Series, Vol 51, 1955, p 1-16 8. R.G. Olsson, V. Koump, and T.F. Perzak, Rate of Dissolution of Carbon Steel in Molten Iron-Carbon

Alloys, Trans. Met. Soc. AIME, Vol 233, 1965, p 1654-1657 9. R.D. Pehlke, P.D. Goodell, and R.W. Dunlap, Kinetics of Steel Dissolution in Molten Pig Iron, Trans. Met.

Soc. AIME, Vol 233, 1965, p 1420-1427 10. R.I.L. Guthrie and P. Stubbs, Kinetics of Scrap Melting in Baths of Molten Pig Iron, Can. Metall. Q., Vol

12, 1973, p 465-473 11. K. Mori and T. Sakuraya, Rate of Dissolution of Solid Iron in Carbon-Saturated Liquid Iron Alloys With

Evolution of CO, J. Iron Steel Inst. Japan, Vol 22, 1982, p 964-990 12. P.T.L. Brian and H.B. Hales, Effects of Transpiration and Changing Diameter on Heat and Mass Transfer to

Spheres, AIChE J., Vol 15, 1969, p 419-425

Purification of Metals

The structure and properties of cast metals are sensitive to numerous impurities. Because purification of melts generallyadds considerable cost to castings, the lowest cost and surest defense against contamination is careful selection of scrap.Purification is generally reserved for elements that are so pervasive that avoidance is impossible. This is exemplified bysulfur and oxygen removal from cast iron and steel, respectively, and removal of alkali and alkaline earth elements fromaluminum. Because of their immediate importance, aspects of the physical chemistry of these processes are reviewed.

Ferrous Melts

One of the most important processes involved in cast iron and steel production is desulfurization. For steels,desulfurization is necessary to reduce the level of inclusions, leading to stronger and more fatigue-resistant steels. For castiron, desulfurization is practiced in the manufacture of ductile iron castings in order to develop spherical graphitemorphology. Ductile iron is used in applications where high fracture toughness is needed.

Sulfur is removed from iron and steel when the metals are liquid. Although a variety of reagents are employed to removesulfur, namely, calcium, magnesium, sodium, and rare earths, the most important is calcium. Common forms of calciuminclude the metal; alloys such as calcium silicon (CaSi); the oxide, calcium oxide (CaO); and the carbide, calcium carbide(CaC2). Despite the use of various forms of calcium, the governing chemical reaction in all cases appears to be the same(Ref 13, 14):

CaO + S = CaS + O (Eq 9)

The equilibrium constant for the reaction (Eq 9) is:

9CaS o

CaO S

a hka h

= (Eq 10)

For both cast iron and steel, target sulfur concentrations after desulfurization are in the range of 0.006 to 0.010% S. Inelectric-melted cast iron and steel, the sulfur levels before desulfurization are 0.02 to 0.03% S, while input sulfur levels tothe cupola are generally much higher--0.1 to 0.2% S.

Requirements for Desulfurization. For reasons related to reaction kinetics and thermodynamics, the final sulfur(%S)f concentration achieved in desulfurization processes depends on three factors:

• Initial sulfur concentrations (%S)i• Amount of desulfurizer used, usually expressed as the weight fraction of desulfurizer to metal, W• Extraction efficiency of the desulfurizer, which is measured by the desulfurization ratio (DR), that is,

the ratio of sulfur concentrations: desulfurizer to metal

The three variables can be related as follows through a mass balance on sulfur:

(% )(% )1 ( )

if

SSW DR

=+ (Eq 11)

For liquid-desulfurizing slags that are not saturated with respect to calcium sulfide (CaS), the maximum value of DR, thatis, the equilibrium value, can be predicted from thermodynamic considerations:

14max

15max

(% )( )

(% )slag s s

f o

S c f kDRS k h

⎛ ⎞= =⎜ ⎟⎜ ⎟⎝ ⎠

(Eq 12)

where CS is the slag sulfide capacity, defined as (Ref 15):

2

2

1/ 2

(% ) Os slag

S

pC S

p⎛ ⎞

= ⎜ ⎟⎜ ⎟⎝ ⎠(Eq 13)

and K14 and K15 are the respective equilibrium constant for the reaction:

12

O2 = O (Eq 14)

12

S2 = S (Eq 15)

and fS is the activity coefficient for 1 wt% S in the standard state.

