Identification of Virulent Isolates of the Entomopathogenic Fungus Nomuraea Rileyi (F) Samson for...

INCLUSION ENGINEERING AND THE METALLURGY OF CALCIUM TREATMENT

Sunday Abraham, Rick Bodnar and Justin Raines

SSAB Americas 1755 Bill Sharp Boulevard

Muscatine, fA 52761 e-mail: [email protected]

[email protected] Jus tin. [email protected]

Tel.: (563) 381-5774

Keywords: Non-metallic inclusions, cleanliness, calcium treatment, inclusion modification, inclusion shape control.

ABSTRACT

Although a great deal of research has been done in the area of inclusion control during steelmaking, a complete description of the

process of inclusion modification and shape control cannot be found in the literature. As a result, the term "inclusion modification and

shape control" is often used loosely in the technical community. To utilize the full potential offered by a calcium treatment, it is important to understand why an inclusion whose composition has been modified may not necessarily have its shape modified. With

the purpose of decoupling the process of inclusion phase change from inclusion shape control, and to provide a complete description

of the metallurgy of calcium treatment, the authors have performed a detailed literature survey on the subject. By coupling t he survey

with the concept developed by the authors to tailor the cleanliness of steels to the intended applications, the hope is that this paper will

provide the reader with valuable information on calcium treatment for inclusion modification and shape control.

INTRODUCTION

The term "clean steel" is used with caution in the metallurgical community. This is due to: a) the varying cleanliness demands for

steels for different applications; b) varying cleanliness in steels produced in different operations and c) our understanding of the word

"clean steel", which some literally interpret as meaning the absence of inclusions in the steel. Steel cleanliness has implications both

from operational and product perfonnance points of view. There is an abundance of infonnation in the literature on the effect of

inclusions on product performance and on the kinetic and thermodynamic phenomena associated with inclusion evolution and

fonnation1.22 With careful analysis of the literature it is possible to develop a good practice at each stage of the steelmaking process

for clean steel production. However, it is not possible or even necessary to eliminate all inclusions, as certain inclusions which are

detrimental to steels for one application may entirely be hannless when present for another. Therefore, steels are expected to have

varying degrees of cleanliness depending on their application. The approach for reducing the harmful effect of inclusions is to tailor

the steelmaking process to avoid the presence of macroinclusions while controlling the population, size, distribution and morphology

of the residual microinclusions in the steel. The application of new technology and the knowledge gained from end users on the

performance of products are valuable infonnation for use in the design of a clean steel strategy. The science of inclusion modification

and shape control stem from the need to change the chemistry of the inclusions to enhance the performance of products in the field

and ensure castability during continuous casting. However, there is a general misconception among the end users of steels about the

impact of calcium treatment on steel cleanliness and the increasing demand to calcium treat steels is bothersome to steelmakers as

calcium treating certain products may not necessarily improve the cleanliness or the performance of the steel. The theme of this paper

is to discuss inclusion modification and shape control as they apply to oxides and sulfides. We will discuss the origin of oxide and

sulfide inclusions, factors affecting inclusion floatation, the nature of steel matrix-inclusion interaction, the use of calcium for

inclusion modification and shape control, and guidelines for calcium treatment.

INCLUSIONS IN STEEL

Non-metallic inclusions can be classified broadly into exogenous and indigenous types. Exogenous inclusions are derived from

external sources like ladle/tundish slags, refractory and re-oxidation products. Exogenous inclusions are always process-related and

consequently, can be eliminated by implementing suitable processing procedures23

The indigenous inclusions result from additives to

the steel. The indigenous inclusions are naturally occurring, and therefore, can only be minimized and not completely eliminated.

1243AISTech 2013 Proceedings

As far as inclusion modification and shape control are concerned, the inclusions of interest are the indigenous type, particularly, the

inclusions resulting from the deoxidation process and sulfide-type inclusions. Oxides and sulfides are the two predominant inclusions

in steel. The sources of oxides and sulfides are inherent to the steelmaking process. Oxygen is employed to react with natural gas and

carbon to generate chemical energy for the melting process. However, a significant amount of the oxygen ends up being dissolved in

the liquid steel. The dissolved oxygen must be removed during the refining stage because of its harmful effect on the structural

integrity of the finished product. Strong deoxidants, like aluminum and silicon, are commonly used in the steel industry to scavenge

oxygen from the steel. However, aluminum-killed steels routinely clog tundish well nozzles and submerged entry nozzles during

continuous casting due to the residual alumina inclusions that remain in the steel. The onset of clogging starts when an alumina

inclusion attaches to the nozzle wall.

Certain types of refractories, especially the graphite-stabilized MgO refractories, have been reported to promote agglomeration of

alumina inclusions. The high contact angle between the alumina inclusions and the steel further promotes the tendency of the

inclusions to agglomerate on refractories. In addition, the presence of significant amounts of alumina and MnS inclusions negatively

impact the performance of steel products. In general, non-metallic inclusions can cause lamellar tearing and degrade the toughness,

bendability, and ductility of steels.

