Inclusion formation afs

Inclusion Formation - AFS http://neon.mems.cmu.edu/afs/afs2/form.html

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 can
be related to the presence of oxides created during the act of steelmaking, ladle metallurgy, casting and
rolling. This last definition causes extreme problems to the steel manufacturer as the definition of "clean" is
not 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 buy
steel 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 by
a customer and, as steel producers are continually striving to produce "cleaner" steels, the cleanliness
standard desired by the customer is continuously changing as a function of time and technological
improvements. The term "clean steel" is therefore continually variable depending upon the application and
the competition between steel suppliers.

Due to the variable nature of the term "clean steel" it is more accurate to talk about "high purity steels" as
steels with low levels of solutes and "low residual steels" as steels with low levels of impurities that originate
from scrap remelting and "clean steels" as those steels with a low frequency of product defects that can be
related to the presence of oxides. For example, there are "high purity, low residual clean steels" such as ultra
deep drawing steel sheets for automobiles which require ultra-low carbon contents (< 30 ppm), low nitrogen
contents (< 30 ppm) and the absence of oxide inclusions with diameters greater than 100 microns; and "low
residual 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 oxide
diameters 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 undergo
significant drawing are very susceptible to variations in inclusion size distribution, as cracks or wire breaks
can be easily counted. For example, in drawn and ironed cans, it is very common to count the number of
cracked flanges per million formed cans and generally a performance of less than 10 ppm would be looked
upon as an excellent performance. Another application which is amenable to measurement is bearing life. It is
well known that total inclusion content (as measured by total oxygen content) has traditionally correlated
with bearing life and decreased total oxygen contents (below 10 ppm) have lead to significant increases in
bearing life and thus the drive to very low inclusion contents in bearing steels. In addition to total oxygen
content, the total length of stringer inclusions after forging correlates well with bearing life and, at low total
oxygen 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 drive
to reduce solute and residual contents in all steels and to control the frequency and size distribution of the
inclusions that are found in all steels. Thus this chapter will focus upon an understanding of the fundamentals
of the production of high purity, low residual clean steels, emphasizing the limits in current technology and
the potential for the production of high purity, low residual steels with a low frequency of inclusions with
average diameters less than 5 microns.

Clean steels are steels with a low frequency of inclusions of average diameter less than 5 microns. The major
problems in clean steel manufacture are incomplete separation of clustered solid inclusions (> 5 microns in
diameter), 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 to
produce clean steel varies greatly between different steel plants; however, current clean steelmaking and
casting practices are based upon the following principles:

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

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

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

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

These practical principles of clean steel manufacture are based upon an understanding of the importance of
maintaining chemical equilibrium between the elements dissolved in liquid steel and the slag and refractory
systems which are in contact with the liquid steel, and, of controlling fluid flow to avoid conditions at liquid
slag-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 a
clean 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 execution
of these four issues.

1. Inclusion Generation

The formation of inclusions during steelmaking is inevitable as oxygen is more soluble in liquid iron than in
solid 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 in
contact with iron oxide at 1600 C is 0.23 wt%, a value which decreases with temperature according to the
equilibrium between iron, oxygen and iron oxide[1-2] The soluble oxygen content of pure liquid iron in
contact with a gas phase can be reduced if the gas phase oxygen partial pressure is less than 6 x 10-9 atm: a
partial 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 the
formation of iron oxide and generally this is accomplished by use of a liquid slag, with a low diffusivity of
oxygen, 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 is
solidified 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 removed
from liquid steel. For example, if it possible to transport the inclusions from the bulk of the liquid steel to a
free 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. Pure
iron solidifies at 1534 C where the maximum solubility of oxygen is approximately 0.17 wt/o. Thus, between
1600 C and 1534 C, 0.06 w/o soluble oxygen will transform to iron oxide. This is maximum possible amount
of oxygen which could removed under these conditions and, regardless of one's processing abilities, it would
be impossible to produce pure solid iron with less than 1700 ppm oxygen in the solid state, if liquid iron was
allowed 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 an
oxide, at liquid steel processing temperatures. Again, if liquid iron is used as an example, the equilibrium
between 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 of
the activity of FeO by dissolution into another chemical species. For example, liquid calcium aluminate could
be added to the liquid iron surface as a sink for the liquid iron oxide. As the iron oxide dissolved in calcium
aluminate, the mole fraction of iron oxide present in the newly formed slag would be reduced and the
equilibrium oxygen activity of the steel would also be reduced as the liquid iron attained equilibrium with the
iron oxide activity in the slag. Thus, modification of the activity of the deoxidation products the first method
that 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 the
equilibrium reaction. In steelmaking, carbon is always present, and the equation [2] often sets dissolved
oxygen activities.