Equation 12 can be derived from Eq 10. Using known input and desired output sulfur values, Eq 11 gives the requireddesulfurization ratios as a function of weight fraction desulfurizer. These data are plotted in Fig. 10 for two cases. In theelectric-melting case, initial and final sulfur concentrations were assumed to be 0.03 and 0.008% S, respectively, Incupola-melting case, the equivalent concentrations were 0.10% S and 0.008% S. The cupola line applies for both cupolairon and ladle-desulfurized cupola iron.

Fig. 10 Weight fraction desulfurizer required to achieve given desulfurization ratios. Plots for electric-meltediron and steel (curve A) assume initial sulfur concentrations [%(S)i = 0.03] different from those of cupola-melted iron [(%S)i = 0.10] (curve B).

The relationships illustrated in Fig. 10 are useful in defining systems that will provide the necessary desulfurizationconditions. Four systems are compared in Table 3. These cover cupola- and electric-melted cast iron and electric-meltedsteel. Also examined are two liquid slag systems: a basic cupola slag (dicalcium silicate saturated) and a CaO-saturatedCaO-Al2O3 slag.

Table 3 Theoretical weight ratios: desulfurizer to iron to achieve 0.008% S in various systems

Type of melt Slag composition T, K CS (×104)(a)

fS hO (×104)(b)

(DR)max(c) W

A. cast iron--cupola 44 CaO-15MgO-5Al2O3-36SiO2

1773 2.7 5 1.3 90 0.13

B. cast iron--cupola (ladledesulfurized)

CaOsat-Al2O3 1773 59 5 5.0 417 0.027

C. cast iron--electric CaOsat-Al2O3 1773 59 5 5.0 417 0.0065

D. steel--electric CaOsat-Al2O3 1873 316 1 0.45 5280 0.00052

(a) CS is obtained from optical basicity correlations (Ref 15). fS is based on iron composition (Ref 16).

(b) Case A: hO is governed by Si/SiO2 equilibrium based on respective concentration in iron and slag (Ref 17). Cases B and C: hO is governed bySi/SiO2 equilibrium with aSiO2 = 1 due to ladle exposure to air (Ref 18). Case D: hO is governed by Al/Al2O3 equilibrium with % Al = 0.05(Ref 19), assumed no contact with air.

(c) K14 and K15 data from Ref 20

Table 3 shows that:

• The best desulfurizing cupola slags have lower sulfide capacity and desulfurization ratio than slags usedin ladle desulfurization; the consequence is the need for much larger quantities of slag*

• Compared to steel, cast iron desulfurization benefits from higher fS because of the presence of relativelyhigh concentrations of carbon and silicon in cast iron

• Ladle desulfurization systems that are exposed to air suffer higher hO and, as a result, poorerdesulfurization than might otherwise be anticipated

For the cupola slag case in Table 3, the thermodynamically predicted value of (DR)max = 90 is in good agreement withmeasured data (Fig. 11). This indicates that the cupola desulfurization process operates close to equilibrium levels.Further evidence for this is given Fig. 12, which plots desulfurization data for a cupola operated with varying amounts ofmunicipal ferrous refuse in the charge. The upper portion of Fig. 12 plots CS and oxygen activity. The latter is expressedas the partial pressure of oxygen. The lower portion of Fig. 12 is a comparison of (DR), measured and calculated. Thegood agreement found is evidence that near-equilibrium conditions existed.