When aluminum is added to liquid steel for deoxidation, the aluminum reacts with the oxygen to form dendritic alumina inclusions

(alumina galaxy), as shown in Figure 1 24 The reaction between aluminum and dissolved oxygen in liquid steel proceeds according to

Equation 1 , and Figure 2 shows the dissolved oxygen as a function of the dissolved aluminum25 From this figure it is clear that there

is an optimal window for aluminum contents in steels in order to achieve the lowest oxygen content. The total oxygen level in steel is

a good indicator of the degree of deoxidation of steel. A total oxygen of 10 ppm is achievable during the refining process (Figures 3

and 4)26

( 1 )

,

•

. •

• • -. ,. •

. .

-- --.6}

Figure 1 : Alumina Clusters: a) Optical Micrograph of Alumina Clusters. Original magnification 1 500 times; b) Electron

Micrograph of an Alumina Cluster in a Sample of As-cast Steel Etched for 6 min in 5 pct Solution of Bromide in Methanol.

Original magnification 1 500 time24

Depending on size, the alumina inclusions formed as a result of deoxidation can be divided into macroinclusions and microinclusions .

Large alumina inclusions are generally buoyant, and as a result, they float readily from the steel into the slag phase. Smaller inclusions

that are not as buoyant as the large inclusions take a long time to float from the steel. As an example, 1 00 micron size alumina

inclusions will float out in 5 minutes while 20 micron size alumina will float out in 1 1 9 minutes 27 Figure 5 further illustrates the

relationship between non-metallic inclusion size and their ability to float. Here, the floatation ratio (the % of reduction in inclusion

population) over a period of time is plotted against the size of the inclusion26

Sulfur is introduced into the steel from coke, natural gas, oil, recycled materials and ore used in steelmaking. Of all the raw materials used in the modem steelmaking process, pig iron contributes most of the sulfur present in the steel. Once dissolved in the steel, sulfur

combines with several dissolved elements to form sulfides. The impact of sulfur content in the steel on its structural integrity depends

on the type of sulfide fonned. For example, FeS is the most deleterious of all sulfides because upon reheating of the steel, low melting

point eutectics ofFeS and FeO are fonned which leads to hot shortness.

1244 AISTech 2013 Proceedings

100

0.1 0.001

- mass ppm 0 _ slgworth e

- activity of 0 _ slgworth et

- - . mass ppm 0 _ Itoh et.a!.

- - . activity of 0 _Itoh et.a!.

0.01 mass ppm AI

Sigworth, et al. (ppm 0)

---.

Sigworth, et al. (0 activity)

0.1

. ,

S 0-� C '" c 0 u I: OJ OJ) » x 0 -;; �

35 O.18%C 30 Al-k

0 0

25

20

15

10 5

Base molten steel

_ Steel equivalent to SCM 420 (Stmctuml

o Steel- 420 MPa)

0 0

0 0 0

0 0 0 0

0 0 0

0 10 20 30

Elapsed time after alloy addition (min)

40

Figure 2: Isotherm of Deoxidation of Steel with Aluminum25 Figure 3 : Change in Oxygen Content of Steel During Al Addition26

� Vl z o

0·06

� 0·04 ...J u Z o z « � 0·02 <!) ,.. >< o

0··········

Inclusions in solid steel

Inclusions in \ iquid steel

[OJ total

:"'"--..... -. .-.-. :::-..:�"":,,.:-.

. .... . . .. .. . .... \�.-.-FURNACE CAST HELD IN THE LADLE TEEMED IN THE

1��1� SOLIDIFI-LADLE PROCESSING STAGES"""" CATION

Figure 4: Change of Oxygen and Inclusion Content

of Steel from Furnace to Ingot25

120

100

�80 0 '';::; (Q a: 1:60 0 '';::; (Q

....

�40 ;:;:

20

0 0 50 100 150 200 250 300

Inclusion Size (�m)

Figure 5: Non-metallic Inclusion Floatation as a Function

of its Size (Redrawn)26

Figure 6 provides the Gibbs free energy of formation of different sulfides28 This figure is a guideline to steelmakers on the choice of

elements to use in alloying the steel to suppress the formation of certain types of sulfides. Note that CaS is the most stable form of

sulfide.

In general, the onset of precipitation of MnS during casting for a given manganese content is a function of the sulfur content in steels

(Figure 7)29 According to this figure, the critical sulfur level for suppression of MnS for a steel containing 1 .0% Mn ranges from

about 12 to 40 ppm depending on the deoxidation practice employed. MnlS ratio of the steel is a better indicator for the onset of

manganese sulfide precipitation and based on Figure 7, it seems this ratio can be as high as 500. With lower Mn content in the steel, a

lower MnlS ratio can be tolerated so long as the sulfur content is low enough that formation ofFeS inclusions is precluded.

