Volume 15 - Casting Part 7 potx

Lambert, Operation of Electron Beam Furnace for Melting Refractory Metals, in Proceedings of the Bakish Conference on Electron Beam Melting and Refining Reno, NV, R.. Table 7 Comparison

Vertical Feeding of Ingots. Refractory and reactive metal ingots of high purity, homogeneity, and smooth surface are remelted by vertical feeding (Fig 35d) The molten metal droplets run down the conical, rotating electrode tip, are refined, and then drop into the pool center The crucible pool is normally of the same diameter as the electrode but is sometimes smaller or larger It is kept in the liquid state to allow final refining and to guarantee ingot homogeneity Because two or more electron guns are used, the entire pool can be equally bombarded; thus, shadow effects of the electrode can be eliminated

Simultaneous melting of horizontally and vertically fed electrodes (Fig 35e) can be used for the production of critical alloys In this case, the feedstock should be of the desired purity

Horizontally Fed Ingots. Drip melting of horizontally fed material with a single electron gun (Fig 36) is used for refining some steel alloys in East Germany and other Soviet bloc countries In this process, the feedstock size is smaller than the pool diameter to minimize the shadow effect of the horizontally fed bar In production units, feeding can be carried out from two opposite sides

Fig 36 Drip melting of 330 mm (13 in.) square steel billet in a 1100 kW single-gun furnace Melt rate: 1000

kg/h (2200 lb/h) Courtesy of VEB-Edelstahlwerk, East Germany

Other Process Considerations. To ensure the production of clean, homogeneous metals and alloys in electron beam

drip melting furnaces, various aspects of material processing and handling must be controlled Key considerations include:

• Dimensions and quality of the feedstock, and the feeding system used

• Ingot cooling and unloading during melting of another ingot

• Passivation and removal of condensates from the melt chamber

• Planning of melt sequences to minimize the number of furnace cleanings required

• Routine preventive furnace maintenance to ensure reliability

• Operator skill in operation of the furnace

• Material yield and energy consumption

Equipment for Drip Melting

The essential equipment groups required for drip melting melting furnaces, control systems, and power supply units are all important for achieving optimum productivity

The melting furnace (Fig 37) includes the electron beam gun as the heat source, material feeding and ingot withdrawal systems, a crucible for material solidification, and a vacuum system to maintain the low pressure Process observation, both visually and with video systems, is possible through viewports The melt chamber flanges are equipped with x-ray absorbing steel boards, and interlocking systems prevent operation failures and accidents

Fig 37 Single-gun 1200 kW furnace for horizontal drip melting of steels Melting rates of up to 1100 kg/h

(2425 lb/h) are possible

The control system allows the adjustment and control of such operating process parameters as electron beam power, operating vacuum level, material feed rate, and ingot withdraw speed The control system also records and logs the process data

Power Supply Units. One or more high-voltage power supply units are needed to supply the electron beam guns with the required continuous voltage (30 to 40 kV) The beam power of each gun can be adjusted between zero and maximum power with an accuracy of ±2%

Other Equipment. Large production furnaces are equipped with lock-valve systems to allow simultaneous melting and unloading of ingots without breaking the vacuum in the melt chamber Production is thus limited only when the condensate remaining in the melt chamber requires cleaning or when a different alloy is to be melted

Characteristics of Electron Beam Drip Melted Metals

Electron beam melted and refined material is of the highest quality The amount of interstitials present is very low, and trace elements of specific high vapor pressure can also be reduced to very low values (Ref 37, 38)

Reactive and Refractory Metals

Tantalum and niobium ingots have smooth surfaces and are of sufficient ductility that they can be cold worked, and sheets and wires can be produced

Tungsten and molybdenum ingots are also of the highest possible purity, but the ingots are brittle because of the very large grain size and the concentration of impurities at grain boundaries

Hafnium. Electron beam melted hafnium is of higher ductility than the vacuum arc remelted metal (Ref 39) The main application of electron beam melted hafnium is as control elements for submarine nuclear reactors

Vanadium is refined by electron beam drip melting The aluminothermically produced feedstock is drip melted in several steps During this procedure, the ingot diameter is reduced at each step by approximately 30 to 40 mm (1.2 to 1.6 in.) to obtain an ingot 30 to 40 mm (1.2 to 1.6 in.) in diameter, regardless of the initial ingot diameter The clean vanadium ingots are primarily used in nuclear reactor applications (Ref 40)

Applications for electron beam melted refractory and reactive metals are listed in Table 6

Table 6 Principal applications for vacuum arc remelted (VAR), electron beam melted (EB), and powder metallurgy (P/M) reactive and refractory metal ingots

Metal Applications

Reactive metals, VAR and EB melting

Hafnium Flash bulbs and glow discharge tubes for the electronics industry; control rods and breakoff elements in submarine

nuclear reactors

Vanadium Targets for high deposition rate sputtering processes in the electronics industry; breakoff elements, fixtures, and fasteners

in nuclear reactors; standards for basic research; alloying element for certain high-purity alloys

Zirconium Getter material in tubes in the electronics industry; stripes for flash bulbs; fuel claddings, fasteners, and fixtures for

nuclear reactors

Titanium Components for bleaching equipment and desalination plants in the chemical industry; superconductive wires; turbine

engine disks, blades and housings, rain erosion boards, landing legs, wing frames, missile cladding, and fuel containers in the aircraft and aerospace industries shape memory alloys; biomedical fixtures and implants; corrosion resistant claddings

Refractory metals, EB melting and P/M

Tungsten Heating elements, punches and dies, and nonconsummable electrodes for arc melting and gas tungsten arc welding for

metal processing equipment; targets for x-ray equipment and high sputtering rate devices such as very large-scale integrated circuits, cathodes and anodes for electronic vacuum tubes in the electronics industry; radiation shields in the nuclear industry; cladding and fasteners for missile and reentry vehicles

Tantalum Condensers, autoclaves, heat exchangers, armatures, and fittings for the chemical industry; electrolytic capacitors for the

electronics industry; surgical implants; fasteners for aerospace applications

Molybdenum Dies for conventional and isothermal forging equipment; electrodes for glass melting; targets for x-ray equipment;

cladding and fasteners for missile and reentry vehicles

Niobium Superconductive wire for energy transmission and large magnets for the electrical and electronics industries; heavy ion

accelerators and radio frequency cavities for nuclear applications; components for aircraft and aerospace applications

Steels

The purity and properties of electron beam melted steels are in some respects better than those of vacuum arc and electroslag remelted steels, but the processing costs are higher The electron beam melting of steel is primarily used in East Germany and other Soviet bloc countries The resulting ingots are up to 1000 mm (40 in.) in diameter and weigh up

to 18 Mg (20 tons) The furnaces used have been in operation since 1965, and have beam powers of up to 1200 kW Larger furnaces for the production of ingots weighing up to 30 to 100 Mg (33 to 110 tons) are under construction (Ref 41)

The essential advantage of the electron beam melting of steel is the drastic reduction of metallic and nonmetallic impurities and interstitial elements (Ref 42, 43) The principal applications for electron beam melted steels are in the machinery industry for parts for which high wear resistance and long service life are required The extended service lives

of the parts and the reduced manufacturing time (for example, less surface polishing is required for electron beam melted steel) can justify the higher material costs

The electron beam melting of steel and superalloys can become much more economical when melting and refining are done by continuous flow melting or cold hearth refining These melting and refining methods reduce energy costs and minimize material losses

References cited in this section

37 R.E Lüders, Tantalum Melting in a 800 kW EB Furnace, in Proceedings of the Bakish Conference on Electron Beam

Melting and Refining (Reno, NV), R Bakish, Ed., 1983, p 230-244

38 J.A Pierret and J.B Lambert, Operation of Electron Beam Furnace for Melting Refractory Metals, in Proceedings of

the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1984, p 208-218

39 H Sperner, Hafnium, Metallwis Technik., Vol 7 (No 16), 1962, p 679-682

40 R Hähn and J Krüger, Refining of Vanadium Aluminium Alloys to Vanadium 99.9% by EB Melting, in Proceedings

of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1986, p 53-67

41 J Lambrecht, D Rumberg, and K.H Werner, Stand der Technologie der Stahlerzeugung im Elektronenstrahlofen mit

Blockmassen bis 100 t, Neue Hütte, Vol 10, Oct 1984

42 C.E Shamblen, S.L Culp, and R.W Lober, Superalloy Cleanliness Evaluation Using the EB Button Melt Test, in

Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1983

43 F Hauner, H Stephan, and H Stumpp, Ergebnisse bei Elektronenstrahl-Schmelzen von gerichtet erstarrten

Knopfproben zur Identifizierung von nichtmetallischen Einschlüssen in hochreinen Superlegierungen, Metall., Vol 2

(No 40), 1986, p 2-7

Continuous Flow Melting

The continuous flow melting process (cold hearth refining process) (Fig 38) was developed approximately 10 years after drip melting (Ref 44) Continuous flow melting is mainly used for refining specialty steels and superalloys and for refining and recycling reactive metal scrap, especially Ti-6Al-4V from high-density tungsten carbide tool tips (Ref 45)

Fig 38 Schematic of the continuous flow melting process

Principles of Continuous Flow Melting

Continuous flow melting (Fig 39) is the most flexible vacuum metallurgical melting process It is a two-stage process in which the first step (material feeding, melting, and refining) takes place in a water-cooled copper trough, ladle, or hearth

In the second step, solidification occurs in one of several round, rectangular, or specially shaped water-cooled continuous copper crucibles Both process steps are nearly independent from each other; they are linked only by the continuous flow

of the liquid metal stream The major refining actions are carried out in the hearth, but some postrefining takes place in the pool of the continuous casting crucible, similar to the drip melting of horizontally fed billets Refinement in continuous flow melting occurs by vacuum distillation in the hearth pool, superheating, and stirring of the molten metal pool