Fig. 11 Desulfurization ratio-basicity comparisons for cupola (closed circles) and laboratory data (open circles).Equilibrium values are indicated by the angled line. The vertical line indicates the slag basicity above whichslags are saturated at 1500 °C (2730 °F), with respect to dicalcium silicate. This is the point at which theobserved DR should equal (DR)max. Source: Ref 14.

Fig. 12 Slag sulfide capacities, oxygen activities, and desulfurization ratios, measured and calculated, for acupola operated with municipal ferrous refuse as a charge material. Source: Ref 21.

For the cases discussed above, the oxygen activity in cupola iron has been found to be governed by the reaction (Ref 17,21):

Mn+ O= MnO (Eq 16)

Therefore, the overall cupola desulfurization reaction is:

CaO + S+ Mn= CaS + MnO (Eq 17)

Considerably lower sulfur could be achieved (Ref 14) if hO were governed by:

C+ O= CO (Eq 18)

However, equilibrium for this reaction has not been observed.

This discussion has concerned CaS-unsaturated slags. However, a desulfurizing slag, saturated with CaS, can in manycases continue to desulfurize as long CaO is present. In this case, the ultimate sulfur levels are not as low as those for

CaS-unsaturated slags. Nevertheless, CaS-saturated slags can produce sulfur levels that are more than sufficient for castiron and steel applications. In fact, CaS-saturated conditions are often employed in industrial applications because muchless desulfurizer is required.

In addition to the CaS-saturated liquid slags, totally solid slags such as CaO and CaC2 are in the same category becauseCaS does not form solid solutions with these materials. Also in this category are desulfurizers containing small amountsof liquid phase. These materials possess the desirable properties of liquid and solid slags. That is, they combine the fastreaction rates of liquid slags with the large desulfurizing capacities (high CaO concentration) of solid slags.

The final sulfur concentrations attainable under CaS-saturated conditions can be obtained from Eq 10 by setting aCaS = 1.For ladle desulfurization, the most important industrial desulfurizers are also CaO-saturated, that is, aCaO = 1. Applyingthe condition of double saturation to Eq 10 yields:

9

(% )s

hoSk f

= (Eq 19)

In the ladle desulfurization of cast iron, hO is governed by the reaction (Ref 14):

12

Si+ O= 12

SiO2 (Eq 20)

where aSiO2 = 1. This is attributed to the exposure of the iron to air (Ref 14). Sulfur concentrations obtained with Eq 19,assuming silicon-silicon dioxide (Si-SiO2) equilibrium and aSiO2 = 1, are given in Fig. 13. In continuous ladle-desulfurization processes, equilibrium sulfur levels are achieved when input sulfur levels are low, but are only approachedwhen input levels are high. This is illustrated in Fig. 14 with desulfurization data for CaC2 (Ref 14). Other desulfurizerssuch as CaO-CaF2 behave similarly.

Fig. 13 Equilibrium sulfur concentrations for CaO desulfurization calculated with Eq 19 and 20, assuming aCaO =aCaS = aSiO2 = 1. Source: Ref 14.

Fig. 14 Comparisons of the sulfur concentrations in production-continuous desulfurization using CaC2 with thesulfur concentrations from Eq 19 and 20. Open circles indicate electric-melted iron. Closed circles indicatecupola iron. Source: Ref 14.

High desulfurization rates are needed for effective desulfurization in the short times required. For CaO-base desulfurizers,the presence of relatively small amounts of liquid phase (<25 vol%) significantly increases the rate of desulfurization.This is illustrated in Fig. 15, in which the rates of desulfurization by CaO with varying amounts of calcium fluoride(CaF2) are compared. The faster rates obtained with CaF2 additions were due to the formation of a CaO-CaF2 liquid phasethat provided a path for CaS diffusion from the reaction interface into the porous CaO particle. This prevented the normalrapid development of an impervious CaS coating on the CaO surface (Ref 23, 24). A similar explanation, involvingliquid-phase formation, was used to account for the higher CaO desulfurization rates of steel when aluminum wasconcurrently added (Ref 25, 26).