Much research has been perfonned in an attempt to understand factors affecting the removal rate of inclusions from the steel. The

removal of a given inclusion from the liquid steel into the slag involves several steps: a) transport of the inclusion to an interface; b)

separation of the inclusion to an interface and c) removal of the inclusion from the interface. Although, it is widely accepted that the

transport of inclusions to the interface is the rate-detennining step it is understood that each of the three steps could play a significant

role depending on the temperature, slag chemistry, inclusion chemistry, inclusion shape and intensity of bath stirring. The initial

studies of inclusion removal were focused on the floatation of inclusions due to buoyancy-driven flow for rigid spherical particles,

where Stokes' law applies30:


2/':.pgr2 ----9 J1

(2)

where VI is the sphere velocity, �p is the differential density between the particle and the liquid, 11 is the liquid dynamic viscosity, g is the acceleration due to gravity and r is the particle diameter. In general, the ability of an inclusion to float from the steel is

detennined by the magnitude of forces acting on the particle as depicted in Figure 8a.

200 eoo 8QO .. 1000 IZ 1 400 TEMPERATURE �C)

ISOC' 2000

Figure 6: Gibbs Free Energy of Fonnation of Sulfides28

� . .,. . - . (i) I '8

.' t

J

1 00� __ ��

80

0 . 20

.. . . ; .� ��.-.,.' -' . Mn-TI

.!..-;.-; .•... :.-� ... : . .. , ......... , .... _ ..... , •• __ .a

./. o· ,'t . :Mn-AI' . ,. ��. -. J.:: i i i i

. '.'!

.'--/f-" -' �

..... i / . " Mn-Zr; ! • Wakoh �t al.'''' .

.} 0 Wakoh et al. (171

:' i • .. Wakoh at al. ("I , ,: , :l

,./' if'

« . W�koh et al.ll1( '. • 10 kg ingot Wakoh at al. ('�

,*Uesh'ima at 01."'(

low caroonsteeJ (0.1% C, 1.0%Mn) 01o.i.,. ....... _ ...... 000!0 ...... _ ..... ____ ""+_� 10 ' 20 50 100 2QO 400 800

S, ppm

Figure 7: Precipitation Ratio ofMnS as a Function of % S29

The driving force for the agglomeration of small spherical particles is larger than the turbulence forces of the melts; and therefore, the

particle clusters will float to the melt surface. Since Stoke's law describes the inclusion floatation mechanism, it can be said that this

law deals only with the aspect of an inclusion transport to an interface. The transport of an inclusion from the metal/slag interface into

the slag phase is a function of the slag temperature, inclusion chemistry and surface tension between the inclusion and the melt.

Lindon, et a131, 32 investigated these factors based on experiments with different alloy additions to an Fe melt. Samples of the melt

were periodically taken for inclusion analysis. The time at which 90% of the deoxidation products have separated from the steel were

detennined. The results are shown in Table 1 . From Lindon's experiment, it was concluded that the separation speed of deoxidation

products in quiet melts increases with increasing surface tensions of the particles. For example, the interfacial tensions between

alumina and an Fe melt is higher than that of silica and an Fe melt and, hence, an alumina inclusion will float out faster than a silica

inclusion. In general, the degree of wetting of an inclusion by liquid steel is a function of the angle � and the relative magnitudes of

(j MS (j sand (j M as shown in Figure 8b32 For a completely non-wetting inclusion, j3 is 1 800 Figure 9 shows metal-inclusion contact

angles for liquid and solid inclusions according to the data by Cramb, et al 33 and Misra, et al 34

llU"" Figure 8: a) Forces Acting on a Separating Particle

30 (where Fu,z is the interfacial capillary force, Fm is damping of acceleration due

to the fluid, Fd is the drag force and Fb is the buoyancy force) and b) Interfacial Tensions Acting on a Slag Drop in Contract with a

Metal Surface32 Where Fu,z is the interfacial capillary force, Fm is damping of acceleration due to the fluid, Fd is the drag force and

Fb is the buoyancy force.


The transport of an inclusion away from the interface is a function of the slag chemistry, solubility of the inclusion in the slag and

inclusion absorption capacity of the slag and slag temperature. Hence, a proper slag design is important for quick inclusion separation

from the interface and dissolution in the slag.

1 QJ c:: u 0 ... . -o .... ... � Ql) 1U c:: c. .- QJ

.2: VI ... c:: C 0 QI) .-c:: III

._ ::J III -IU u QJ .E ... ... u .E .E

0

I Liquid Inclusions I

laO-AI,03

J6/14)

CaO-AI,03 -SiO, • ecap-AI,03 (40/40/20) I (5�/50)

',O-AIP, -51°1 I (26/26/49)

• 30 60

Solid Inclusions

.1 •

liO,

90

CaO .AI,03 • • ·MgO I

MnO

120 Metal-Inclusion Contact Angle (0)

150

Figure 9: Metal - Inclusion Contact Angles (adapted from Cramb, et al 33 and Misra, et al 34)

Table 1 : Separating Rate of Deoxidation Products

Sample Composition of Deoxidizer Deoxidation products t90e;o' min I 100% AI alumina 8.0 2 100% AI alumina 8.0 3 100% AI alumina 2.3 4 100% AI alumina 2.5 5 100% Si silica 12.5 6 100% Si silica 12.8 7 90.2% Si, 3.8% al and 6% Ca CaO-SiO,-Al,03 18.0 8 90.2% Si, 3.8% al and 6% Ca CaO-SiO,-Al,03 18.0 9 85,9% Si, 4.1 % AI and 10% Ca CaO-SiO,-Al,03 20.5