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

In the BOF, for example, carbon content sets the lower limit for the oxygen activity at a given temperature
and as the CO pressure approximates to 1 atmosphere during refining, lower carbon contents lead to higher
oxygen activities and higher soluble oxygen levels. Oxygen content in carbon deoxidized steels can be further
modified by vacuum or inert gas treatment which will reduce the equilibrium carbon monoxide pressure and
lead 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 CO
are evolved during solidification of steels containing only carbon as a deoxidant. This "rimming" phenomena
was utilized in ingot casting to produce a cleaner ingot surface; however, the porosity associated with even
mild rimming is undesirable in many castings and the rimming action can be controlled by addition of
deoxidants to "kill" the action in the steel during solidification. Elements such as silicon, manganese and
aluminum are routinely added to liquid steel to reduce oxygen activities to below that which will cause CO
evolution 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 manganese
silicate; in aluminum killed steels, by the formation of solid alumina; and, in calcium treated, aluminum killed
steels by the formation of a liquid calcium aluminate. To further reduce soluble oxygen content at a given
temperature, manganese silicate or alumina can be dissolved into a slag. For example, dissolution of alumina
into a slag can result in the soluble oxygen content of an aluminum killed steel being reduced to less than 2
ppm. Dissolution of silica into a slag of calcium aluminate, for example, can significantly reduce the activity
of 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 be
effective, all phases in contact with the liquid steel must be at the same equilibrium condition. If aluminum
killed steels are at equilibrium, reaction [1] must also be at equilibrium; therefore, FeO activity in slags must
be very low or aluminum will oxidize by reaction with the slag. Similarly equilibrium MnO and SiO2 slag
activities in the slag are uniquely set by the addition of aluminum and any increases above the equilibrium
slag levels will result in reaction with the slag and the formation of alumina. Therefore, in practice, the slag
systems must be designed to be very low in FeO and MnO after the steel has been killed. Currently, furnace
slags 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 alumina
refractories are not 100% alumina and numerous binders are added which can contain oxides which are less
stable than alumina. Under these conditions the container refractories will react with the liquid steel, if not
adequately designed. Future clean steel developments will be aimed at developing steels which are in
equilibrium 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 refractories
has caused significant processing problems due to low residual levels (< 5ppm) causing the primary
deoxidation inclusion to be transformed into a high magnesia containing inclusion which often is a magnesium
aluminate spinel, a highly refractory inclusion. Often, this causes nozzle clogging during steel pouring. Recent
studies by Itoh et al.[5] have fully documented the problem and magnesium residuals must be reduced to less
than 1 ppm.

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

A second source of cleanliness problems is incomplete separation of emulsified slags[7]. This occurs during
processing events where there is a high dissipation of energy at the slag-metal interface. Events such as vessel
filling, vessel drainage and level fluctuations at the slag metal interface are generally responsible for the
generation of such inclusions. The size range of entrapped slag inclusions varies from less than 20 to greater
than 200 microns; however, their frequency can be low. Generally, these type of inclusions are practice
specific 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 well
sand 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 inclusion
formation. 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 refractory
system will enable the lowest possible inclusion contents in the shortest processing time. Reoxidation of liquid
steel 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 which
enables 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 can
allow the prediction of the chemistry of the most stable inclusion, it does not allow a prediction of the size of
an inclusion and, when inclusion removal is a goal, the inclusion size distribution is a key to understanding the
rate at which an inclusion can be transported to a surface where it has the opportunity to separate from liquid
steel.