Fig. 15 Desulfurization rates of carbon-saturated iron, containing 0.4% Si, with CaO and varying amounts ofCaF2 at 1450 °C (2640 °F). Source: Ref 22.

Another rate-controlling variable in ladle desulfurization is bath agitation. For gas-stirred melts, the desulfurization rate

constant k is a function of gas flow rate Q•

, that is, k ∝ Q•

n. For well-dispersed solid-liquid mixtures, the rate of

diffusion-controlled interphase reaction is a relatively weak function of Q•

, that is, n ~0.2 to 0.4 (Ref 27, 28, 29). For apoorly dispersed system, such as a ladle of iron with a cover slag of desulfurizer, agitation has a much greater influenceon the rate constant, with n typically in the range of 1.0 to 2.5 (Ref 27, 28, 29). This is due to the entrainment of

increasing amounts of top slag into the liquid metal with increasing Q•

. This effect is illustrated in Fig. 16, which plotsthe apparent mass transfer coefficient for desulfurization versus the gas flow rate.

Fig. 16 Dependence of CaO desulfurization rate constant on the rate of gas flow through a reactor. Source: Ref29.

Deoxidation. In the manufacture of steel, FeO-saturated conditions (~0.2% O) are approached as a result of oxygenblowing to remove carbon and silicon as the respective oxides. To make a steel product, the oxygen levels aresubsequently lowered to 0.005 to 0.02% O by reactive alloy additions of electropositive elements. The low oxygenconcentrations are necessary to maintain the number of oxide inclusions at a suitable low level, to prevent the formationof CO blow-holes, and to ensure effective desulfurization.

For single-element deoxidation, expressed by:

xM + yO = MxOy (Eq 21)

the equilibrium oxygen concentration in the liquid steel, obtained by rearranging the equilibrium constant expression, is:

1

22

(% )(% )

yax y

x x yM O

M OO

K M f f

−

⎛ ⎞= ⎜ ⎟⎜ ⎟⎝ ⎠

(Eq 22)

The calculated equilibrium values of oxygen solubility in liquid iron at 1600 °C (2910 °F), based on Eq 22, are given inFig. 17 for several elements as a function of the concentration of the element, assuming aMxOy = 1. As seen among thecommon deoxidizers, aluminum produces the lowest oxygen. Although not shown, rare-earth metals produce even loweroxygen (Ref 30). Figure 17 plots data for oxygen concentration and oxygen activity. In all cases, oxygen activitydecreases with increasing levels of deoxidizer, but the oxygen concentration can go through a minimum. This occurs incases where the interaction coefficient M

Oe is a large negative number, and as a result, fO decreases significantly even atrelatively low concentrations of M. To obtain exact values for [%O] or hO, equations for K22 can be found in Ref 13 and20.

Fig. 17 Deoxidation equilibria in liquid iron alloys at 1600 °C (2910 °F). Source: Ref 30.

As indicated by Eq 22, the oxygen concentration can be reduced if a complex deoxidation product is formed so thataMxOy < 1. Common deoxidation systems of this type are Si-Mn, Al-Si-Mn, or Al-CaO. An advantage, in addition tolower oxygen, is that less deoxidizer is required in solution to achieve a given level of oxygen. This is illustrated in Fig.18, which plots the oxygen activity of iron as a function of the concentration of CaO in the calcium aluminate inclusion.Separate lines are given for aluminum concentrations ranging from 0.001 to 0.05% Al. Also indicated is the oxygenactivity when pure aluminum oxide (Al2O3) is the reaction product. The data in Fig. 18 show that an oxygen activity of 4ppm can be produced at three aluminum levels--0.002%, 0.005%, or 0.01%--depending on whether the respectivedeoxidation product was CaO-saturated calcium aluminate, CaAl2O4-saturated calcium aluminate, or Al2O3.