10 87.8% Si, 3.9% al and 8.3% Ca CaO-SiO,-Al,03 17.5 II 64% Si, 6% AI and 30% Ca CaO-SiO,-Al,03 11.8 12 67% Si, 3% AI and 30% Ca CaO-SiO,-Al,03 11.0 13 92% Si, 3% AI and 5% Mg MgO-Si02-Al203 17.0

In addition to the slag condition, and inclusion chemical and physical properties, an adequate refining time and a good bath

homogenization are important for a complete removal of large inclusions from the bath. Microinclusions, on the other hand, can be

tolerated and may even be beneficial so long as they are present in the steel in a desirable fonn and are widely dispersed. The

detrimental effect of microinclusions on the structural integrity of finished product depends on whether they are "hard" or "soft"

relative to the steel matrix35 It has been reported that fatigue failure is a consequence of the structural tessellated stresses (stresses

resulting from radial compression and circumferential tension at inclusion sites) in and around the inclusions27 The tessellated stresses can be expressed as:

(3)

where <l> is a variable function depending on (a) the elastic moduli of the inclusions and the steel matrix, (b) the inclusion size, shape,

and distribution, and (c) the position and direction of the individual stress considered. The sign depends on the type of stress, e.g.,

positive for circumferential and negative for radial. The strain potential of structural tessellation is (am-aj)� T, where am is the mean

linear coefficient of thennal expansion of the matrix, aj is the mean linear coefficient of thennal expansion of the inclusion, and 1'17 is the temperature change in the system. Another way of looking at inclusion/steel matrix interaction is to consider the relative

elongation of the inclusion to the matrix. Malkiewicz and Rudnik36 suggested the following relationship as an index of defonnability

for inclusions:


&1 V= -

&2 b

£1 = In A, = In -

a

(4)

(5)

(6)

where £1 is the true elongation of the inclusion and £2 is the true elongation of the steel matrix, b is the inclusion final length, a is the

thickness of the inclusion, F 0 is the initial cross sectional area of steel and F 1 is the final cross sectional area of steel.

If V is 0, stress concentration at the inclusion/matrix interface is unavoidable and this may result in cracking at the interface. If V is

greater than 1 , stress concentration at inclusion/matrix interface is equally unavoidable, particularly during cold deformation. An ideal

condition is when V is equal to 1 (i.e. the inclusion and steel have the same plastic behavior).

Figures l Oa, b and c show examples of the interaction of particles with the steel matrix37

.38

.39

Non-metallic inclusions which set up

stresses in the matrix are primarily responsible for fatigue failures. The inclusions with lower coefficients of expansion than the steel

are the most deleterious. Overall, the deformability and thermal properties of the inclusion in reference to that of the steel matrix

detennines the nature of its effect on the structural integrity of the product (Figure 1 1 ) . For example, the behavior of "hard" inclusions

depends on whether they take the form of the globular high-silica, vitreous silicates and rare earth oxysulfides - Types ("a" and "b");

crystalline aggregates such as the high-alumina calcium aluminates and the crystalline silicates - Type (c); or alumina clusters - Type

(d). In each case, the matrix has to defonn around the inclusions, creating cavities or causing deformation damage which can act as

sites for fatigue crack initiation in engineering components and zones of internal weakness in processes or service conditions

involving tensile stresses. This type of inclusion which does not deform with the matrix is particularly hannful during wire-drawing

operations.

Stringer formation Types (c) and (d), increase the directionality of mechanical properties, adversely affecting toughness and ductility

in particular. The inclusions which have the most deleterious effect on toughness and ductility, particularly in the through-thickness

direction of flat-rolled products, are those which more or less deform with the matrix. As shown in Figure 1 1 40, inclusions of Type (e)

include manganese sulfides stiffened by a "skeleton" of oxide, and the partially crystallized silicates, while vitreous silicates and the

purer manganese sulfides are examples of Type (t). Defonnable oxides and sulfides can also cause lamellar tearing in steel weldments

where the fusion boundary is approximately parallel to the rolling plane of the plate. Most inclusion modification techniques are

designed to replace inclusions of Type (t) with those of Type (a), as inclusions of the latter type do not have such a harmful effect on

ductility and toughness. In such applications as wire-drawing, however, inclusions of Type (t) are preferred. For good machining

properties, macroinclusions of the sulfide type are preferable. Therefore, inclusion modification should be tailored to the application of

the product and desired final properties.

In summary, with good steelmaking practices, gross contamination of the steel by harmful macroinclusions can be avoided but the

presence of microinclusions can be tolerated in some applications provided they are small in size and widely dispersed within the

matrix. A successful inclusion engineering strategy requires a good knowledge of the relationship between inclusions and properties of

steel as well as how they interplay with the enviromnent under which the steel is being used. With this approach it becomes clear

which inclusion type requires modification to render it less harmful to the intended application of the product.

radial compcoS$lOn

..