Fig 39 Four-gun 1200 kW combined electron beam drip melting and continuous flow melting furnace

Removal of Impurities. Most impurities with densities lower than that of the melt (for example, metalloids in steels and superalloys) can be segregated by flotation and formed into a slag raft The raft is then held in place by either mechanical or electrothermal means Impurities denser than the melt, such as tungsten carbide tool tips in titanium, are removed by sedimentation Inclusions with densities such that efficient flotation or sedimentation does not occur can be partially removed by adhesion to the slag raft

Hearth dimensions are based on the type and amount of refining required For example, hearths for vacuum distillation should be nearly square and relatively deep to allow sufficient melt stirring For flotation refining, the hearth should be long and narrow (for superalloys, approximately 10 mm, or 0.4 in., of hearth length for each 100 kg/h, or 220 lb/h, of melt rate is recommended) Hearths for titanium alloy scrap recycling can be relatively short if all the materials can be transported to the pool of the hearth rather than to the ingot pool

Feeding. Material feeding criteria include 100% homogenous material transportation to avoid uncontrolled evaporation

of alloying elements and correct feeding into or above the hearth pool Horizontal feeding of compacted, premelted, or cast material is most often used Loose scrap and raw material are used only when compaction is too expensive Feeding

of liquid metal was used in one of the first continuous flow melting furnaces to produce a ferritic steel in a vacuum induction furnace (Ref 46) Postrefining was carried out in a cascade of five hearths 1.5 m (60 in.) long and 1 m (40 in.) wide

Casting and Solidification. The criteria for material casting and solidification include the shape of the final product and the solidification rate required to avoid ingot tears or other defects and to ensure a homogeneous ingot structure The multiple casting of small ingots is sometimes used, especially when forging is impossible because of the brittleness of the solidified material (for example MCrAly wear-resistant coating alloys) The casting of round and rectangular ingots and slabs is common practice, and the continuous casting of hollow ingots is also being used (Ref 47) The casting of segregation-free ingots and ingots with a fine grain size is under development to improve the workability of superalloys (Ref 48, 49)

Continuous Flow Versus Drip Melting

Table 7 compares the essential features of drip melting and continuous flow melting Generally, continuous flow melting

is used for all refractory metals, superalloys, and specialty steels, especially when flotation or sedimentation of inclusions

is required Drip melting is used for refractory metals because of their high melting points and the resulting high heat losses to the water-cooled copper crucible Depending on production quantity, double or triple drip melting may require less energy than a single continuous flow melt of some materials, such as niobium

Table 7 Comparison of the characteristics of drip melting and continuous flow melting

Characteristic Refractory metals Reactive metals, superalloys, and specialty steels

Power density High Soft; smoothly distributed

Ingot shape and structure Round; coarse grain Round or flat; fine grain, segregation-free

Competitive economical processes Vacuum arc remelting Vacuum arc remelting; electroslag remelting

Preferred method Drip melting Continuous flow melting

Refining and Production Data

Data on continuous flow electron beam melting and refining in laboratory and pilot production furnaces are given in Table 8 The data demonstrate the effectiveness of the process in reducing impurities and interstitial elements It can also

be seen that the selective evaporation of chromium from superalloys can be controlled by the distribution of beam power

at the trough pool and by controlling trough pool area and melt rate The selective evaporation of aluminum from 4V alloy is much more difficult to control; additional aluminum must be used to compensate for the aluminum evaporated

Ti-6Al-Table 8 Refining and production data for the continuous flow melting of reactive and refractory metals and stainless steels in laboratory and pilot production furnaces

Composition of feedstock and product

size,

mm (in.)

Trough size,

mm (in.)

Ingot size,

mm (in.)

Ingot weight,

kg (lb)

Melt rate, kg/h (lb/h)

Electron beam power,

kW

Operating pressure,

Pa (torr)

Specific melting energy,

kW · h/kg C, ppm O, ppm N, ppm H, ppm Al, % V, % Cr, %

900

square

120 × 250 (5 × 10)

(5 × 12)

150 (6) 90.5 (200) 42 (92.5) 185 3.5 × 10-2

(2.6 × 104

100 (4) 40.2 (89) 80 (176) 140 3.5 × 10-2

(2.6 × 104

-)

1.75

1520 210 3

1045 210 10 99 Vanadium 50 (2) square 120 × 300

(5 × 12)

100 (4) 20 (44) 130 1.5 × 10-2

(1.1 × 104

-)

6.5

277 50 3 99

400 2600 110 84 6.0 4.0 Ti-6Al-4V Swarf 120 × 300

(5 × 12)

150 (6) 62.6 (138) 40 (88) 122 2 × 10-2

(1.5 × 104

-)

2.0 1520 1520 75 15 6.0 4.0

(5 × 12) )

1320 76 8 4.8 4.1

6.0 4.0 Ti-6Al-4V 125 (5) diam 150 × 400

(6 × 16)

2 × 75 (3) diam

2 × 32 (70.5)

91 (200) 147 6 × 10-2

(4.5 × 104

-)

1.61

3.6 4.3

Commercially pure

titanium

160 (6.3) 150 × 250

(6 × 10)

100 × 400 (4 × 16)

96.4 (213) 86.3

(190)

148 6 × 10-2

(4.5 × 104

103.0 (227) 41.2 (91) 226 8 × 10-2

(6 × 10-4)

5.5

701 97 155 18.25 Stainless steel 150 (6) diam 150 × 400

(6 × 16)

2 × 75 (3) diam

2 × 55 (121) 136 (300) 144 6 × 10-2

(4.5 × 104

2 × 75 (3) diam

2 × 57 (126) 136 (300) 156 6 × 10-2

(4.5 × 104

3 × 41.5 (91.5)

136 (300) 156 6 × 10-2

(4.5 × 104

-)

1.15

Source: Ref 50

References cited in this section

44 C.d'A Hunt and H.R Smith, Electron Beam Processing of Molten Steel in Cold Hearth Furnace, J Met., Vol 18,

1966, p 570-577

45 H.R Harker, Electron Beam Melting of Titanium Scrap, in Proceedings of the Bakish Conference on Electron Beam

Melting and Refining (Reno, NV), R Bakish, Ed., 1983, p 187-190

46 C.d'A Hunt, H.R Smith, and B.C Coad, The Combined Induction and Electron Beam Furnace for Steel Refining and

Casting, in Proceedings of the Vacuum Metallurgy Conference (Pittsburgh, PA), June 1969, p 1-22

47 H.R Harker, The Present Status of Electron Beam Melting Technology, in Proceedings of the Bakish Conference on

Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1986, p 3-7

48 C.d'A Hunt, J.C Lowe, and T.H Harrington, Electron Beam, Cold Hearth Refining for the Production of Nickel and

Cobalt Base Superalloys, in Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno,

NV), R Bakish, Ed., 1984, p 295-304

49 H Stephan, R Schumann, and H.J Stumpp, Production of Superclean Fine-Grained Superalloys for Improvement of

the Workability and Engine Efficiency by EB Melting and Refining Methods, in Proceedings of the Eighth International Conference on Vacuum Metallurgy (Linz), 1985, p 1219-1309

50 H Stephan, Production of Ingots and Cast Parts From Reactive Metals by Electron Beam Melting and Casting, in

Proceedings of the Third Electron Beam Processing Seminar (Stratford, UK), 1974, p 1b1-1b69

Equipment for Continuous Flow Electron Beam Melting

The equipment required for continuous flow melting is different from that used in drip melting mainly because of the trough and the somewhat larger melting chamber In addition, because of the materials often melted in the continuous flow process (superalloys and titanium alloys), additional instrumentation is often provided This may include an ingot pool level control system, metal vapor and partial pressure analyzers, a two-color temperature control system, and a data logging system

Accurate beam power distribution is achieved in two- or three-gun furnaces by microprocessor control, which allows the splitting of a single beam to 64 locations and the adjustment of dwell time at each location between 0.01 and 1000 s The beam spot at each of the 64 locations can be scanned over an elliptical or rectangular area With such systems, the required refining can be achieved without unnecessary power consumption and evaporation of alloying elements (Ref 51)

Process observation is accomplished with a video monitoring system Samples can be obtained from both the trough pool and the ingot pool for nearly continuous control of material quality

Feeding systems for continuous flow furnaces must maintain homogeneity along the length of the feed material The trough and crucible should be easily accessible for convenient maintenance, especially when different alloys are to be melted in the same furnace

Reference cited in this section

51 H Ranke, V Bauer, W Dietrich, J Heimerl, and H Stephan, Melting and Evaporation With the Newly Developed

Leybold-Heraeus 600 kW EB Gun at Different Pressure Levels, in Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1985

Characteristics of Continuous Flow Melted Materials

Titanium ingots and slabs can be produced from titanium scrap contaminated with tungsten carbide tool tips The electron beam melted product contains tungsten carbide particles no larger than 0.7 mm (0.028 in.) in diameter Oxygen content can be reduced by fitting titanium sponge compacts around a forged ingot or slab Continuous flow melted titanium ingots can be directly remelted in a VAR furnace

Superalloys. Continuous flow electron beam melted superalloy ingots 150 to 200 mm (6 to 8 in.) in diameter are often used in VIM investment casting furnaces Such ingots are nearly free of nonmetallic inclusions and trace elements (Ref 52) The simultaneous continuous casting of twin, triple, or multiple ingots is under development in a 200 kW pilot

furnace Multi-ingot casting can become economical when 90 Mg (100 tons) or more of an alloy is being produced annually (Ref 53)