Fig. 18 Concentration of oxygen and aluminum in liquid steel in equilibrium with calcium aluminate at 1600 °C(2910 °F); arrows indicate oxygen values when the reaction product is pure Al2O3. Source: Ref 31.

The levels of total oxygen measured in steel melts considerably exceed equilibrium levels (Ref 20). Two possibleexplanations are the kinetic limitations in the deoxidation reaction and the ineffective separation of deoxidation productsfrom the melt. Consideration of the first possibility suggests that the most important kinetic limitation would be oxygenand/or deoxidant solute diffusion to inclusions in the melt (Ref 20). Considering the presence of 105 to 107 inclusionparticles in a cubic centimeter of melt, deoxidation to equilibrium levels of 10 to 20 ppm would require a few minutes.Appreciably longer times would be needed to reduce oxygen to less than 1 ppm, as with rare-earth additions. For mostcases, oxide formation appears faster than oxide particle separation from the bath. Thus, for applications that depend ondissolved oxygen, such as desulfurization or CO blowhole formation, equilibrium oxygen values can be assumed to existafter relatively short reaction periods. For considerations of the inclusion content of steel, the kinetics of particleseparation need to be considered.

The limiting case of particle flotation at velocities obtained from Stokes' law calculations is shown in Fig. 19. Inclusionsize versus time of flotation is plotted. The particles produced during deoxidation are generally small (<5 μm, or 200μin.); accordingly, separation times are long. The times indicated in Fig. 19 are representative of conditions in a quiescentbath. Times are considerably shorter for stirred melts because articles grow by collision and coalescence to sizes of 30 to100 μm (Ref 20). The efficiency of coalescence of particles is material dependent. The energy of adhesion of particles isdetermined by the wetting characteristics of the particle, measured by wetting angle. For example, the force of adhesionof Al2O3 particles with a wetting angle of 140° is twice that for SiO2 with a wetting angle of 115°. On this basis, Al2O3inclusions are expected to cluster more easily than SiO2 inclusions, a phenomenon that has been observed in practice (Ref20).

Fig. 19 Calculated time of flotation of inclusions in stagnant melts as a function of inclusion size. Melt depth: A,50 mm (2 in.); B, 500 mm (20 in.); C, 2000 mm (80 in.). Source: Ref 20.

Aluminum Melts

The principal alloying elements in all cast irons, namely, carbon and silicon, are always the same, but the situation foraluminum is different because there are a number of different important aluminum alloy systems, such as Al-Si, Al-Cu,Al-Mg, and Al-Zn (Ref 32). As a result, the preparation of casting alloys from aluminum scrap holds greater risk ofcontamination than cast iron alloy preparation. In addition to the different elements present in aluminum alloy classes,minor elements are commonly present that are beneficial to some alloy systems but harmful to others. These elementsmust also be controlled in alloy preparation.

An example of minor element control is the case of silicon modifiers: sodium, strontium, antimony, and phosphorus.These elements favorably alter the nucleation and growth kinetics of silicon in aluminum-silicon alloys to produce a morecompact form of the precipitate. Contamination of sodium- or strontium-modified alloys with alloys containingphosphorus or antimony negates the modification. Therefore, control must be exercised over these elements.

Another contaminant that can have positive and negative impact on aluminum castings is hydrogen. Because aluminumreacts readily with atmospheric moisture, hydrogen contamination is common in aluminum alloys (Ref 32). Although thesolubility of hydrogen in liquid aluminum is only of the order of 1.0 ppm, most of the hydrogen precipitates duringsolidification, producing 0.03 volume fraction of gas porosity. This is a serious problem for castings having high strengthrequirements. On the other hand, for castings that are not subjected to high stress, hydrogen-generated porosity is often adesirable condition because it is used to counter the large solidification shrinkage (0.065 volume fraction) associated withaluminum castings.

