oxide inclusion can be removedfrom liquid steel. For example, if it possible to transport the inclusions from the bulk of the liquid steel to afree surface, where the inclusions can separate from the liquid steel, the steel will become cleaner. Of course,the only inclusions which can be removed are those which have formed while the liquid iron is molten. Pureiron solidifies at 1534 C where the maximum solubility of oxygen is approximately 0.17 wt/o. Thus, between1600 C and 1534 C, 0.06 w/o soluble oxygen will transform to iron oxide. This is maximum possible amountof oxygen which could removed under these conditions and, regardless of one's processing abilities, it wouldbe impossible to produce pure solid iron with less than 1700 ppm oxygen in the solid state, if liquid iron wasallowed to become saturated with oxygen before solidification.

To produce liquid iron of lower than saturation oxygen content it is necessary to precipitate the oxygen, as anoxide, at liquid steel processing temperatures. Again, if liquid iron is used as an example, the equilibriumbetween liquid iron, oxygen dissolved in liquid iron and liquid iron oxide can be written as follows:

[Fe] + [O] = (FeO) ... [1]

where [O] is oxygen dissolved in liquid iron and (FeO) is pure liquid iron oxide. The equilibrium constant K1= 1/ao for pure liquid iron and iron oxide and the oxygen level at saturation is a function of temperature only.The equilibrium oxygen content at a given temperature in pure liquid iron can be reduced by modification ofthe activity of FeO by dissolution into another chemical species. For example, liquid calcium aluminate couldbe added to the liquid iron surface as a sink for the liquid iron oxide. As the iron oxide dissolved in calciumaluminate, the mole fraction of iron oxide present in the newly formed slag would be reduced and theequilibrium oxygen activity of the steel would also be reduced as the liquid iron attained equilibrium with theiron oxide activity in the slag. Thus, modification of the activity of the deoxidation products the first methodthat can be used to reduce dissolved oxygen levels during processing.

A second method to reduce dissolved oxygen levels is to modify chemical equilibrium by changing theequilibrium reaction. In steelmaking, carbon is always present, and the equation [2] often sets dissolvedoxygen activities.

[C] + [O] = CO ...[2]

In the BOF, for example, carbon content sets the lower limit for the oxygen activity at a given temperatureand as the CO pressure approximates to 1 atmosphere during refining, lower carbon contents lead to higheroxygen activities and higher soluble oxygen levels. Oxygen content in carbon deoxidized steels can be furthermodified by vacuum or inert gas treatment which will reduce the equilibrium carbon monoxide pressure andlead to the production of steels with lower oxygen contents at a given carbon level.

The equilibrium constant for reaction [2] increases with decreasing temperature and large quantities of COare evolved during solidification of steels containing only carbon as a deoxidant. This "rimming" phenomenawas utilized in ingot casting to produce a cleaner ingot surface; however, the porosity associated with evenmild rimming is undesirable in many castings and the rimming action can be controlled by addition ofdeoxidants to "kill" the action in the steel during solidification. Elements such as silicon, manganese andaluminum are routinely added to liquid steel to reduce oxygen activities to below that which will cause COevolution in a mold.

The practical limit of steel cleanliness in normal steels is set by chemical equilibrium and, to reach this value,all inclusions formed during the deoxidation reaction must be separated before solidification[4] . In general, in


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steels deoxidized with silicon and manganese, the equilibrium is set by the formation of a liquid manganesesilicate; in aluminum killed steels, by the formation of solid alumina; and, in calcium treated, aluminum killedsteels by the formation of a liquid calcium aluminate. To further reduce soluble oxygen content at a giventemperature, manganese silicate or alumina can be dissolved into a slag. For example, dissolution of aluminainto a slag can result in the soluble oxygen content of an aluminum killed steel being reduced to less than 2ppm. Dissolution of silica into a slag of calcium aluminate, for example, can significantly reduce the activityof silica and lead to soluble oxygen contents of silicon killed steels in the range of aluminum killed steels.