References cited in this section

52 M Krehl and J.C Lowe, Electron Beam Cold Hearth Refining for Superalloy Revert for Use in Foundry Production,

in Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1986, p

286-296

53 M Romberg, R Schumann, H Stephan, and H Stumpp, Electron Beam Melting and Refining of Superalloys for

Ingot and Barstick Production, in Proceedings of the Bakish Conference on Electron Beam Melting and Refining

(Reno, NV), R Bakish, Ed., 1986

Investment Casting Using Electron Beam Melting

Investment casting with electron beam heat sources from water-cooled copper skull crucibles has been used for the production of superalloy turbine parts with directionally solidified or monocrystal structures and for titanium parts with equiaxed structures In both applications, the feedstock is clean material that is melted with no major refining effect and without picking up any contamination The process competes for superalloys with vacuum induction melting and investment casting and for titanium with vacuum arc skull melting and casting

Characteristics of Electron Beam Investment Casting

With electron beam melting, it is possible to superheat the casting material just before and during pouring to increase the fluidity of the metal This is not possible with vacuum arc skull melting The extremely short metal flow path between the crucible spout and the mold funnel reduces erosion of the mold, and the controllable pouring speed allows the use of small mold funnels for reducing revert material An essential advantage of electron beam titanium casting is the possibility of using premelted VAR ingots and solid, clean in-house recycling scrap separately or simultaneously to reduce material costs The decisive advantage of electron beam casting technology is the possibility of process automation, which guarantees reproducibility of quality and a high production rate Therefore, electron beam melting is used only when a high production rate is demanded The obvious disadvantage is the decreasing metal temperature during pouring caused by the temperature gradient of the melt in the water-cooled copper crucible; for this reason, the process cannot be used for the production of superalloy cast parts with very specific equiaxed structures

Process Technology and Equipment for Superalloy Castings

For the economical mass production of superalloy turbine parts, the electron beam casting process and equipment shown

in Fig 40 and described in Table 9 meet the required process and product specifications The pool temperature before casting can be adjusted between 1650 and 1900 °C (3000 and 3450 °F), depending on part requirements The adjusted temperature can be reproduced with an accuracy of 0.3%

Table 9 Processing data for the mass production of directionally solidified and monocrystal superalloy castings in 120 and 300 kW electron beam casting machines

Pa (torr)

Leak and degassing rate, mbar · L/s

Electrode dimensions,

mm (in.)

Number of electrodes per magazine

Casting weight,

kg (lb)

Average weight deviation,

%

Cycle time,

s

Pool surface temperature

at pouring,

°C ( °F)

Average temperature deviation,

%

Material loss by evaporation and

splattering, %

120 kW 120 2 × 10-4 (1.5 × 10-6) 2 × 10-4 125 (5) diam; 1600 (63) long 12 0.45 (1) 1.8 60 1850 (3360) 0.3 1.5

300 kW 300 2 × 10 -4 (1.5 × 10 -6 ) 2 × 10 -4 125 (5) diam; 1600 (63) long 12 8 (18) 1.5 440 1850 (3360) 0.3 1.5

Specific electron beam energy

s

Electron beam power, kW

kW/kg kW/lb

A Heating of electrode tip and positioning into melting area 4 80 0.20 0.090

B Distribution of electron beam power to the electrode only 0.5 80 0.02 0.009

C Drip melting of rotating electrode with simultaneous droplet counting and

weight control

D Retraction of electrode and distribution of increased power to the melt pool 0.5 100 0.74 0.33

F Pouring with simultaneous pool superheating 0.7 100 0.04 0.33

G Washing of crucible for reduction of skull weight 6.0 60 0.22 0.10

H Cleaning of the spout lip to obtain reproducible pouring streams; tilting back of

crucible

1.0 60 0.04 0.02

Fig 40 Procedure and process data for the electron beam melting and casting of 0.45 kg (1 lb) superalloy

parts Cycle time: 60 s; total electron beam energy: 3.01 kW · h/s

The accuracy of the part weight was 1.8% for small parts (450 g, or 1 lb) and 1.5% for larger parts (8 kg, or 18 lb) The cycle time for smaller parts was 58 ± 1.5 s; for larger parts, 300 ± 10 s

Reduction in chromium content was less than 0.1% Feedstock load is twelve electrodes 125 mm (5 in.) in diameter and

1600 mm (63 in.) long The electrode magazine can be reloaded without venting the melt chamber Exchange of cooled copper crucibles, for changing alloys or part weight, can also be accomplished without venting the melt chamber

water-More than 50% of the evaporating and splattering material can be condensed and collected at the condensate plate This plate can be replaced through a lock-valve system

Vacuum pressure during melting and pouring is in the range of 0.001 to 0.0001 Pa (10-5 to 10-6 mbar, or 7.5 × 10-6 to 7.5 ×

10-7 torr) Leak and wall degassing rates are below 2 × 10-4 mbar · L/s

Melting and pouring within one electrode cycle can be done automatically Movement of an electrode from the storage position to the melting position is semiautomatic

The distance between the copper crucible pouring spout and the mold funnel was less than 40 mm (1.6 in.) during pouring The mold funnel cross section was 20 to 30 mm (0.8 to 1.2 in.) for casting small parts

These results can be achieved with a computer-controlled electron beam melting furnace using the process data given in Table 9 The fundamental equipment groups are the means for accurate beam power distribution at the electrode tip and

crucible pool and its programmable power level during the many steps of process procedure, and integrated droplet counter and balance to control the material to be drip melted, and the weight of the remaining skull, which will be remelted again and again at each step The high degree of process reproducibility is achieved by controlling pool surface temperature, crucible washing, and crucible spout lip cleaning Crucible washing and spout lip cleaning minimize the skull weight remaining in the crucible and ensure a reproducible beam pattern for pouring in a small mold funnel A computer-controlled production casting system that uses electron beam melting is described in detail in Ref 54

Process Technology and Equipment for Titanium Castings

The production of titanium castings in electron beam furnaces is economical when large quantities of the same size and weight must be manufactured and when the entire specification range of alloy compositions can be used Electron beam casting is economical and superior to the competing VAR skull melting and nonconsumable arc melting processes; a blend of solid scrap and consumable electrodes can be used because the crucible spout can be cleaned by electron beam melting after each pour Figure 41 illustrates the beam control possible with microprocessor control Another advantage

of this process is the possibility of superheating the pool just before pouring and the metal stream during pouring During the melting and casting process, the feedstock can be loaded into the scrap feeder or into the electrode lowering device without influencing the process cycle Mold loading, transportation, positioning below the crucible spout, spinning there with up to 500 rpm, and unloading of the cooled casting are completely automated

Fig 41 Procedure and process data for the electron beam melting and casting of titanium and titanium alloy

parts starting from scrap and consumable electrodes (a) Melting of skull with two beams, each scanning and jumping on two circles (500 kW, 7 min) (b) Melting of scrap with two beams, each scanning and jumping on different circles (500 kW, 7 min) (c) Melting of consumable electrode with two concentrated beams (500 kW, 6 min) (d) Superheating of the pool with two beams, each scanning and jumping on two different circles (500

kW, 3 min) (e) Superheating of the metal stream during pouring with two concentrated beams (500 kW, 10 s) (f) Cleaning of the mold spout with two slightly scanning beams (10 kW, 2 min)

Reference cited in this section

54 J Mayfield, Computer-Controlled Production Gains, Aviat Week Space Technol., 3 Dec 1979, p 1-5

Plasma Melting and Casting

H Pannen and G Sick, Leybold AG, West Germany

Plasma melting is a material processing technique in which the heat of a thermal plasma is used to melt the feed material

A thermal plasma is considered a suitable heat source if high temperature and a defined gas atmosphere are needed to

melt the material before subsequent processing, such as solidification or atomization The products are usually ingots, slabs, castings, or powders

Plasma Torch

Plasma Generation. An electric current passing through an ionized gas leads to a phenomenon known as gaseous discharge To generate this kind of plasma, electrical breakdown of the gas must be accomplished Breakdown (that is, the creation of charge carriers) establishes a conducting path between a pair of electrodes in gases that are insulators at room temperature Other plasmas can be created by electrodeless radio frequency (RF) discharges, microwaves, shock waves, lasers, or high-energy particle beams

In this section, only thermal (hot) plasmas will be considered The term thermal differentiates these plasmas from cold (nonequilibrium) plasmas, which are characterized by high electron temperatures and low sensible temperatures of the heavy particles Cold plasmas are produced in various types of glow discharges, low-pressure RF discharges, and corona discharges Cold plasmas are not suitable for melting processes

Most plasma generators (plasma torches) for melting processes use an electric arc to produce gaseous discharges The characteristics of an electric arc include relatively high current densities, low cathode fall, and high luminosity of the column

A typical potential distribution along an arc is shown in Fig 42 (Ref 55) The column cross section at the cathode is smaller than at the anode, with contrary current and power densities This fact influences the design principle of plasma torches

Fig 42 Typical potential distribution along a plasma arc Va, anode voltage; Vc, cathode voltage; da , anode

current density; dc , cathode current density

Torch Design. Plasma torches can be used in either the transferred or the nontransferred mode In the nontransferred mode, the cathode and the anode are inside the torch This mode is generally suitable only for the melting of nonconductive materials and has the disadvantage of lower efficiency compared to the transferred mode In the transferred mode, one electrode is inside the torch, and the counter-electrode is the material to be melted Figure 43 shows two typical design principles for torches in the transferred mode: the tungsten tip design and the hollow copper electrode design