Deoxidation and reoxidation inclusions must first be nucleated within the liquid steel and once nucleated
these inclusions will grow until the rate of growth is limited by diffusion in the liquid. The smallest measured
deoxidation inclusions are of the order of 15 nanometers[13,14] however, these quickly grow to between 1
and 5 microns in diameter in practical steelmaking conditions. Liquid turbulence can aid in agglomeration of
inclusions and solid inclusions can easily cluster and form large three dimensional rafts of sintered small
inclusions. Thus it is not unusual to find a high frequency of small inclusions less than 5 microns in diameter
and 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 5
microns and to eliminate all clustered inclusions before casting.

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

where VT is the sphere velocity, (( is the differential density between the particle and the liquid, ( is the liquid
dynamic 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 density
differences, in a static bath. Stokes Law only holds for rigid, spherical particles within the viscous flow
regime, where the Reynolds number is less than 0.1. For spherical alumina particles in liquid steel the
minimum particle radius for adherence to Stokes Law is 33.2 microns. For flows where the Reynolds number
is greater than 0.1, it is common to view the problem as a balance between the gravity force and the
combination of a buoyancy and a frictional force. Details of friction factors for different flow conditions and
particle types are given by Schwerdtfeger[15] and corrections for slip, interfacial tension, particle rigidity and
wall effects are given by Iyengar et al.[16].

Unfortunately, actual metallurgical vessels are not quiescent and are subject to natural convection, as has
been pointed out by Szekely[17] and Iyengar[16] and inclusion removal rates in industrial vessels cannot be
explained solely by Stokes' Law and its modifications. For example, Iyengar calculated that the critical
temperature gradient to avoid natural convection, for steel held in a small crucible, was 0.08 C/cm and that
fluid velocities of 0.015 cm/s were sufficient to ensure that the drag force due to the fluid flow was an order
of magnitude larger than the buoyancy force. In actual steelmaking practice, temperature gradients are high
against the refractory container walls and natural convection occurs leading to natural stirring patterns within
the ladle or the tundish. Liquid steel velocity calculations, by Joo et al.[18] indicate fluid velocities which are
greater 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 can
be viewed as a standard boundary layer problem where fluid flow leads to a momentum boundary layer which
has a thickness which can be related to the flow field and the physical properties of the fluid. Inclusions are
carried by fluid motion to a boundary layer at an interface where, due to buoyancy forces, they may traverse
this boundary layer. In order to calculate, from first principles, inclusion removal in a stirred system, it is
necessary to calculate the exact form of the boundary layer and then calculate the fraction of inclusions
which can traverse the boundary layer and separate during their contact time with the boundary layer. The
Navier Stokes equations must be solved for the coupled heat and mass transfer conditions to take into account
natural convection, to determine the boundary layer thickness against the container walls and the slag-metal
surface (taking into account the fact that this is a liquid-liquid boundary). The local flow conditions for each
particle at the boundary layer and the frequency of separation must then be determined as a function of
particle size. Calculations of this type are quite numerically complex; however, results have been given by
Joo et al.[18] where large inclusions with a large rising velocity separate more completely within the tundish
and the efficiency with which inclusions are removed decreases with decreases in inclusion size. Joo's work
indicated that even large inclusions can pass through a tundish and not be removed due to the effect of fluid
flow and that particles with diameters less than 40 microns ( float out velocities of 0.5 mm/s) have removal
efficiencies of less than 50%.

Inclusion removal in ladles has not, as yet, been computer modeled for inclusion removal in the same manner
as 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 practical
results:

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 apparent
flotation rate constant for a given type and intensity of stirring. Turkdogan models inclusion removal is a first
order process where the removal rate is proportional to concentration. Schwerdtfeger[21] has also analyzed
plant results of inclusion removal from aluminum-killed steels and found that a similar first order equation
could 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 transport
coefficient related to surface area, A, and V is steel volume. Schwerdtfeger has shown that keff is a function
of stirring intensity and that effective mass transfer coefficients from inclusion removal studies are similar to
those seen in ladle desulfurization and aluminum oxidation.