The control of cast aluminum alloy composition is difficult. Few but the largest foundries attempt to process purchasedscrap; instead they rely on secondary alloy producers with special equipment to perform this function (Ref 33). Becauseof the wide range of elements involved in aluminum alloy casting, the first line in composition control is the separation ofscrap into alloy type and the accurate characterization of the composition of the batch by chemical analysis. Next,compatible batches are combined, and necessary alloy additions are made. This process avoids the difficult task ofremoving unwanted elements. Chemical purification is commonly practiced for hydrogen and for elements that are moreelectropositive than aluminum, namely, the alkali and alkaline earth elements. These separations are usually performed ina single process, with small bubbles of inert or reactive gas removing the undesired elements by evaporation and/oroxidation (Ref 34).

Refining Aluminum by Evaporation Treatment. Some elements dissolved in aluminum have higher vaporpressures than aluminum and can therefore be separated from aluminum melts by inert gas flushing or vacuum treatment.For the evaporation reaction:

m = Mm(g) (Eq 23)

the equilibrium constant is:

( )24

pmm

M

MKh

= (Eq 24)

and for the condition hM = 1, that is, at the hypothetical 1 wt% standard state, the vapor pressure is only a function of K.Using the thermodynamic data given in the article "Thermodynamic Properties of Aluminum-Base and Copper-BaseAlloys" in this Volume, the vapor pressures of selected elements were calculated. The results are plotted in Fig. 20. Thevapor pressure of aluminum is also shown because it represents a practical lower limit to which impurity vapor pressurescan be reduced without incurring significant losses of aluminum. Making this assumption with regard to limiting theimpurity vapor pressure, it can be seen in Table 4 that elements with higher vapor pressures than calcium can beeffectively removed.

Table 4 Approximate theoretical minimum metal compositions to be realized by the vacuum treatment ofaluminumT = 1000 K.

Element Composition, ppm

Sodium 3.7 × 10-7

Cadmium 7.0 × 10-6

Zinc 3.3 × 10-4

Magnesium 1.3 × 10-2

Lithium 3.4 × 10-2

Lead 0.11

Bismuth 0.41

Calcium 39

Indium 110

Antimony 310

Source: Ref 35

Fig. 20 Calculated vapor pressures of selected elements dissolved in aluminum at the hypothetical 1 wt%standard state. Source: Ref 35.

Volatile impurities approach their equilibrium vapor pressures in purge gas bubbles (Ref 35); therefore, the rate ofimpurity removal depends on the vapor pressure of the element and the volumetric gas flow rate (see the article "Gases inMetals" in this Volume). For elements other than hydrogen, the volumetric ratio of purge gas to impurity vapor is so largethat the economics of the process are jeopardized (Ref 35). For alkali and alkaline earth metals, this situation is rectifiedby adding a reactive gas to the purge gas.

Refining Aluminum With Reactive Gases. Alkali and alkaline earth metals form more stable halides and oxidesthan aluminum; therefore, by adding F2, Cl2, or O2 to the purge gas, these elements can be separated from aluminum.

Because the reaction products in this case are condensed phases, the rate of impurity removal no longer depends on buton the gas/liquid interfacial area, where the reactions take place. For the generalized reaction between a metal M andhalogen or oxygen X2 gas:

2 21

m mxM xX M Xm

+ = (Eq 25)

the equilibrium constant is:

/2

262

( )( )

a l mm mx

xM x

M XKh p

= (Eq 26)

and the equilibrium X2 pressure is:

( )1/1/

22

26

ymam mx

xM

M Xp

K h

⎡ ⎤⎢ ⎥=⎢ ⎥⎣ ⎦

(Eq 27)

If px2 for an impurity element is less than px2 for the corresponding reaction with aluminum, then impurity removal istheoretically possible. Conversely, if p x2 for the impurity element is greater than p x2 for aluminum, the aluminumcompound will separate in preference to the impurity. Figure 21 and 22 plot p x2 versus hM, respectively, for reactions withchlorine and fluorine. The oxide data were not included, because oxygen is much less effective than the halides (Ref 35).The data in Fig. 21 and 22 assume aMmX2mx = 1.