Fig 43 Design concepts for plasma arc torches in the transferred mode (a) Torch with tungsten tip and

concentric gas flow (b) Torch with hollow copper electrode and vortex generator

In the tungsten tip design (Fig 43a), the torch electrode is connected as the cathode because high current densities are required The cathodes are usually made of tungsten, along with small additions of thoria to lower the thermionic work function of tungsten Still, electron emission requires high electrode temperatures (3500 to 6000 K) at the attachment of the arc Therefore, the cathode material is liquid at the arc attachment point even with water cooling at the rear of the cathode For this reason, oxidizing gases generally cannot be used in direct contact with the tungsten tip Argon shielding is one solution Further, it should be kept in mind that portions of the small liquid pool at the cathode tip are ejected into the material being melted The high melting point of tungsten can result in inclusions in the product; such defects are not tolerable in high-strength metals such as titanium alloys and superalloys for aircraft applications

With argon as the plasma gas, the service life of the cathode in the tungsten tip design is typically between 30 and 150 h The inert gas consumption and plasma column stabilization of such torches are low

The hollow copper electrode design is shown in Fig 43(b) The electrode material, usually copper alloyed with chromium or zirconium, is intensively water cooled Copper is a field emitter of electrons (as opposed to tungsten, a thermionic emitter) because its boiling point is substantially below that required for thermionic electron emission Current densities at the attachment points on cold cathodes are double those for thermionic emission on hot cathodes Cathode spots are of high intensity, rapidly moving on the surface of the cold cathode and causing erosion by locally vaporizing copper For this reason, the polarity is usually reversed in plasma torches with cold electrodes The rear electrode acts as the anode, and the workpiece or melt is the cathode The anodic arc attachment is smoother, and erosion is reduced The plasma-forming gas is blown in tangentially between the rear electrode and the nozzle to create a swirl or vortex to stabilize the arc and to rotate the arc attachment in the rear electrode With the movement of the arc attachment, a substantially lower power density can be achieved in the rear electrode

Electrode life of 1000 h has been demonstrated in torches using argon as the plasma gas All inert and reactive gases, or mixtures of these, can generally be used in such torches, but the influence of the gases on electrode erosion varies The gas consumption of vortex-stabilized torches is higher than that of other designs

The efficiency of plasma torches is less affected by design principles than by process parameters and by the type of

plasma gas used Efficiency is defined as Pw/Pt, where Pt is the electrical power input of the torch and Pw is the thermal power applied to the workpiece or melt The efficiency of a 650 kW torch measured in a cold furnace ranges from approximately 20 to 40%, depending on the plasma gas used, furnace pressure, and the distance between torch and workpiece

The highest efficiency is obtained with helium as the plasma gas and a short distance between material and torch The power loss in torches with hollow water-cooled electrodes is approximately 20% Remaining losses are due to radiation and convection of the plasma column between torch and workpiece These losses are primarily a function of the distance between torch and workpiece, but radiative losses of the plasma column also increase with pressure

Reference cited in this section

55 E Pfender, M Boulos, and P Fauchais, Methods and Principles of Plasma Generation, in Plasma Technology in

Metallurgical Processing, J Feinman, Ed., Iron and Steel Society, 1987, p 27-47

Furnace Equipment

Plasma melting furnaces usually consist of double-wall, vacuum-tight, water-cooled constructions that can withstand the radiative and convective heat transfer from the plasma column Furnaces also usually have material feeding systems These can consist of vibratory or rotary feeders with bins (for bulk materials) or bar feeding systems If the furnace is to operate continuously, the feed material must be "locked in" to prevent disturbance of the furnace atmosphere

The furnace must also have a vacuum pump so that it can be evacuated before backfilling If the furnace is operated under reduced pressure, an offgas pump with an on-line gas control valve is required

Atmosphere Control

Plasma melting furnaces are usually operated under slightly positive pressure to prevent the potential atmospheric contamination by oxygen and nitrogen However, state-of-the-art furnaces are vacuum tight and can be operated at pressures between 5 and 200 kPa (50 and 2000 mbar, or 38 and 1500 torr) Vacuum tightness is essential because the back diffusion of oxygen, nitrogen, and moisture through small leaks can be easily demonstrated even with positive pressure in the furnace Plasma torches are normally operated with argon plasma gas A typical gas purity of 99.999% indicates 10 ppm gaseous impurities and corresponds to an absolute pressure of 1 Pa (0.01 mbar, or 0.0075 torr) in a vacuum process

Additional sources of atmospheric contamination are the moisture and gases in the feed material, desorption of the furnace wall, and leaks in the furnace construction In the case of metals with high affinities for oxygen or nitrogen, these gases are totally absorbed by the melt and will be found as oxide or nitride inclusions in the product

To obtain a clean initial furnace atmosphere, the vacuum-tight plasma furnace is evacuated with a mechanical pump to final pressure (typically about 1 Pa, or 0.01 mbar, or 0.0075 torr) The furnace walls can be degassed with the help of warm water running between the double walls of the furnace to enhance moisture desorption The furnace is then backfilled with high-purity inert gas up to 1 kPa (10 mbar, or 7.5 torr), evacuated again to 1 Pa (0.01 mbar, or 0.0075 torr), and backfilled to operating pressure This procedure dilutes residual gas impurities by a factor of 1000

Processes

Application of Plasma Melting Processes. The selection of a suitable heat source for melting, remelting, casting, and atomizing reactive metals, titanium, and superalloys is of great interest to the manufacturer and user of plasma furnace equipment Skull melting processes in water-cooled copper crucibles are often the only melting technology for these metals (see the section "Vacuum Arc Skull Melting and Casting" in this article) Plasma torches, however, are the only nonconsumable heat sources for melting under high inert gas pressures in skull crucibles High pressure is an

essential requirement for preventing the selective evaporation of alloying elements that are characterized by high activity coefficients and/or vapor pressures, such as chromium and manganese in superalloys and aluminum in titanium alloys

Various plasma melting processes have been developed, including plasma consolidation, plasma arc remelting, plasma cold hearth melting, and plasma casting Plasma cold hearth melting and plasma cold crucible casting are thought to have the most potential for industrial application Table 10 lists data on feed materials, operating parameters, and products of furnaces for some of these processes

Table 10 Characteristics and operating parameters of furnaces for plasma melting processes

Typical

industrial

furnace

Feed material

Product form and size,

mm (in.)

Typical melt rate, kg/h (lb/h)

Specific power consumption,

kW · h/kg

Plasma gas

Typical furnace pressure

Torch design principle and power

Ingot 355-430 (14-17) diam,

3000 (120) long; density

>90%

250-260 (550- 570)

1.4-1.6 Argon ≥Atmospheric Tungsten tip

with external ignition bar; 6 ×

Ingot 700 (27.5) diam,

3800 (150) long; density 98%

250-670 (550- 1480)

0.89-2.4 Argon ≥Atmospheric Hollow copper

electrode with graphite liner;

600 kW

57

Plasma arc remelting furnaces

Soviet type I Titanium,

titanium alloys

Ingot 125 (5) diam, 500 (20) long

50-70 (110- 155)

heat-Ingot 150-200 (6-8) diam,

1200 (47) long; slabs 70

× 300 × 1200 (2.8 × 12 × 47)

90-135 (200- 300)

Ingots 650 (25.5) diam, 1500-2300 (60-90) long

400-900 (880- 1985)

Slab 245 ×

1125 × 2450 (9.6 × 44.3 ×

330 (730)

9 1-13 Pa

(0.01-0.13 mbar)

Six hollow cathode torches;

700 kW on hearth; 1170 kW

59

Steel, Japan 96.5) on crucible

Retech

pilot-scale furnace,

United States

Titanium scrap and sponge, superalloys

Ingot 195 (7.7) diam

90-227 (200- 500)

1.3-3 Helium;

gas recycling

≥Atmospheric Hollow copper

electrode, vortex stabilized; 200

kW on hearth,

100 kW on crucible

60

Plasma cold crucible melting and casting furnaces

Leybold AG Titanium

scrap and sponge, superalloys

Casting ingots

380 (840)

gas recycling

50 kPa (500 mbar)

Hollow copper electrode, vortex stabilized; 650

kW

61

Plasma Consolidation. Consolidation implies that low-density feed material is converted into a high-density product Plasma consolidation is used for titanium scrap and sponge to produce an electrode for further remelting in vacuum arc remelting (VAR) furnaces The material is usually fed directly into a water-cooled copper crucible and melted by plasma heat from one or several plasma torches The water-cooled copper baseplate of the crucible is continuously withdrawn The shape and depth of the liquid pool in the crucible depend on feed rate and power input Under these conditions, 100% melting cannot be ensured; some material can be trapped at the liquid/solid interface before melting Figure 44 shows two typical plasma consolidation furnaces

Fig 44 Plasma consolidation furnaces (a) Japanese furnace with six torches and total power of 540 kW

Source: Ref 60 (b) Retech furnace in the United States with one torch Source: Ref 57

Figure 45 shows ingot density as a function of the raw material feed rate and power consumption in the plasma consolidation of titanium Plasma consolidation processes can reduce manufacturing costs for the preparation of titanium VAR electrodes, which are usually plasma melted from pressed sponge or chip compacts

Fig 45 Relationship among raw material feeding rate, ingot bulk density, and specific power consumption in

the plasma consolidation of titanium Source: Ref 56

With plasma consolidation, metal chlorides can be removed from titanium sponge to sufficiently low levels Tungsten carbide tool tips in titanium chips cannot be removed and eventually form high-density inclusions in the product Therefore, the conventional x-ray inspection of scrap is still necessary Low-density inclusions such as titanium nitride particles have not been found after plasma consolidation of contaminated titanium scrap and subsequent VAR remelting (Ref 57) Plasma consolidation has also been successfully applied to the processing of titanium aluminide (TiAl), a low-density alloy with excellent high-temperature service properties (Ref 57)