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

Mixing in the liquid also has a strong effect on inclusion size and turbulence can result in more frequent
contact between particles and inclusion agglomeration. The level of turbulence necessary to optimize
inclusion 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 and
on refractory surfaces and small bubbles with float out velocities lower than that of the liquid steel
recirculating velocities are excellent sites for inclusion agglomeration, as is seen in continuous cast
product[22] and there have been many attempts to use fine bubble generation as a mechanism of inclusion
removal[24]. Another important method of inclusion agglomeration is on refractory surfaces, the most
obvious example being nozzle clogging. Agglomeration on refractories can be particularly troublesome as the
clog 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 for
liquid steels over the last 10 years[25-28]; however, as yet, no long life filter or method of in-situ filter
replacement 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. It
can be shown from thermodynamics[29-32] that all inclusions have a lower energy when separated from the
liquid steel to either a liquid steel-slag, liquid steel-gas or liquid steel refractory surface. Thus if the inclusion
can 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 must
drain and then, a hole must spontaneously form and grow between the two interfaces in order for the particle
to complete separation at the interface. The energy of hole formation is related to the interfacial energy
between the interfaces and the liquid steel and the distance between the particle and the interface. Clearly
energy must be supplied to create the hole which when created will spontaneously increase in size as the total
interfacial 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 have
sufficient energy to overcome the interfacial energy separating the two liquids, exhibiting a rest time
phenomenon where the particles or droplets are stabilized for significant times before separation. In flowing
systems, this can lead to droplets or particles which reach the surface moving across the interface due to the
velocity gradients across the boundary layer and eventually being re-entrained. Thus in particle or droplets
which exhibit rest time phenomena, the separation efficiency at an interface is less than unity and droplets
and particles under this condition can be extremely difficult to remove completely from liquid steel. Thus it is
important to increase particle or droplet size in liquid steels to the point that the buoyancy plus inertial force
of the particle can overcome the interfacial forces to ensure complete separation at the interface. Thus fluid
flow 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, it
is possible for inclusions stabilized at refractories to be released back into the fluid during times of fluid
turbulence. In filtration or during nozzle clogging the refractory interface acts as a accumulator and
agglomerator of inclusions and the presence of large clusters of inclusions in cast product can often be related
to 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 bubbles
can cause inclusion stabilization and agglomeration on the bubble surface. If the bubble buoyancy causes the
bubble and its associated bubble raft to separate to a slag-metal interface then the inclusions can be
completely removed from the steel; however, if the bubble size is too small top overcome the natural
convection currents during process the bubble itself becomes an inclusion agglomeration site within the liquid
and can also be responsible for quite large inclusion rafts that appear in castings. Thus separation to an
interface 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 major
technique that is used to completely remove inclusions and that is dissolution into a liquid slag. For liquid
inclusions this step is not a problem as most liquid are fully miscible in other like liquid, i.e., liquid slag
inclusions tend to be completely miscible in the covering slags found in the ladle tundish and mold. Solid
inclusions; 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 gradient
and volume.

To date most models of inclusion removal have assumed that inclusion transport to an interface is the rate
determining 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 into
the slag phase is diffusion controlled and strongly dependent upon slag chemistry. Thus, in a manner similar to
solute partition into a slag, inclusion separation can be controlled in a mixed mode where diffusion in the slag
layer 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 be
greater than the flux of solid inclusions to the slag metal interface so that the interfacial concentration is less
than the saturation limit for the slag. Once saturation concentrations in the slag are reached, inclusion will pile
up at the interface, agglomerate and then dissolve at a rate determined by diffusion and the amount of mixing
in the slag. Temperature gradients in the slag are also important as solubility decreases with temperature and
the total capacity of the slag to dissolve inclusions and remain liquid is defined by slag chemistry and
temperature.

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 and
Treating 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 in
Ultra 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-deoxidized
Iron 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", Kinetics
of 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 Inclusion
Transportation 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 in
Practice" 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 das
Eisenhuttenwessen, 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. 8th
Japan-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, The
Iron and Steel Institute.

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

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Front Page

Introduction

Inclusion Stability

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Inclusion formation afs

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