G Metallurgy of Hie / sse 00------

::�::= :: Hz 10

• H2 accumulates a\ particle interfaces (notab� MnS, TIN)

Figure 1 0: Inclusion-steel Matrix Interaction: a) Schematic Diagram Showing Primary Tessellated Stresses37

; b) Cavities Created

around Inclusions and c) Promotion of Localized Shear and Void Coalescence by Flattened Type II MnS Inclusions38

,39


,s-cast After defonnation

a) A 'HARD' INCLUSION

UNDER ROLLING CONDITIONS •

(a) non-defonnable viscous inclusion

b) A 'HARD' INCLUSION UNDER ROLLING

CONDITIONS (b) beginning disintegration

c) A'HARD'CRYSTALLINE INCLUSION BROKEN

DURING ROLLING (c) crushing of a crystal

d) A 'HARD' INCLUSION ..... -CLUSTER 'STRUNG OUT'

DURING ROLLING (d) dislocation of agglomerates

e) AN INCLUSION

COMPOSED OF 'HARD' CRYSTALS DISPERSED

IN A 'SOFT MATRIX' (e) flow of a heterogeneous inclusion

t) A "SOFT" INLCUSION UNDER ROLLING

CONDITIONS o (f) flow of a liquid inclusion '------'-------- -.-.--

Figure 1 1 : Schematic Representation of Inclusion Morphologies Before and After Deformation40

CALCIUM TREATMENT FOR INCLUSION MODIFICATION AND SHAPE CONTROL

The process of reducing the harmful effect of microinclusions by controlling their size, shape, and properties is known as inclusion

modification27 A common approach to modifYing oxide and sulfide inclusions to prevent clogging and minimize any negative effects

on the structural integrity of steel is through calcium injection during secondary refining of the steel.

Rare earth metals like cerium, lanthanum etc., have also been used by steel makers to modifY inclusions, but they are not as efficient as

calcium due to the slow floatation (due to their weight) of the modified inclusions35 In addition, lanthanum and cerium readily

corrode the ladle refractories. Calcium treatment is performed to achieve two primary objectives27:

1 ) During calcium treatment, the alumina and silica inclusions are converted to molten calcium aluminate and calcium silicate

which are globular in shape because of a surface tension effect. This change in inclusion composition and shape is known as

inclusion morphology control27

2) The calcium aluminate inclusions retained in liquid steel suppress the fonnation of MnS stringers during solidification of

steel. This change in the composition and mode of the precipitation of sulfide inclusions during solidification of steel is

known as sulfide morphology or sulfide shape control27

The conversion of inclusions to globular shape has been theorized to play a significant role on the separation rate of inclus ions41. It

was observed in Lindon's experiment that Ab03 inclusions are non-wetting in liquid steel and tend to have a higher separation rate

compared to CaO-Si02-Ab03, for example31,3 2

This implies that by modifYing the alumina inclusions with calcium, their ability to

cluster is impeded as the liquid globular inclusions formed and as a result are wetted by the liquid steel. However, the authors believe

that the high vapor pressure of calcium with the associated intense bath stirring promotes collision and coalescence of the alumina

inclusions in the melt. With the aid of calcium vapor and the resulting coalescence of the alumina inclusions through collision, their

removal from the steel is enhanced compared to the small non-buoyant alumina inclusions which must first cluster on their own

(without forced convection) before they are able to separate from the liquid steel. This is why umnodified small alumina inclusions

will only separate from the liquid steel and get attached to the refractory in the tundish well after refining is complete in the ladle.

The ultimate goal of calcium injection is to reduce the ability of the small harmless solid oxide and oxysulfide inclusions to clog

tundish nozzle to ensure castability and reduce their harmful effects on the perfonnance of the steels by changing their composition


and physical properties. In general, calcium - bearing agents (CaSi, CaFe, CaAl, CaC, etc.) are usually introduced into the steel at the

end of the steel refining in the fonn of powder or wire injection through hollow metallic tubes. Irrespective of the Ca-bearing agent

employed, the quantity of calcium required for treatment in a given weight of steel depends on the alumina content, and the oxygen

and sulfur levels of the steel. A sufficient amount of calcium must be added to react with the alumina inclusions to fonn calcium

aluminate compounds that are liquid at steelmaking temperatures. For completely modified inclusions, the equilibrium reactions are as

follows42,43

:

rCa] +[0] = (CaO) (7)

[Ca]+[S] = (CaS) (8)

7(AI,03) + 12[Ca] + 12[0] = 12CaO· 7 AI,03 (9)

[MnS] + 2[0] + CaSi = (CaS) + (SiOJ + [Mn] ( 10)

Figure 1 2 shows the binary phase diagram of CaO-Ab03 . The highlighted region in the figure shows the desirable composition of the

calcium aluminate inclusions. Outside the highlighted region the phases are solid at steelmaking temperatures. These phases can be the

prominent constituents when there is an over- or under-injection of Ca. While MnS inclusions are undesirable in the steel, the

fonnation of solid CaS inclusions is equally undesirable. In terms of clogging, solid calcium aluminate or pure CaS inclusions are just

as detrimental as the alumina inclusions; they also sinter and agglomerate on nozzle refractories. Agglomeration of alumina, calcium

aluminate and CaS inclusions on tundish nozzle refractories during continuous casting can result in a premature termination of casting

due to a completely clogged nozzle. Depending on the population of the inclusions in the steel, complete clogging of the nozzle can

occur within minutes of the start of casting.

u 2-UJ no ::> !;:r '" � :::'" u.J I-

�600 2570 2500 "-

'" 2400 "-2300 �

'\ 2200 '\ 2100 \ 2000

1900 LI ME • �IOUID 1800

1700

1600

1(;00

,1400

1)00

1200

CoO

\ \

\ \

\ \ \

� "IQ 3CoO'AI;>03

Figure 12 : CaO - A1 P3 Phase

Inclusions are Highlighted.