Plasma Arc Remelting. An electrode that has already been melted by a primary melting process such as electric arc or induction melting can be plasma arc remelted into a water-cooled withdrawal crucible The major objectives for plasma arc remelting are:

• To obtain directional solidification without changing the chemical composition of the feed material

• To improve cleanliness by removal, size reduction, shape control, and redistribution of inclusions

• To lower gas impurity content

• To alloy with nitrogen

Plasma arc remelting is used most frequently in the Soviet Union to produce high-temperature alloys, bearing steels, superalloys, and titanium alloys Plasma arc remelting can be advantageous for the nitrogen alloying, deoxidation, or desulfurization of nonreactive alloys An example of successful reactive melting is the nitrogen alloying of steel during plasma arc remelting with argon-nitrogen mixtures as the plasma gas (Ref 58) Figure 46 shows nitrogen concentrations

in various metals as a function of nitrogen partial pressure in the gaseous phase after plasma arc remelting The data indicate that excited molecules of nitrogen in the plasma react with the metal (Ref 58)

Fig 46 Nitrogen concentrations in various metals as a function of the partial pressure of nitrogen after plasma

arc remelting (a) Iron (b) 16-25-6 stainless steel (c) Austenitic stainless steel Curves A and B, induction heating; minimum and maximum nitrogen concentrations, respectively Curve C, plasma arc remelting Source: Ref 58

In plasma cold hearth melting, bars, sponge, or scrap can be continuously fed into a water-cooled copper trough, usually called a hearth, and melted by means of plasma heat The liquid metal flows over the lip of the hearth into a withdrawal mold The plasma cold hearth melting process is adapted from the electron beam cold hearth refining process, which is a well-developed refining and casting method for superalloys Plasma heat, instead of an electron beam, is employed to minimize selective evaporative losses that occur during electron-beam melting in the pressure range of 0.01

Figure 47 shows a pilot plasma cold hearth furnace (Fig 47a) and the only industrial-scale plasma cold hearth melting furnace currently in operation (Fig 47b) The plasma heat source of the industrial furnace is unconventional in that there

is no mechanism provided for torch movement The torches, called plasma electron beam guns, use hollow tantalum cathodes and operate at pressures of 1 to 10 Pa (0.01 to 0.1 mbar, or 0.0075 to 0.075 torr) The unusually high specific power consumption of this furnace reflects the relatively low efficiency of the heat source used

Fig 47 Two plasma cold hearth melting furnaces (a) Pilot-scale furnace with two tiltable plasma torches Total

power: 400 kW (b) ULVAC plasma beam furnace with six fixed plasma beam guns Total power: 2000 kW Source: Ref 57

Figure 48 shows a modern plasma cold hearth melting furnace concept with a double-wall, water-cooled, vacuum-tight furnace chamber and three plasma torches for the production of ingots 710 mm (28 in.) in diameter Two torches serve the hearth, and the third the ingot mold The torches use hollow water-cooled copper electrodes The high gas consumption of this torch design means that an effective gas recycling system must be provided, especially if helium is used as the plasma gas

Fig 48 Plasma cold hearth melting concept for the recycling of titanium scrap

The operating principle of a gas-recycling system is shown in Fig 49 Key technology in this system includes reactive filters, oil-free high-capacity pump sets, and compressors If necessary, the recycling system can be supplemented with cleaning systems for oxygen, carbon dioxide, nitrogen, hydrogen, and moisture

Fig 49 Schematic of plasma gas-recycling system 1, plasma furnace; 2, off-gas heat exchanger; 3, cyclone

separator; 4, mechanical filter; 5, motor-actuated valve; 6, vacuum pump set; 7, chemical filter; 8, fine filter;

9, buffer; 10, compressor; 11, heat exchanger; 12, gas storage tank

Plasma cold crucible casting is an alternative process to vacuum arc skull melting and casting or electron beam skull melting and casting, and it can overcome some of the disadvantages of these processes The disadvantages of vacuum arc skull melting in the melting and casting of titanium and superalloys are:

• Control of superheating is not possible; increased power results in increased melting rate instead of superheating

• Expensive electrodes must be used instead of the high-quality unconsolidated revert material that can be used for plasma cold crucible casting

For small batches of superalloys (up to 8 kg, or 18 lb), electron beam skull casting is a well-established technique Electron beam melting and casting is rarely used for large castings of either superalloys or titanium alloys This is because

of the shallow melt pool in the crucible due to lack of stirring action and the potential for selective evaporation of alloying elements during melting

Figure 50 shows a sectional view of an industrial-size plasma cold crucible casting furnace The furnace consists of a cylindrical, horizontally installed mold chamber with lockchambers at both sides and the melt chamber on top of the mold chamber A programmable torch-moving device is installed on top of the melt chamber.The pressure in the furnace during the melting operation can be preselected; pressures typically range from 30 to 50 kPa (300 to 500 mbar, or 225 to 375 torr) The melting operation can be started at every pressure in this range Solid material can be charged continuously into the water-cooled copper crucible, and additional material can be fed into the crucible during melting The melt is poured into stationary molds or centrifugal casting molds situated in the mold chamber The pouring stream can be superheated during casting

Fig 50 Industrial-size plasma cold crucible casting furnace with 650 kW torch power

The typical casting weight of this furnace is 27 kg (60 lb) of titanium with a 650 kW plasma torch To achieve a deep molten pool, a high heat flow from the plasma torch to the melt must be established in combination with high stirring velocity in the melt to enhance the effective heat transfer coefficient The stirring action in the melt can be regulated by controlling current, gas flow, and torch-to-melt distance Continuous feeding of the metal to be melted is required in order

to achieve a deep pool

In contrast to vacuum arc skull melting, the crucible need not be inspected after each melt The power densities in the plasma attachment area are adjusted to avoid damaging the crucible if it is inadvertently exposed to the plasma for short periods of time

References cited in this section

56 T Yagima et al., Development of the Plasma Progressive Casting Process and Its Application for Titanium Melting,

in Titanium 1986: Products and Applications, Vol II, Proceedings of the Technical Program from the 1986

International Conference, Titanium Development Association, 1987, p 985-993

57 S Stocks and D Hialt, Plasma Consolidation of Large Diameter Titanium Electrodes, in Titanium 1986: Products

and Applications, Vol II, Proceedings of the Technical Program from the 1986 International Conference, Titanium

Development Association, 1987, p 918-927

58 G.K Bhat, Plasma Arc Remelting, in Plasma Technology in Metallurgical Processing, J Feinman, Ed., Iron and Steel Society, 1987, p 163-174

59 K Murase, T Suzuki, T Kijima, H Takei, and Y Yoneda, Production of Titanium Slab Ingot in Vacuum Plasma

Electron Beam Furnace, in Proceedings of the Electron Beam Melting and Refining State of the Art 1986 Conference,

Bakish Materials Corporation, 1986, p 184-194

60 R.C Eschenbach, in Proceedings of the 1986 Vacuum Metallurgy Conference on Specialty Metals Melting and

Processing (Pittsburgh, PA), Iron and Steel Society, 1986

61 G Sick, in Proceedings of the 1986 Vacuum Metallurgy Conference on Specialty Metals Melting and Processing (Pittsburgh, PA), Iron and Steel Society, 1986, p 179-186

References

Vacuum Induction Melting (VIM)

1 J.W Pridgeon et al., in Superalloys Source Book, American Society for Metals, 1984, p 201-217

2 W.B Kent, Int Voc Sci Technol., Vol 11 (No 6), 1974, p 1038-1046

3 P.P Turillon, in Transactions of the Sixth International Vacuum Metallurgy Conference (Boston), American Vacuum

Society, 1983, p 88

4 G.A Simkovich, Int Met., Vol 253 (No 4), 1966, p 504-512

5 H Katayam et al., in Proceedings of the Seventh International Conference on Vacuum Metallurgy (Tokyo), The Iron

and Steel Institute of Japan, 1982, p 933-940

6 A Choudhury et al., World Steel and Metalworking Manual, Vol 9, 1987-1988, p 1-6

7 P Hupfer, Fachberichte Hüttenpraxis Metallverarbeitung, 1986, p 773-781

8 O Kamado et al., Method of Producing Electrical Conductor, European Patent 0121152, 1986

9 J.G Krüger, Proceedings of the Fifth International Vacuum Metallurgy Conference (Munich), 1976, p 75-80

Electroslag Remelting (ESR)

10 R.K Hopkins, Method and Apparatus for Producing Cast Metal Bodies, U.S Patent 2,380,238, 1945

11 B.E Raton, B.I Medower, and W.E Raton, Ein neues Verfahren Das Elektroabgiessen von Blöcken, Bull tech Informazil NTO, No 1, Maschprom., Kiew 1956

12 G Hoyle, Electroslag Processes, Applied Science, 1983

13 W Holzgruber and E Plöckinger, Metallurgische und verfahrenstechnische Grundlagen des

Elektroschlacke-14 M Wahlster, A Choudhury, and K Forch, Einfluss des Umschmelzens nach Sonderverfahren auf Gefüge und

einige Eigenschaften von Stählen, Stahl Eisen, Vol 88, 1968, p 1193-1202

15 M Wahlster and A Choudhury, Beitrag zum Elektroschlacke-Umschmelzen von Stählen, Rheinstahl Technik, Vol

5, 1967, p 31-37

16 M Wahlster, G.H Klingelhöfer, and A Choudhury, Neue metallurgusche und technologische Ergebnisse einer 10

t ESU-Anlage, Radex-Rundsch., Vol 2, 1970, p 99-111

17 G.H Klingelhöfer, A Choudhury, and E Königer, Etude comparée des caractéristique des aciers e aborés à l'air et des aciers refondus selon le procédé sous laiter electroconducteur (ESR) pour les cylindres de laminage à froid et

les cylindres calandreurs, Rev Metall., Vol 67, 1970, p 512-522

18 H.J Klingëlhofer, P Mathis, and A Choudhury, Ein Beitrag zur Metallurgie des

Elektroschlackeumschmelzverfahrens, Arch Eisenhüttenwes., Vol 42, 1971, p 299-306