Fig. 21 Calculated equilibria for the ternary Al-M-Cl system at 1000 K. Source: Ref 35.

Fig. 22 Calculated equilibria for the ternary Al-M-F system at 1000 K. Source: Ref 35.

Table 5 compares the minimum impurity concentrations that are possible by treatment with Cl2, F2, and O2. For all thecases examined, chlorine appears to be the best reagent. Chlorine is the most commonly used reagent for removing theseelements from aluminum melts. In practice, the thermodynamically predicted levels are not achieved (Ref 35).

Table 5 Approximate theoretical minimum contents to be realized by refining with reactive gasesT = 1000 K

Material content, ppm,remaining after oxidation with:

Element

Oxygen Chlorine Fluorine

Calcium <4000 3 × 10-5 5.7 × 10-4

Lithium <270 0.01 0.046

Magnesium 1160 1.4 4.9

Sodium <950 9 × 10-5 0.19

Source: Ref 35

The use of halogens has a disadvantage in terms of the environmental precautions that need to be taken. This hasprompted the development of numerous electrochemical separation processes.

Electrochemical Refining of Aluminum. An electrochemical process for separating magnesium from moltenaluminum is illustrated in Fig. 23. Through the application of a potential, magnesium in the aluminum melt (anode)dissolves in the electrolyte, while pure magnesium deposits at the cathode. As seen at left in Fig. 23, densities of the threephases form a convenient system for cell construction, with the aluminum phase having the highest density and themagnesium phase having the lowest. Therefore, liquid magnesium is easily removed from the top of the cell. Anadvantage of the electrochemical method is that the recovered metallic magnesium has high commercial value comparedto the MgCl2 produced by reactive gas separations.

Fig. 23 Schematic of an electrochemical magnesium separation apparatus. Source: Ref 36.

The open-circuit cell potential ε for this system can be estimated from the Nernst equation:

(Eq 28)

where is Faraday's constant and (aMg)cathode = 1. For producing alloys with less than 0.1% Mg, as required for a numberof important casting alloys, such as 319 and 380, Eq 28 predicts that a cell potential of 0.32 V is required (Ref 36). Thisvoltage must not be greatly exceeded, because aluminum will be transferred from anode to cathode at a potential of 0.5 V.

References cited in this section

13. R.J. Fruehan, Ladle Metallurgy: Principles and Practices, Iron and Steel Society of AIME, 1985, p 7 14. S. Katz and C.F. Landefeld, Cupola Desulfurization, in Cupola Handbook, American Foundrymen's

Society, 1984, p 351-363 15. I.D. Sommerville, The Capacities and Refining Capabilities of Metallurgical Slags, in Foundry Processes:

Their Chemistry and Physics, S. Katz and C.F. Landefeld, Ed., Plenum Press, 1988, p 101-133 16. G.K. Sigworth and J.F. Elliott, The Thermodynamics of Liquid Dilute Iron Alloys, Met. Sci., Vol 8, 1974, p

298-31017. S. Katz and H.C. Rezeau, The Cupola Desulfurization Process, Trans. AFS, Vol 87, 1979, p 367-376 18. S. Katz, D.E. McInnis, D.L. Brink, and G.A. Wilkinson, The Determination of Aluminum in Malleable Iron

From Measured Oxygen, Trans. AFS, Vol 88, 1980, p 835-844 19. R.J. Fruehan, Ladle Metallurgy: Principles and Practices, Iron and Steel Society of AIME, 1985, p 8 20. E.T. Turkdogan, Physical Chemistry of High Temperature Technology, Academic Press, 1980 21. S. Katz and V.R. Spironello, Effect of Charged Aluminum on Iron Temperature, Silicon Recovery and