70 eo AIP3 c.:.O· AI.,o3 J CoO'6A1:tJ3

Diagram27 Desirable Range for the

An example of a clogged nozzle is depicted in Figure 1 3 . Two phases of inclusions were identified in the nozzle: inclusions of the

pure alumina family and calcium aluminate rich in alumina with a melting point of about 1 850°C. Therefore, these inclusions are solid

at steelmaking temperatures causing them to tend to adhere to the tundish nozzle refractory. D. Janke, et a1 44 found a relationship

between total oxygen and castability in aluminum-killed steel (Figure 14) . From this figure, it is clear that the lower the total oxygen

content of the steel the greater the castability index. However, 20 ppm total oxygen is optimal to maintain a good CaO to Ab03 ratio

and for a reasonable castability index.


Figure 1 3 : Clogged Tundish Nozzle During Continuous

Casting (The buildup in the Nozzle shown by the arrow

comprises Ah03 and CaO.2Ah03).

� 0 'iii :J '0 .<;; '0 0 � 'r 6 ' .. � 0 01 (J

0,5 50 1.3 ;;\! 1: .QI 0,4 - 41 40 � �aOI�03 - 12

· vi )( Q) '" '0 0 0,3 'iO 30 .S :J ,� �. 11 II

0,2 - '0 20 � c' � Q) 10 C 0,1 � 10

� 0,0 (,) 0 9

10 15 4Il 2S 35

. totaJ.l0I-eo�te.nt, ppm

Figure 14 : Relationship between Total Oxygen and Castabilitl6

The efficiency of calcium treatment is dependent on a number of factors, which include the type, the amount and the injection rate of

the calcium-bearing agent used for the treatment. Overall, by classifying the alumina and MnS inclusions according to their

compositions and shapes, the efficiency of calcium treatment can be evaluated45

• Class A non-metallic inclusions are present when high levels of calcium have been added to the molten steel and are liquid

throughout processing (see Figure 1 5) . The intermingled sulfide and aluminate phases of these inclusions indicate that both

phases solidified at about the same time. The sulfide phase tends to be a CaS composition. The calcium aluminate phase is

either CaO· AI,03 or 12CaO· 7 AI,03• This indicates the presence of calcium aluminates with the lowest melting points and

with high levels of calcium.

• Class B non-metallic inclusions are the "bulls-eye" type most prevalent in calcium-treated steels. The central, dark aluminate

phase has solidified first and then the outer sulfide phase precipitated onto it. In this instance the sulfide phase tends to be a

(Ca, Mn)S. The calcium-aluminate is of the CaO· AI,03 or CaO· 2Al203 composition.

• Class C non-metallic inclusions are indicative of incomplete calcium treatment. As shown in Figure 14 these duplex

inclusions have an unmodified MnS phase which is deformable during hot rolling. The dark central calcium aluminate tends

to be of the CaO· 6AI,03 composition, which has the lowest calcium content and remains undeformed during hot rolling.

• Class D non-metallic inclusions are alumina-like oxide inclusion clusters which may have some calcium associated with

them. However, there is not enough calcium present to result in complete fluxing of the alumina galaxy.

• Class E non-metallic inclusions are MnS inclusions, which are present when sulfur has not been completely tied up by

calcium.

• Class F non-metallic inclusions are interdendritic MnS inclusions, which are present when sulfur is not completely tied up by

calcium and the oxygen potential of the steel is high.

The end results of an optimized calcium treatment are: (a) the alumina is modified to form liquid calcium aluminate, sulfur is tied up

as CaS, which will precipitate on the calcium aluminate inclusions, and (b) floatation of the inclusions is enhanced through the

fonnation and agglomeration of spherical oxide and sulfide inclusions.


Figure 1 5 : Inclusion Classes based on the Degree ofModification23

A good knowledge of the chemical and physical properties of calcium is required to ensure it is effectively used for inclusion

modification. Calcium has a low solubility in steel (-300 ppm), high vapor pressure (Pea = 1 .84 atm. at l 600°C), and reacts violently

with air due to its high affinity for oxygen. The free energies of fonnation for CaO and CaS are -203,000 and - 1 65,000 cal/mol,

respectively at 1 ,600°C, and thus, the CaO compound is thermodynamically more stable32

Therefore, at certain oxygen concentrations

it is unlikely that CaS can precipitate directly from the steel melt and then accumulate at the surface of CaO-Alz03 particles.

Therefore, given the higher affinity of calcium for oxygen compared to sulfur, care must be taken to avoid introduction of oxygen into

the bath during calcium treatment.