The use of chemical equilibrium to reach low soluble oxygen contents is widespread; however, to beeffective, all phases in contact with the liquid steel must be at the same equilibrium condition. If aluminumkilled steels are at equilibrium, reaction [1] must also be at equilibrium; therefore, FeO activity in slags mustbe very low or aluminum will oxidize by reaction with the slag. Similarly equilibrium MnO and SiO2 slagactivities in the slag are uniquely set by the addition of aluminum and any increases above the equilibriumslag levels will result in reaction with the slag and the formation of alumina. Therefore, in practice, the slagsystems must be designed to be very low in FeO and MnO after the steel has been killed. Currently, furnaceslags tend to be high in FeO and MnO and this has led to the development of slag-free tapping techniques and,also, slag killing using calcium carbide or aluminum.

Equilibrium between the refractories and the steel must also be considered. For example, high aluminarefractories are not 100% alumina and numerous binders are added which can contain oxides which are lessstable than alumina. Under these conditions the container refractories will react with the liquid steel, if notadequately designed. Future clean steel developments will be aimed at developing steels which are inequilibrium with both their container refractories and the slag present on top of the liquid steel.

Recently, residual magnesium from ferro-alloys, scrap or recycled aluminum or from reduction of refractorieshas caused significant processing problems due to low residual levels (< 5ppm) causing the primarydeoxidation inclusion to be transformed into a high magnesia containing inclusion which often is a magnesiumaluminate spinel, a highly refractory inclusion. Often, this causes nozzle clogging during steel pouring. Recentstudies by Itoh et al.[5] have fully documented the problem and magnesium residuals must be reduced to lessthan 1 ppm.

It is interesting that originally the term slag denoted "any non-metallic substance, formed together with ametal during a metallurgical process" [6]; however, this very general definition has been refined with time torefer only to the covering liquid oxides that are present on top of the liquid steel, i. e., ladle slags, tundishslags and mold slags. The primary and secondary deoxidation or reoxidation product are not referred to asslag inclusions but as simply non-metallic inclusions or inclusions.

A second source of cleanliness problems is incomplete separation of emulsified slags[7]. This occurs duringprocessing events where there is a high dissipation of energy at the slag-metal interface. Events such as vesselfilling, vessel drainage and level fluctuations at the slag metal interface are generally responsible for thegeneration of such inclusions. The size range of entrapped slag inclusions varies from less than 20 to greaterthan 200 microns; however, their frequency can be low. Generally, these type of inclusions are practicespecific and lead to cleanliness problems which are apparently random, until their source is recognized[8-13].

The last source of inclusions is the erosion and corrosion of refractories and carry-over of the nozzle wellsand during processing. These problems are solved through better process design.

The final level of cleanliness is a balance between the rate of inclusion removal and the rate of inclusionformation. Clearly to refine liquid steel the rate of removal must be greater than the rate of creation;therefore, reductions in the rate of reaction between the steel and the slag and the steel and the refractorysystem will enable the lowest possible inclusion contents in the shortest processing time. Reoxidation of liquidsteel by reaction with air causes very high rates of inclusion formation and, often, causes the rate of formation


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of inclusions to be greater than the rate of inclusion removal. Thus, a natural limit to any process whichenables inclusion removal is the point at which air reoxidation begins.

2. Transport to an Interface

In the above section on generation of inclusions it must be recognized that although thermodynamics canallow the prediction of the chemistry of the most stable inclusion, it does not allow a prediction of the size ofan inclusion and, when inclusion removal is a goal, the inclusion size distribution is a key to understanding therate at which an inclusion can be transported to a surface where it has the opportunity to separate from liquidsteel.