19 H Löwenkamp, A Choudhury, R Jauch, and F Regnitter, Umschmelzen von Schmiedeblöcken nach dem

Elektro-Schlacke-Umschmelzverfahren und dem Vakuumlichtbogenofenverfahren, Stahl Eisen, Vol 93, 1973, p 625-635

20 R Jauch, A Choudhury, H Löwenkamp, and F Regnitter, Herstellung grosser Schmiedeblöcke nach dem

Elektroschlacke-Umschmelzverfahren, Stahl Eisen, Vol 95, 1975, p 408-418

21 A Choudhury, R Jauch, and H Löwenkamp, Primärstruktur und Innenbeschaffenheit herkömmlicher und nach dem Elektro-Schlacke-Umschmelzverfahren hergestellte Blöcke mit einem Durchmesser von 2000 and 2300 mm

BA für Forschung und Technologie, Kennzeichen NTS 23

22 R Jauch, A Choudhury, F Tince, and H Steil, Herstellung und Verarbeitung von ESU-Blöcken bis zu 160 t, Stahl Eisen, Vol 101, 1981, p 41-44

23 A Choudhury, R Jauch, and F Tince, Low Frequency Electroslag Remelting of Heavy Ingots With Low

Aluminium Content, in Proceedings of the Sixth International Conference on Vacuum Metallurgy (San Diego),

1979, p 785-794

24 A Choudhury, R Jauch, H Löwenkamp, and F Tince, Application of the Electroslag Remelting Process for the

Production of Heavy Turbine Rotors from 12 %-Cr-Steel, Stahl Eisen, Vol 97, 1977, p 857-868

25 Ch Kubisch, Druckstickstoffstähle eine neue Gruppe von Edelstählen Berg und Hüttenmännische Monatshefte,

1971, p 84-88

26 A Choudhury, "Vacuum Electroslag Remelting A New Process for Better Products," Paper presented at the Vacuum Metallurgy Conference, Pittsburgh, PA, June 1986

Vacuum Arc Remelting (VAR)

27 W.A Tiller and J.W Rutter, Can J Phys., Vol 311, 1956, p 96

28 W.H Sutton, in Proceedings of the Seventh International Vacuum Metallurgy Conference (Tokyo), The Iron and

Steel Institute of Japan, 1982, p 904-915

29 J Preston, in Transactions of the Vacuum Metallurgy Conference, American Vacuum Society, 1965, p 366-379

30 A.S Ballentyne and A Mitchell, Iron-making Steelmaking, Vol 4, 1977, p 222-238

31 S Sawa et al., in Proceedings of the Fourth International Vacuum Metallurgy Conference (Tokyo), The Iron and

Steel Institute of Japan, 1974, p 129-134

32 J.W Troutman, in Transactions of the Vacuum Metallurgy Conference, American Vacuum Society, 1968, p

599-613

33 R Schlatter, Giesserei, Vol 61, 1970, p 75-85

34 A Mitchell, in Proceedings of the Vacuum Metallurgy Conference, Pittsburgh, PA, 1986, p 55-61

35 F.J Wadier, in Proceedings of the Vacuum Metallurgy Conference, Pittsburgh, PA, 1984, p 119-128

36 J.W Pridgeon, F.M Darmava, J.S Huntington, and W.H Sutton, in Super-alloys Source Book, American Society

for Metals, 1984

Electron Beam Melting and Casting

37 R.E Lüders, Tantalum Melting in a 800 kW EB Furnace, in Proceedings of the Bakish Conference on Electron

Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1983, p 230-244

38 J.A Pierret and J.B Lambert, Operation of Electron Beam Furnace for Melting Refractory Metals, in Proceedings

of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1984, p 208-218

40 R Hähn and J Krüger, Refining of Vanadium Aluminium Alloys to Vanadium 99.9% by EB Melting, in

Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1986, p

53-67

41 J Lambrecht, D Rumberg, and K.H Werner, Stand der Technologie der Stahlerzeugung im Elektronenstrahlofen

mit Blockmassen bis 100 t, Neue Hütte, Vol 10, Oct 1984

42 C.E Shamblen, S.L Culp, and R.W Lober, Superalloy Cleanliness Evaluation Using the EB Button Melt Test, in

Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1983

43 F Hauner, H Stephan, and H Stumpp, Ergebnisse bei Elektronenstrahl-Schmelzen von gerichtet erstarrten

Knopfproben zur Identifizierung von nichtmetallischen Einschlüssen in hochreinen Superlegierungen, Metall., Vol 2

(No 40), 1986, p 2-7

44 C.d'A Hunt and H.R Smith, Electron Beam Processing of Molten Steel in Cold Hearth Furnace, J Met., Vol 18,

1966, p 570-577

45 H.R Harker, Electron Beam Melting of Titanium Scrap, in Proceedings of the Bakish Conference on Electron Beam

Melting and Refining (Reno, NV), R Bakish, Ed., 1983, p 187-190

46 C.d'A Hunt, H.R Smith, and B.C Coad, The Combined Induction and Electron Beam Furnace for Steel Refining

and Casting, in Proceedings of the Vacuum Metallurgy Conference (Pittsburgh, PA), June 1969, p 1-22

47 H.R Harker, The Present Status of Electron Beam Melting Technology, in Proceedings of the Bakish Conference on

Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1986, p 3-7

48 C.d'A Hunt, J.C Lowe, and T.H Harrington, Electron Beam, Cold Hearth Refining for the Production of Nickel

and Cobalt Base Superalloys, in Proceedings of the Bakish Conference on Electron Beam Melting and Refining

(Reno, NV), R Bakish, Ed., 1984, p 295-304

49 H Stephan, R Schumann, and H.J Stumpp, Production of Superclean Fine-Grained Superalloys for Improvement

of the Workability and Engine Efficiency by EB Melting and Refining Methods, in Proceedings of the Eighth International Conference on Vacuum Metallurgy (Linz), 1985, p 1219-1309

50 H Stephan, Production of Ingots and Cast Parts From Reactive Metals by Electron Beam Melting and Casting, in

Proceedings of the Third Electron Beam Processing Seminar (Stratford, UK), 1974, p 1b1-1b69

51 H Ranke, V Bauer, W Dietrich, J Heimerl, and H Stephan, Melting and Evaporation With the Newly Developed

Leybold-Heraeus 600 kW EB Gun at Different Pressure Levels, in Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1985

52 M Krehl and J.C Lowe, Electron Beam Cold Hearth Refining for Superalloy Revert for Use in Foundry Production,

in Proceedings of the Bakish Conference on Electron Beam Melting and Refining (Reno, NV), R Bakish, Ed., 1986,

p 286-296

53 M Romberg, R Schumann, H Stephan, and H Stumpp, Electron Beam Melting and Refining of Superalloys for

Ingot and Barstick Production, in Proceedings of the Bakish Conference on Electron Beam Melting and Refining

(Reno, NV), R Bakish, Ed., 1986

54 J Mayfield, Computer-Controlled Production Gains, Aviat Week Space Technol., 3 Dec 1979, p 1-5

Plasma Melting and Casting

55 E Pfender, M Boulos, and P Fauchais, Methods and Principles of Plasma Generation, in Plasma Technology in

Metallurgical Processing, J Feinman, Ed., Iron and Steel Society, 1987, p 27-47

56 T Yagima et al., Development of the Plasma Progressive Casting Process and Its Application for Titanium Melting,

in Titanium 1986: Products and Applications, Vol II, Proceedings of the Technical Program from the 1986

International Conference, Titanium Development Association, 1987, p 985-993

57 S Stocks and D Hialt, Plasma Consolidation of Large Diameter Titanium Electrodes, in Titanium 1986: Products

and Applications, Vol II, Proceedings of the Technical Program from the 1986 International Conference, Titanium

Development Association, 1987, p 918-927

58 G.K Bhat, Plasma Arc Remelting, in Plasma Technology in Metallurgical Processing, J Feinman, Ed., Iron and Steel Society, 1987, p 163-174

59 K Murase, T Suzuki, T Kijima, H Takei, and Y Yoneda, Production of Titanium Slab Ingot in Vacuum Plasma

Electron Beam Furnace, in Proceedings of the Electron Beam Melting and Refining State of the Art 1986 Conference, Bakish Materials Corporation, 1986, p 184-194

60 R.C Eschenbach, in Proceedings of the 1986 Vacuum Metallurgy Conference on Specialty Metals Melting and

Processing (Pittsburgh, PA), Iron and Steel Society, 1986

61 G Sick, in Proceedings of the 1986 Vacuum Metallurgy Conference on Specialty Metals Melting and Processing

(Pittsburgh, PA), Iron and Steel Society, 1986, p 179-186

62 N.A Barcza, Application of Plasma Technology to Steel Processing, in Plasma Technology in Metallurgical

Processing, J Feinman, Ed., Iron and Steel Society, 1987, p 131-148

Degassing Processes (Converter Metallurgy)