Many researchers have attempted to detenrune the required amount of calcium addition for optimal cleanliness results . For example,

Ca/S ratios have been correlated to reduction of area in the Z direction and impact properties of steel as shown in Figures 1 6 and 17 ,

respectively. An optimal Ca/S i s a function of the oxygen potential and sulfur level of the steel. At certain sulfur and oxygen levels

Ca/S becomes meaningless. A good knowledge of how total oxygen and sulfur contents of the steel affect inclusion modification is

necessary to avoid over injection of calcium, which will result in the formation of undesirable single-phase CaO 'Alz03-CaS

compounds. In contrast, under-injection of calcium will result in the fonnation of clusters of alumina inclusions in addition to MnS

inclusions. A good refining practice in the ladle and an efficient calcium treatment will result in the majority of the alumina inclusions

being converted to liquid calcium aluminate while most of the sulfur will be tied up as CaS. The CaS will precipitate on the calcium

aluminate to produce the desirable bulls-eye shape (Figure 1 8) .

---e 80 .---....:....----------------------.

'-o 20 c:: .S .... u .g �

o 00 gb 0 0 080.,00 • � • 0 •• � I <Y l! 0 08 : � ��\J�ellllllllllllll� � �� \ / �EndPart

�;J .. • Axial center • •

o 0.4 0.8 CatS ratio

Figure 16 : Relationship between Ca/S and Through

thickness Tensile Reduction in Area32

a

'8 � �

:>

1.0 .----------------,

O.S

0.6

2S

24

20

16

12

S

--0---«--0' -_.--

. __ :'_rr:+-L direction

C direction

4 2��-��--�--��

o 100 200 500 1 000 2 000 5 000 Total length of A type inclusions (x 400, 100 field)

Figure 1 7: Correlation between the MnS Inclusion Length, the

Ca/S Ratio and the Impact Properties of Steel Plate2 E - the

top of the ingot, M- the middle of the ingot, L-longitudinal to

rolling direction and C-transverse to rolling direction32


As cast

MnS segregated at grain boundaries

Rolled MnS stringer in the

roiling direction

Numerous broken angular crystals in the rolling direction

Ca wire treated

CaS-MnS ring lormed around C 12A7

D

Figure 1 8: Schematic Illustration of Modification ofInclusion Morphology with Calcium Treatment of Stee127

GUIDELINES FOR OPTIMIZING CALCIUM TREATMENT AND APPLICATION OF CRITICAL PRODUCTS BASED

ON CLEANLINESS PERFORMANCE

The authors have performed an extensive study of the effects of the type of Ca-bearing agent, amount of injection, rate of injection,

oxygen and sulfur levels of steel on steel cleanliness. The results are detailed in another paper entitled "Development of an Inclusion

Classification Methodology for Improving Steel Product Cleanliness" by S. Abraham, et aI., also published in the AISTech 20 1 3

conference proceedings. Interested readers are referred to this paper, but here a summary o f findings from the study i s given.

Type of Ca-bearing agent: Three different Ca-bearing agents were evaluated including 30% CaSi wire (or powder) injected through

a 1 3 mm metal tube. It was found that the 30% CaSi consistently produced better cleanliness results . The reason for the improved

performance of the 30% CaSi may be related to the beneficial effect of silicon on the nature of the reactions at slag/metal interface 46

Oxygen potential of steels: Prior to calcium treatment, the activity of oxygen in the melt should be reduced to a low level to avoid

unnecessary fonnation of CaO inclusions and to utilize the injected calcium effectively to modifY alumina and sulfide inclusions. A

total oxygen content in the range of 1 5 to 30 ppm can be achieved in the ladle by maintaining aluminum level within a certain range

and ensuring adequate refining in the ladle. Based on our evaluation, aluminum content in the range of 0.04 to 0.06% is considered

optimal.

Sulfur level: It is not very effective to inject calcium for sulfide modification at high sulfur levels. A significant amount of calcium is

required to achieve low sulfur and, in fact, such a reaction cannot proceed because excessive calcium injection will generate a high

vapor pressure, and, given the affinity of calcium for oxygen, a reaction with oxygen from the air is favored over a reaction with

sulfur. Such a scenario will result in poor steel cleanliness due to the formation of CaO and precipitation of sulfur as MnS inclusions

during solidification. For this reason, for critical products, where MnS inclusions need to be avoided, it is important to first reduce the

sulfur level to a low value by employing good slag practice prior to the calcium treatment.

It was found that the formation of gross MnS inclusions can be avoided at sulfur levels below 30 ppm (Figures 1 9a, b, c and d). At

such low sulfur levels, the recommended Ca/S is typically in the range of 8 to 10 for the best cleanliness results (Figure 20). For

applications where MnS inclusions are harmless, higher sulfur levels can be tolerated and the calcium injection amount should just be

sufficient to modifY alumina inclusions to avoid clogging the tundish well nozzle during casting. By carefully detennining the

injection amount and rate, Ca/S can be maintained consistently within a tight range and formation of a significant number of CaO can

be avoided.