Deoxidation and reoxidation inclusions must first be nucleated within the liquid steel and once nucleatedthese inclusions will grow until the rate of growth is limited by diffusion in the liquid. The smallest measureddeoxidation inclusions are of the order of 15 nanometers[13,14] however, these quickly grow to between 1and 5 microns in diameter in practical steelmaking conditions. Liquid turbulence can aid in agglomeration ofinclusions and solid inclusions can easily cluster and form large three dimensional rafts of sintered smallinclusions. Thus it is not unusual to find a high frequency of small inclusions less than 5 microns in diameterand a much smaller frequency of inclusion rafts which are from 5 to 200 microns in diameter in liquid steels.The goal of clean steel manufacture is to minimize the total numbers of inclusions with diameters less than 5microns and to eliminate all clustered inclusions before casting.

Initial studies of inclusion removal were focused upon the removal of inclusions due to buoyancy driven flowwhere Stokes' Law applies[15-16]:

where VT is the sphere velocity, (( is the differential density between the particle and the liquid, ( is the liquiddynamic viscosity, g is acceleration due to gravity and r is the particle diameter.

Stokes law describes the velocity of a solid sphere under the influence of buoyancy forces, due to densitydifferences, in a static bath. Stokes Law only holds for rigid, spherical particles within the viscous flowregime, where the Reynolds number is less than 0.1. For spherical alumina particles in liquid steel theminimum particle radius for adherence to Stokes Law is 33.2 microns. For flows where the Reynolds numberis greater than 0.1, it is common to view the problem as a balance between the gravity force and thecombination of a buoyancy and a frictional force. Details of friction factors for different flow conditions andparticle types are given by Schwerdtfeger[15] and corrections for slip, interfacial tension, particle rigidity andwall effects are given by Iyengar et al.[16].

Unfortunately, actual metallurgical vessels are not quiescent and are subject to natural convection, as hasbeen pointed out by Szekely[17] and Iyengar[16] and inclusion removal rates in industrial vessels cannot beexplained solely by Stokes' Law and its modifications. For example, Iyengar calculated that the criticaltemperature gradient to avoid natural convection, for steel held in a small crucible, was 0.08 C/cm and thatfluid velocities of 0.015 cm/s were sufficient to ensure that the drag force due to the fluid flow was an orderof magnitude larger than the buoyancy force. In actual steelmaking practice, temperature gradients are highagainst the refractory container walls and natural convection occurs leading to natural stirring patterns withinthe ladle or the tundish. Liquid steel velocity calculations, by Joo et al.[18] indicate fluid velocities which aregreater than 1 cm/sec due to natural convection in ladles and tundishes.


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The effect of fluid flow in practical vessels is to provide a transport path to an interface and the problem canbe viewed as a standard boundary layer problem where fluid flow leads to a momentum boundary layer whichhas a thickness which can be related to the flow field and the physical properties of the fluid. Inclusions arecarried by fluid motion to a boundary layer at an interface where, due to buoyancy forces, they may traversethis boundary layer. In order to calculate, from first principles, inclusion removal in a stirred system, it isnecessary to calculate the exact form of the boundary layer and then calculate the fraction of inclusionswhich can traverse the boundary layer and separate during their contact time with the boundary layer. TheNavier Stokes equations must be solved for the coupled heat and mass transfer conditions to take into accountnatural convection, to determine the boundary layer thickness against the container walls and the slag-metalsurface (taking into account the fact that this is a liquid-liquid boundary). The local flow conditions for eachparticle at the boundary layer and the frequency of separation must then be determined as a function ofparticle size. Calculations of this type are quite numerically complex; however, results have been given byJoo et al.[18] where large inclusions with a large rising velocity separate more completely within the tundishand the efficiency with which inclusions are removed decreases with decreases in inclusion size. Joo's workindicated that even large inclusions can pass through a tundish and not be removed due to the effect of fluidflow and that particles with diameters less than 40 microns ( float out velocities of 0.5 mm/s) have removalefficiencies of less than 50%.