Desulfurization in an Iron-Producing Cupola, Trans. AFS, Vol 92, 1984, p 161-172 22. C.F. Landefeld and S. Katz, Kinetics of Iron Desulfurization by CaO-CaF2, in Proceedings of the Fifth

International Iron and Steel Congress, Vol 6, Iron and Steel Society of AIME, 1986, p 429-439 23. S. Katz and C.F. Landefeld, Plant Studies of Continuous Desulfurization with CaO-CaF2-C, Trans. AFS,

Vol 93, 1985, p 215-228 24. S. Katz and B.L. Tiwari, A Critical Overview of Liquid Metal Processing in the Foundry, in Foundry

Processes: Their Chemistry and Physics, S. Katz and C.F. Landefeld, Ed., Plenum Press, 1988, p 1-52 25. J. Niederinghaus and R.J. Fruehan, Desulfurization Mechanisms for CaO-Al and CaO-CaS in Carbon

Saturated Iron, Metall. Trans. B, to be published26. E.T. Turkdogan, Physiochemical Phenomena of Mechanisms and Rates of Reaction in Melting, Refining

and Casting of Foundry Irons, in Foundry Processes: Their Chemistry and Physics, S. Katz and C.F.Landefeld, Ed., Plenum Press, 1988, p 53-100

27. S. Asai and I. Muchi, Fluid Flow and Mass Transfer in Gas Stirred Ladles, in Foundry Processes: TheirChemistry and Physics, S. Katz and C.F. Landefeld, Ed., Plenum Press, 1988, p 261-292

28. S.-H. Kim and R.J. Fruehan, Physical Modelling of Liquid/Liquid Mass Transfer in Gas Stirred Ladles,Metall. Trans. B, Vol 18B, 1987, p 381-390

29. J. Ishida, K. Yamaguchi, S. Sugiura, N. Demukai, and A. Notoh, Denki Seiko, Vol 52, 1981, p 2-8 30. E.T. Turkdogan, Ladle Deoxidation, Desulfurization and Inclusions in Steel--Part I: Fundamentals, Arch.

Eisenhüttenwes., Vol 54, 1983, p 1-10 31. E.T. Turkdogan, Slags and Fluxes for Ferrous Ladle Metallurgy, Ironmaking Steelmaking, Vol 12, 1985, p

64-7832. Aluminum Casting Technology, American Foundrymen's Society, 198633. Recycled Metals in the 1980's, National Association of Recycling Industries, 1982 34. J.H.L. Van Linden, R.E. Miller, and R. Bachowski, Chemical Impurities in Aluminum, in Foundry

Processes: Their Chemistry and Physics, S. Katz and C.F. Landefeld, Ed., Plenum Press, 1988, p 393-409 35. G.K. Sigworth and T.A. Engh, Refining of Liquid Aluminum--A Review of Important Chemical Factors,

Scand. J. Metall., Vol 11, 1982, p 143-149 36. B.L. Tiwari and R.A. Sharma, Electrolytic Removal of Magnesium From Scrap Aluminum, J. Met., Vol 36

(No. 7), 1984, p 41-43

Note cited in this section

* The actual differences are less than those indicated. Calcium sulfide has only limited solubility in CaO-Al2O3slags. As a result, a minimum W = 0.01 is needed to maintain the CaS-unsaturated condition.

References

1. R.I.L. Guthrie, Addition Kinetics in Steelmaking, in Proceedings of the 35th Electric Furnace Conference,Iron and Steel Society of AIME, 1978, p 30-41

2. F.A. Mucciardi and R.I.L. Guthrie, Aluminum Wire Feeding in Steelmaking, Trans. ISS, Vol 3, 1983, p53-59

3. J.W. Robison, Jr., "Ladle Treatment With Steel-Clad Metallic Calcium Wire," Paper 35, presented atScaninject III, Part II, MEFOS, Lulea, Sweden, 1983

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