CaSi injection amount: It was established through trials that for fully killed steels with a total oxygen content in the range of 1 5 to

30 ppm, the amount of CaSi required for complete inclusion modification can be written as a function of the volume of steel and its

sulfur level as in Equations 1 1 and 12 . These equations are useful up to a sulfur level of 30 ppm. At sulfur levels above this value,

modification of sulfide inclusions becomes difficult. Figure 2 1 shows the recommended CaSi injection amount as a function of %S .


1. 07 + %S

(meter per ton of steel) 19.69*10-4

0.38 + 182%S (kg per ton of steel)

( 1 1 )

( 12)

CaSi injection rate: As indicated above, given the violent nature of the injection process, CaSi injection should be performed in a

controlled manner to avoid re-introduction of air into the already killed steel or formation of undesirable CaO inclusions. Based on

cleanliness results and the nature of bath agitation, 30% CaSi injection rates of 1 .07 m and 0.38 kg per ton of steel are optimal for a

1 3 mm diameter core wire and powder, respectively.

Particle size: The efficiency of calcium treatments also depends on the particle size of CaSi. The required particle size range for CaSi

powder is defined in ASTM A46S47 However, to improve the efficiency of CaSi, it has been found that a tighter range for size distribution is required. Very fine and coarse particles are not desirable due to high and low rates of reactivity, respectively, resulting

from an effect of surface area. It was found that particle size in the range of 200 to 700 �m works best for calcium treatment.

Product disposition based on cleanliness performance: Good steelmaking practices for products destined for critical applications

should be coupled with clear cleanliness criteria. For such criteria, the distribution, interparticle spacing, proximity to the surface,

position of as-cast slab in a sequence of steel, type and size of the inclusions, and product thickness should be taken into consideration.

In addition, continuous inclusion engineering and re-engineering using current technology is key for improved product performance.

>...

0.10

. � 0.08 QI c � NE 0.06 .;: E QI-� � 0.04 1: I-� 0.02 :!!:

0.00

• a

•

0.0000 0.0010 0.0020 0.0030 0.0040 %S

0.012 C N E E 0.010 • '#" ;: 0.008 .� III I: � 0.006 "*' • � 0.004 QI

J: VI 0.002 I: :!!:

0.000 .----0.0000 0.0010 0.0020 0.0030 0.0040

%S

0.80 >-.� 0.70 I: � 0.60 QI� Ji N

E 0.50 .2 � 0.40 jl � 0.30 1:-� 0.20 � 0.10 :!!: 0.00

•

/ /

/ .

b

0.0000 0.0010 0.0020 0.0030 0.0040 %S

0.25 N d E • E E 0.20 -== >-

.� 0.15 III I: QI c QI 0.10 N Vi � • "' QI 0.05 J:

VI I: :!!: .-0.00 -

0.0000 0.0010 0.0020 0.0030 0.0040 %S

Figure 1 9 : a) MnS Thin Series # Density vs. % S; b) MnS Thin Series Size Density vs. % S; c) MnS Heavy Series # Density vs. % S

and d) MnS Heavy Series Size Density vs. % S .


QI) s::: 'u ro Co I/) QJ U � ro

600.00

500.00

400.00 Co ,-.. ... S .fl =. 300.00 .E -QJ "

t;:::: "5 � " Ox o

200.00

100.00

0.00

-- . -..... r • :--... � . . . . . ""-

.-- . . . . .

-II: • � . _ ; • .., .�

�!, . ...-' -.

-

o 5 10 15 20 Ca/S

25 30 35

Figure 20: Steel Cleanliness as a Function ofCa/S Ratio.

3 Qj QJ .: 2.5 o s::: o .:!:::. 2 .§. .... § 1.5 o E � 1 o � u °2 0.5

Vi � 0

o

For products for critical applications and to avoid

tundish well nozzle clogging

0.001 0.002 0.003 % S

CaSi injection amount - Sulfur line ICS line) for alumina modification

To avoid tundish well nozzle clogging

0.004 0.005 0.006

40

0.007

Figure 2 1 : Recommended CaSi Injection Amount for Inclusion Shape Control and Avoiding Tundish Well Nozzle Clogging During

Casting of a Completely Killed (AI-killed) Steel.

In summary, inclusions and product properties interact in a complex manner and systematic work is required to establish tangible

relationships. Good research work has been perfonned in different institutions and steel industry on inclusion fonnation and evolution

during steelmaking process. While it may not be necessary to re-invent the wheel, a careful assessment of the information in the

literature could provide guidelines with respect to inclusion engineering in steel products.

The authors hope the material presented in this paper will help clarify some of the key aspects of inclusion modification and shape

control, particularly to the end users of steels.

SUMMARY AND CONCLUSIONS

This paper provides a synopsis of the types of inclusions in steel, inclusion modification and sulfide shape control during refining in

the ladle. The mechanism of inclusion floatation in relation to their physical and chemical properties was discussed. An overview of

the nature and effects of steel matrix-inclusion interaction on product perfonnance was also provided. In addition, several key aspects

of the metallurgy of calcium treatment and inclusion evolution during refining in the ladle were discussed and summarized. Further,


guidelines for optimizing calcium treatment for clean steel production were provided. Overall, the key to an effective inclusion

modification is a controlled addition of calcium to steels during refining in the ladle.

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