Inclusion removal in ladles has not, as yet, been computer modeled for inclusion removal in the same manneras tundishes; however, the basic phenomenon of liquid phase mass transfer is well recognized.Turkdogan[19,20] suggested an equation of the following type is an adequate representation of practicalresults:

Ct = Co exp ( - kt ) .... [3]

where Co is the initial inclusion content, Ct is the concentration of inclusions at a time t and k is the apparentflotation rate constant for a given type and intensity of stirring. Turkdogan models inclusion removal is a firstorder process where the removal rate is proportional to concentration. Schwerdtfeger[21] has also analyzedplant results of inclusion removal from aluminum-killed steels and found that a similar first order equationcould be used to describe the data from total oxygen analysis:

[O] = < [o]i - [o]e > exp [ - A keff t / V ] + [O]e .... [4]

where [O]i and [O]e are the initial and equilibrium oxygen concentrations, Keff is an effective mass transportcoefficient related to surface area, A, and V is steel volume. Schwerdtfeger has shown that keff is a functionof stirring intensity and that effective mass transfer coefficients from inclusion removal studies are similar tothose seen in ladle desulfurization and aluminum oxidation.

Stirring of any kind in a ladle will increase fluid velocities, decrease boundary layer thickness and promoteimproved efficiency of inclusion removal (as long as no other phenomena occur which promote inclusiongeneration). Mass transfer coefficients are often plotted as a function of energy input via stirring andcorrelations of inclusion removal trends can be developed in this manner; however, such an approach is onlyuseful when small, closely sized particles (such as alumina) are being removed. Fortunately, this approach isappropriate for ladle metallurgy.

Mixing in the liquid also has a strong effect on inclusion size and turbulence can result in more frequentcontact between particles and inclusion agglomeration. The level of turbulence necessary to optimizeinclusion agglomeration is currently an active field of research; however, as yet, no clear method of


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optimizing turbulence has been developed22,23]. Inclusion agglomeration also occurs on bubble surfaces andon refractory surfaces and small bubbles with float out velocities lower than that of the liquid steelrecirculating velocities are excellent sites for inclusion agglomeration, as is seen in continuous castproduct[22] and there have been many attempts to use fine bubble generation as a mechanism of inclusionremoval[24]. Another important method of inclusion agglomeration is on refractory surfaces, the mostobvious example being nozzle clogging. Agglomeration on refractories can be particularly troublesome as theclog is easily removed from the nozzle surface, freeing quite large agglomerated inclusions into the casting.Filtration of steel, an method of increasing inclusional separation to refractories, has been studied in detail forliquid steels over the last 10 years[25-28]; however, as yet, no long life filter or method of in-situ filterreplacement has been developed.

3. Separation or Stabilization at an Interface

The next step in clean steel manufacture is the separation of the inclusion from liquid steel to an interface. Itcan be shown from thermodynamics[29-32] that all inclusions have a lower energy when separated from theliquid steel to either a liquid steel-slag, liquid steel-gas or liquid steel refractory surface. Thus if the inclusioncan separate to a surface it will be stabilized on the surface.

In order that an inclusion separates to an interface the liquid between the inclusion and the interface mustdrain and then, a hole must spontaneously form and grow between the two interfaces in order for the particleto complete separation at the interface. The energy of hole formation is related to the interfacial energybetween the interfaces and the liquid steel and the distance between the particle and the interface. Clearlyenergy must be supplied to create the hole which when created will spontaneously increase in size as the totalinterfacial area decreases with adsorption of the particle or droplet.

This last step in inclusion transport can result in inclusions, which arrive at the interface but do not havesufficient energy to overcome the interfacial energy separating the two liquids, exhibiting a rest timephenomenon where the particles or droplets are stabilized for significant times before separation. In flowingsystems, this can lead to droplets or particles which reach the surface moving across the interface due to thevelocity gradients across the boundary layer and eventually being re-entrained. Thus in particle or dropletswhich exhibit rest time phenomena, the separation efficiency at an interface is less than unity and dropletsand particles under this condition can be extremely difficult to remove completely from liquid steel. Thus it isimportant to increase particle or droplet size in liquid steels to the point that the buoyancy plus inertial forceof the particle can overcome the interfacial forces to ensure complete separation at the interface. Thus fluidflow at the interface is also important in inclusion removal.