Chairman: Gerhard Kienel, Leybold AG, West Germany

Introduction

CONVERTERS are used in secondary metallurgy to refine melts outside the primary metallurgical melting unit Various converters are available that apply the bottom blowing of oxygen-inert gas mixtures These bottom-blowing converters use different types, numbers, and arrangements of injection nozzles and the following gases:

• Argon as a cooling inert gas with a purity ranging from 85 to 99.99%

• Nitrogen as a cooling inert gas with a purity of 99 to 99.9%

• Argon-oxygen and nitrogen-oxygen mixtures in the case of the argon oxygen decarburization converter

• Dry air as a diluted reactive gas

• Oxygen as a reactive gas with a purity of 80 to 99.5%

• Steam as a cooling reactive gas

• Carbon dioxide as a diluted reactive gas

This article will review three converter designs These are the argon oxygen decarburization vessel; the oxygen top and bottom blowing converter, which is an extension of argon oxygen decarburization technology; and the vacuum oxygen decarburization converter Additional information on degassing processes can be found in the article "Degassing Processes (Ladle Metallurgy)" in this Volume

Argon Oxygen Decarburization (AOD)

Ian F Masterson, Union Carbide Corporation Linde Division

Argon oxygen decarburization (AOD) is a secondary refining process that was originally developed to reduce material and operating costs and to increase the productivity of stainless steel production In addition to its economic merits, argon oxygen decarburization offers improved metal cleanliness, which is measured by low unwanted residual element contents and gas contents; this ensures superior mechanical properties The AOD process is duplexed, with molten metal supplied from a separate melting source to the AOD refining unit (vessel) The source of the molten metal is usually an electric arc furnace or a coreless induction furnace Foundries and integrated steel mills utilize vessels ranging in nominal capacity from 1 to 160 Mg (1 to 175 tons)

Although the process was initially targeted for stainless steel production, argon oxygen decarburization is used in refining

a wide range of alloys, including:

• Stainless steels

• Tool steels

• Silicon (electrical) steels

• Carbon steels, low-alloy steels, and high-strength low-alloy steels

• High-temperature alloys and superalloys

Fundamentals

In the AOD process, oxygen, argon, and nitrogen are injected into a molten metal bath through submerged, side-mounted tuyeres The primary aspect of the AOD process is the shift in the decarburization thermodynamics that is afforded by blowing with mixtures of oxygen and inert gas as opposed to pure oxygen References 1, 2, and 3 contain detailed discussions of the thermodynamics and the dilution principle and its influence on the refining of chromium-bearing steel

To understand the AOD process, it is necessary to examine the thermodynamics governing the reactions that occur in the refining of stainless steel, that is, the relationship among carbon, chromium, chromium oxide (Cr3O4), and carbon monoxide (CO) The overall reaction in the decarburization of chromium-containing steel can be written as:

a a

where a and P represent the activity and partial pressure, respectively

At a given temperature, there is a fixed, limited amount of chromium that can exist in the molten bath that is in equilibrium with carbon By examining Eq 2, one can see that by reducing the partial pressure of CO, the quantity of chromium that can exist in the molten bath in equilibrium with carbon increases The partial pressure of CO can be reduced by injecting mixtures of oxygen and inert gas during the decarburization of stainless steel Figure 1 illustrates the relationship among carbon, chromium, and temperature for a partial pressure of CO equal to 1 and 0.10 atm (1000 and

100 mbar, or 760 and 76 torr) The data shown in Fig 1 indicate that diluting the partial pressure of CO allows lower carbon levels to be obtained at higher chromium contents with lower temperatures

Fig 1 Carbon-chromium equilibrium curves

In refining stainless steel, it is generally necessary to decarburize the molten bath to less than 0.05% C Chromium is quite susceptible to oxidation; therefore, prior to the introduction of the AOD process, decarburization was accomplished by withholding most of the chromium until the bath had been decarburized by oxygen lancing After the bath was fully decarburized, low-carbon ferrochromium and other low-carbon ferroalloys were added to the melt to meet chemical specifications

Dilution of the partial pressure of CO allows the removal of carbon to low levels without excessive chromium oxidation This practice enables the use of high-carbon ferroalloys in the charge mix, avoiding the substantially more expensive low-carbon ferroalloys Figure 2 compares the refining steps in the two processes

Fig 2 Composition changes in refining type 304-L stainless steel using electric arc furnace practice and argon

oxygen decarburization

Equally important, however, are the processing advantages offered by submerged blowing These include excellent slag/metal and gas/metal contact, superior decarburization kinetics, and 100% utilization of the injected oxygen by reaction with the bath In addition, submerged injection allows the accurate control of end point nitrogen and the removal

of dissolved gases (nitrogen and hydrogen) and nonmetallic inclusions

References cited in this section

1 F.D Richardson and W.E Dennis, Effect of Chromium on the Thermodynamic Activity of Carbon in Liquid

Iron-Chromium-Carbon Metals, J Iron Steel Inst., Nov 1953, p 257-263

2 R.J Choulet, F.S Death, and R.N Dokken, Argon-Oxygen Refining of Stainless Steel, Can Metall Q., Vol 10 (No 2), 1971, p 129-136

3 R.B Aucott, D.W Gray, and C.G Holland, The Theory and Practice of the Argon-Oxygen Decarburizing Process, J

W Scot Iron Steel Inst., Vol 79 (No 5), 1971-1972, p 98-127

Equipment

The processing vessel consists of a refractory-lined steel shell mounted on a tiltable trunnion ring (Fig 3) With a removable, conical cover in place, the vessel outline is sometimes described as pear shaped Several basic refractory types and various quality levels of the refractories have gained widespread acceptance (Ref 4, 5) Dolomite refractories are used

in most AOD installations; magnesite chromium refractories are predominant in small (<9 Mg, or 10 ton) installations and are used almost exclusively in Japan

Fig 3 Schematic of argon oxygen decarburization vessel

As seen in Fig 3, process gases are injected through submerged, side-mounted tuyeres The number and relative positioning of tuyeres are determined in part by vessel size, range of heat sizes, process gas flow rates, and the types of alloys refined Process gases are oxygen, nitrogen, argon, and in some cases carbon dioxide (CO2) The most recent AOD installations include the use of top-blown oxygen (one-third of all AOD installations have this capability) Oxygen top- and bottom-blowing converters are described in the discussion "Oxygen Top and Bottom Blowing" in this section

The specific volume of AOD vessels ranges from 0.4 to 0.8 m3/Mg (16 to 32 ft3/ton) The gas control system supplies the process gases at nominal rates of 0.5 to 3 m3/min/Mg (15,000 to 6000 ft3/h/ton) The system accurately controls the flow rates and monitors the amount of gas injected into the bath to enable the operator to control the process and keep track of the total oxygen injected

Normally, a shop has three interchangeable vessels At any given time, one of the vessels is in a tiltable trunnion ring refining steel, a second vessel is at a preheating station, and the third vessel is being relined The vessel in the trunnion ring can be replaced with a preheated vessel in less than 1 h

The control of process gases, vessel activities, and ancillary equipment can range from manual to fully automated Most installations are equipped with a computer to assist in process control by calculating the required amount of oxygen as well as alloying additions Some installations have computer control systems capable of sending set points and flow rates

to the gas control systems

References cited in this section

4 D.A Whitworth, F.D Jackson, and F.F Patrick, Fused Basic Refractories in the Argon-Oxygen Decarburization

Process, Ceram Bull., Vol 53 (No 11), 1974

5 D Brosnan and R.J Marr, "The Use of Direct Burned Dolomite Brick in the AOD Vessel," Paper presented at the Electric Furnace Conference, Detroit, MI, Iron and Steel Society, Dec 1977

Processing

Stainless Steels. Charge materials (scrap and ferroalloys) are melted in the melting furnace The charge is usually

melted with the chromium, nickel, and manganese concentrations at midrange specifications The carbon content at meltdown can vary from 0.50 to 3.0%, depending on the scrap content of the charge Once the charge is melted down, the heat is tapped, and the slag is removed and weighed prior to charging the AOD vessel

In the refining of stainless steel grades, oxygen and inert gas are injected into the bath in a stepwise manner The ratio of oxygen to inert gas injected decreases (3:1, 1:1, 1:3) as the carbon level decreases Once the aim carbon level is obtained,

a reduction mix (silicon, aluminum, and lime) is added If extra-low sulfur levels are desired, a second desulfurization can

be added Both of these steps are followed by an argon stir After reduction, a complete chemistry sample is usually taken and trim additions made following analysis Figure 4 illustrates the relationships among carbon, chromium, temperature, and the various processing steps for refining a typical type 304 stainless steel using argon oxygen decarburization

Fig 4 Schematic of stainless steel refining cycle for small (<27 Mg, or 30 ton) vessels showing the relationship

between carbon and chromium contents as a function of time and temperature

Another critical aspect of AOD refining is the ability to predict when to change from nitrogen to argon to obtain the aim nitrogen specification Table 1 lists the process parameters, carbon removal efficiencies, and carbon removal rates for the AOD refining of type 304 stainless steel The point during refining when the oxygen-to-inert-gas ratio is lowered is based

on carbon content and temperature The ratios and carbon switch points are designed to provide optimum carbon removal efficiency without exceeding a bath temperature of 1730 °C (3150 °F)

Table 1 Processing parameters for AOD refining of type 304 stainless steel

Total gas flow rate: ~1.5 m3/min/ton (3500 ft3/h/ton)