Separation of inclusions to a refractory interface can lead to inclusion stabilization at an interface; however, itis possible for inclusions stabilized at refractories to be released back into the fluid during times of fluidturbulence. In filtration or during nozzle clogging the refractory interface acts as a accumulator andagglomerator of inclusions and the presence of large clusters of inclusions in cast product can often be relatedto disintegration of inclusion build-ups stabilized against refractories.

Separation of inclusions to a bubble interface can be either helpful or deleterious to steel quality. The bubblescan cause inclusion stabilization and agglomeration on the bubble surface. If the bubble buoyancy causes thebubble and its associated bubble raft to separate to a slag-metal interface then the inclusions can becompletely removed from the steel; however, if the bubble size is too small top overcome the naturalconvection currents during process the bubble itself becomes an inclusion agglomeration site within the liquidand can also be responsible for quite large inclusion rafts that appear in castings. Thus separation to aninterface without complete removal from the system, can be deleterious to steel quality.

4. Removal From an Interface


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Complete separation inclusions from liquid steel includes removal from the interface. There is one majortechnique that is used to completely remove inclusions and that is dissolution into a liquid slag. For liquidinclusions this step is not a problem as most liquid are fully miscible in other like liquid, i.e., liquid slaginclusions tend to be completely miscible in the covering slags found in the ladle tundish and mold. Solidinclusions; however, tend to have limited solubility in the covering slags and the production of clean steel,where the inclusions are solid, is dependent upon slag chemistry, mixing in the slag, slag temperature gradientand volume.

To date most models of inclusion removal have assumed that inclusion transport to an interface is the ratedetermining step and that separation to the interface and removal from the interface are trivial; however,although inclusion separation to and emersion from the interface is thermodynamically favored transport intothe slag phase is diffusion controlled and strongly dependent upon slag chemistry. Thus, in a manner similar tosolute partition into a slag, inclusion separation can be controlled in a mixed mode where diffusion in the slaglayer and the solubility of the dissolving particle may become important.

If one views this process kinetically, ideally the flux of material away from the slag/metal interface should begreater than the flux of solid inclusions to the slag metal interface so that the interfacial concentration is lessthan the saturation limit for the slag. Once saturation concentrations in the slag are reached, inclusion will pileup at the interface, agglomerate and then dissolve at a rate determined by diffusion and the amount of mixingin the slag. Temperature gradients in the slag are also important as solubility decreases with temperature andthe total capacity of the slag to dissolve inclusions and remain liquid is defined by slag chemistry andtemperature.

References

1. L. S. Darken and R. W. Gurry: Physical Chemistry of Metals, McGraw-Hill, 1953

2.. N. A. Gockcen: Trans AIME, 206, 1558 (1956)

3. A. Muan: Am. Ceram. Soc. Bull., 37, [2], 81, (1958)

4. R. J. Fruehan and E. T. Turkdogan: "Physical Chemistry of Iron and Steelmaking" , Making Shaping andTreating of Steel, USS, 1984.

5.. C. Itoh:" Thermodynamics of the formation of Magnesium Aluminate Spinel in Liquid Iron", PhD Thesis,Tohoku University, Japan 1996

6. C. Benedicks and Helge Lofquist: Non -Metallic Inclusions in Iron and Steel, John Wiley, 1931.

7. A. W. Cramb and I. Jimbo: "Interfacial Considerations in Continuous Casting", Iron and Steelmaking, Vol.16, No.6, 1989, p 43 - 55.

8. M. Byrne, A. W. Cramb and T. W. Fenicle: "The Sources of Exogenous Inclusions in Continuous Cast,Aluminum Killed Steels," Iron and Steelmaker, June 1988, Vol. 15, p 41 - 50.

9. A. W. Cramb and M. Byrne: "Tundish Slag Entrainment at Bethlehem's Burns Harbor Slab Caster,"Transactions of ISS, Vol 10, (1989), p 121 - 128.

10. M. Byrne and A. W. Cramb: "Operating Experiences with Large Tundishes", Transactions of ISS, Vol 10,(1989), p 91 - 100.

11. A. W. Cramb: "Directions in the Production of Clean Steels", Trans AFS, (1994) , p 3 - 9.


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12. W. H. Emling, T. A. Waugman, S. L. Feldbauer and A. W. Cramb:"Subsurface Mold Slag Entrainment inUltra Low Carbon Steels", Steelmaking Conference 1994, p 371 - 379.

13. K. Wasai and K. Mukai: J. Japan Ist. Metals, 52, (1988), p 1088

14. K. Wasai, A. Miyanaga and K. Mukai: " Observation of Non-Metallic Inclusion in Aluminum-deoxidizedIron Alloy", Proceedings of the Joint US-Japan Seminar - Clean Steel for the 21 st Century", ed. Y. Iguchi,Tohoku University, 1996, p 87-90.

15. K. Schwerdtfeger: "Rates of Movement of Solid Particles, Drops and Bubbles in Static Liquids", Kineticsof Metallurgical Processes in Steelmaking, Verlag StahlEisen, 1975, p 192 - 218.

16. R. K. Iyenger and W. O. Philbrook: "Motion of Droplets and Inclusions in Liquid Steel", ibid, p 219 - 233

17. J. Szekely and V. Stanek: Met. Trans., 1, 1970, p 119.

18. S. Joo, R. I. L. Guthrie and C. J. Dobson: "Modelling of Heat Transfer, Fluid Flow and InclusionTransportation in Tundishes", Steelmaking Conference Procedings, 1989, p 401 - 408.

19. E. T. Turkdogan: "Ladle Deoxidation, Desulfurization and Inclusions in Steel - Part 1 : Fundamentals"Archiv fur Das Eisenhuttenwessen, 1983. No. 1., Vol 54, p 1-10.

20. E. T. Turkdogan: "Ladle Deoxidation, Desulfurization and Inclusions in Steel - Part 2 : Observations inPractice" Archiv fur das Eisenhuttenwessen, 1983, No. 2, Vol 54, p 45 - 52.

21. K. Schwerdtfeger: "Present State of Oxygen Control in Aluminum Deoxidized Steel", Archiv fur dasEisenhuttenwessen, 1983, No. 3, Vol 54, P 87 - 98.

22. S. Tanginuchi and A. Kikuchi: Tetsu-to-Hagane, 78, (1992), p 527

23. S. Tanginuchi: "Kinetics of Inclusion Agglomeration in Liquid Steel, Tetsu-to-Hagane,1996, p 81 - 111.

24. Y. Kikuchi, H. Matsuno, S. Maeda, M. Komatsu, M. Arai, K. Watanabe and H. Nakanishi: Proc. 8thJapan-Germany Seminar, (Oct 1993, Sendai, ISIJ, Tokyo), p 66.

25. D. Apelian and R. Mutharasan: J. Metals, 32, 1980, p 14-18.

26. D. Apelian, S. Luk, T. Piccone, R. Mutharasan: Steelmaking Conference Proceedings, Vol 69, 1986,p957-968.

27. P. F. Weiser, Steelmaking Conference Proceedings: ISS-AIME, Vol. 69, p969-976.

28. L. S. Aubrey, J. W. Brockmeyer and M. A. Mahaur: Steelmaking Conference Proceedings, ISS-AIME,Vol. 69, p977-991.

29. P. Kozakevitch and L. D. Lucas: Rev. Metallurg., 65, 1968, S. 589-98.

30. P. Kozakevitch and M. Olette: Revue de Metallurgie, October 1971, p 636 -646.

31. P. Kozakevitch and M. Olette: in "Production and Application of Clean Steels", 42, 1972, London, TheIron and Steel Institute.

32. P. V. Riboud and M. Olette: Proc. 7th ICVM, 1982, Tokyo, Japan, p 879 - 889.


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Introduction

Inclusion Stability


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Inclusion formation afs

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