Gas ratios Partial pressures, atm Temperature Carbon level, %

O 2 N 2 Ar Ar N 2 CO

CRE (a) ,

%

CRR (b) , ppm/min

(a) CRE, carbon removal efficiency

(b) CRR, carbon removal rate

Carbon and Low-Alloy Steels. The refining of carbon and low-alloy steels involves a two-step practice: a carbon removal step, followed by a reduction/heating step The lower alloy content of these steels eliminates the need for injecting less than a 3:1 ratio of oxygen to inert gas Once the aim carbon level is obtained, carbon steels are processed similarly to stainless steels Figure 5 illustrates carbon content and temperature relationship for the AOD refining of carbon and low-alloy steels Because the alloy content of these grades of steel is substantially lower than that of stainless steel and because the final carbon levels are generally higher, there is no thermodynamic or practical reason for using an oxygen, inert-gas ratio of less than 3:1

Fig 5 Schematic of carbon and low-alloy steel argon oxygen decarburization refining

Oxidation measurements indicate that all of the oxygen reacts with the bath and that none leaves the vessel unreacted By monitoring and recording the oxygen consumption during refining, very close control of end point carbon is achieved Because the oxygen and inert gases are introduced below the bath and at sonic velocities, there is excellent bath mixing and intimate slag/metal contact As a result, the reaction kinetics of all chemical processes that take place within the vessel are greatly improved

Decarburization. In both stainless and low-alloy steels, the dilution of oxygen with inert gas results in increased carbon removal efficiencies without excessive metallic oxidation In stainless grades, carbon levels of 0.01% are readily obtained

Chemistry Control. The excellent compositional control of AOD-refined steel is indicated in Table 2 for a ten-heat series of high-strength low-alloy steel The injection of a known quantity of oxygen with a predetermined bath weight enables the steelmaker to obtain very tight chemical specifications

Table 2 Composition control of AOD-refined high-strength low-alloy steel

Element Aim Mean Standard

deviation

Carbon 0.27 0.264 0.006

Manganese 0.80 0.83 0.03

Silicon 0.50 0.49 0.04

Nitrogen Control. Degassing in argon oxygen decarburization is achieved by inert gas sparging Each argon and CO bubble leaving the bath removes a small amount of dissolved nitrogen and hydrogen Final nitrogen content can be accurately controlled by substituting nitrogen for argon during refining Nitrogen levels as low as 25 to 30 ppm can be obtained in carbon and low-alloy steels, and 100 to 150 ppm N can be obtained in stainless steels The ability to obtain aim nitrogen levels substantially reduces the need to use nitrided ferroalloys for alloy specification, and this also minimizes the use of argon Hydrogen levels as low as 1.5 ppm can be obtained

Property Improvements. It has been well documented that AOD-refined steels exhibit significantly improved ductility and toughness, along with impact energy increases of over 50% (Ref 6, 7, 8) These improved properties result from a decrease in the number and size of inclusions The capability to produce low gas content steel with exceptional microcleanliness, along with alloy savings, is the primary factor for the growth of the AOD process for refining stainless, carbon, and low-alloy steels

References cited in this section

6 P.A Tichauer and L.J Venne, AOD, a New Process for Steel Foundries, Promises Better Properties, Less Repair, 33

Oxygen Top and Bottom Blowing

Oxygen top and bottom blowing (OTB), also referred to as combined blowing, is an extension of AOD technology for refining steel During OTB, oxygen is injected into the molten steel through a top lance during carbon removal, and inert gases, such as argon, nitrogen, or carbon dioxide, are injected with oxygen through submerged tuyeres or alternate forms

of gas injectors such as canned bricks, porous bricks, or thin pipes set in the refractory brick The top-blown oxygen can react with the bath, reducing refining times, or with CO The combustion of CO above the surface of the bath increases the thermal efficiency of the refining process and decreases the quantity of silicon and aluminum required for reduction of metallic oxides

If the top-blown oxygen system is designed so that more than 65% of the oxygen reacts with the bath, the system is referred to as hard blown; if less than 65% of the oxygen reacts with the bath, it is referred to as soft blown As a general guideline, a top-blown oxygen system installed in AOD vessels with a nominal capacity greater than 45 Mg (50 tons) will

be hard blown, smaller vessels will be soft blown In hard-blown top oxygen systems, the flow rate of oxygen through the top lance will be between 50 and 150% of the oxygen injected through tuyeres or injectors In soft-blown systems, the top oxygen flow rate is between 50 and 100% of the oxygen flow through the tuyeres or injectors

Because foundry AOD vessels have a nominal capacity less than 45 Mg (50 tons) the top-blown oxygen systems are designed to be soft blown This design maximizes the thermal efficiency of the smaller AOD vessels, reducing the quantity of fuel and reduction material required during refining Top-blown oxygen flow rates in foundry AOD vessels range from 50% of the bottom oxygen flow rate to 120%

More detailed information on OTB technology can be found in Ref 9 and 10 Figures 8 and 9 in the following section

"Vacuum Oxygen Decarburization (VODC)" compare OTB with various refining processes

References cited in this section

9 S.K Mehlman, Ed., Mixed Gas Blowing, Proceedings of the Fourth Process Technology Conference, Iron and Steel

Society of AIME, American Institute of Mechanical Engineers, 1984

10 L.G Kuhn, Ed., Mixed Gas Blowing A New Era of Pneumatic Steelmaking, Issue of Iron and Steel Maker, Aug

1983

Vacuum Oxygen Decarburization (VODC)

Wilhelm Burgmann, Leybold AG, West Germany

Vacuum oxygen decarburization can be carried out in a converter vessel (VODC) or in the ladle (VOD) The parameters that favor the choice of ladles as reaction vessels include the following:

• Tapping, treatment, and teeming are done in the same reaction vessel Thus, there are no temperature losses due to any final transfer of the melt, and the high level of cleanliness achieved during the treatment can be preserved up to teeming

• The use of electric power as an inexpensive energy source permits the highest flexibility in the melt shop The ladle unit can act as a time buffer between the melting unit and the casting stand

Information on VOD ladle technology can be found in the article "Degassing Processes (Ladle Metallurgy)" in this Volume

The primary reasons for selecting the converter as the VOD reaction vessel are as follows:

• Initial carbon contents are as high as possible, together with high oxygen blow rates

• Ease of deslagging and strong stirring action

• Shop restrictions such as a limited number of ladles, prohibited use of slide gates, and restricted crane capacity in the casting bay make converter technology more attractive

This section will review both the equipment and the processing parameters for vacuum oxygen decarburization, which is used for the production of low-alloy steels, stainless steels, and superalloys

Equipment and Processing Parameters

Equipment. Vacuum oxygen decarburization converters are similar to argon oxygen decarburization (AOD) and oxygen top blown (OTB) converters in terms of design and the tilting device used (AOD and OTB converters were discussed in the previous section "Argon Oxygen Decarburization (AOD)" in this article.) Bottom blowing is, however, restricted to the introduction of small amounts of inert gas through simple pipes, thus avoiding the special erosion-resistant refractory material used around tuyeres Flue gas handling is easier and is incorporated into the vacuum system In terms of vessel design, the conical converter top is closed by a vacuum hood with an oxygen lance feed through and vacuum addition lock (Fig 6)

Fig 6 Schematic layout of a vacuum oxygen decarburization converter

Because the VODC system is closed and no air enters the vessel, permanent control of the decarburization rate and the carbon level in the bath can be maintained and monitored with a flue gas analyzing device (Ref 11, 12) Pollution control for carbon monoxide (CO) and dust is also incorporated into the system

Processing Parameters. Figures 7 and 8 show the relationship among CO partial pressure, gas blow rates, and carbon

and chromium oxidation rates for the VODC, VOD, AOD, and OTB processes As shown, low chromium losses and high carbon removal rates are possible using vacuum oxygen decarburization converters The high carbon removal rates are due to the higher oxygen yield of the vacuum oxygen decarburization converter and the fact that oxygen dilution is not required The amount of bottom blown inert gas is only about 2 to 5% of the oxygen volume

Fig 7 Comparison of CO partial pressure for the AOD, VODC, and VOD processes

Fig 8 Carbon removal rates and chromium oxidation rates for the VODC, OTB, and AOD processes

With a specific oxygen blow rate of 0.7 m3/min/Mg (~25 ft3/min/ton), the decarburization rate exceeds 600 ppm/min at high carbon levels (Fig 8) In smaller VODC heats, higher blow rates of the order of 10 m3/min (~350 ft3/min) are maintained, leading to a carbon removal of 1500 ppm/min in 5 Mg (5.5 ton) units

Figures 7 and 8 demonstrate clearly the low chromium losses and the high ratio of carbon-to-chromium removal rate due

to the very low CO partial pressure achieved in the VODC process As in any vacuum process, the lowest carbon levels (from 50 to 500 ppm) are reached in stainless steel heats very quickly (6 to 10 times the carbon removal rate of AOD) without applying excessive heat and excessive amounts of inert gas As shown in Fig 8, temperatures of about 1650 °C (3000 °F) are common Inert gas volumes are generally less than about 40 L/min/Mg (1.5 ft3/min/ton)

Figure 9 shows the relationship between the oxygen blow rates and the inert gas addition for various refining techniques,

as well as the resulting chromium and carbon levels achieved at specific process steps Figure 9 also shows the extent to

which decarburization is accompanied by chromium oxidation Carbon and nitrogen levels below 100 ppm are achieved with corresponding chromium yields of over 95% before reduction and 98% after reduction (Ref 13, 14) Sulfur levels are

at 10 to 30 ppm for any grade of steel without any sacrifice in the degassing effect Hydrogen levels are below 2 ppm

Total gas flow Technique

m 3 /min/Mg ft 3 /min/ton

Inert gas, %

Lance bubbling equilibrium (LBE) 2.5 95 1

Argon oxygen decarburization (AOD) 1.0 40 20-90