Fracture at Elevated Temperatures
The constant load creep rupture test is the basis for design data for both creep strength (minimum creep rate or time to a specific creep strain) and failure (time to rupture). The various ways in which such data are presented, correlated, and extrapolated are addressed in subsequent sections.

At this point, it is appropriate to consider the processes leading to fracture. Plastic instability in ductile materials has already been reviewed. This process may lead directly to fracture in pure metals and contribute significantly to fracture in engineering materials at moderately high stresses. However, of much greater concern are the processes leading to intergranular fracture with reduced ductility at low stresses and high temperatures. Here again, many of the basic studies have been conducted on pure metals and solid-solution alloys.

Crack Nucleation and Morphology. Two types of cracking have been identified: wedge-shaped cracks emanating from grain-boundary triple points and the formation of cavities or voids on grain-boundary facets often oriented perpendicular to the applied tensile stress. Although much work continues to model the nucleation and growth of these cracks and cavities, there are uncertainties in the mechanism of nucleation and in the identification of a failure criterion.

Another major problem is the effect of temperature and stress on the extent of cracking at failure. Most theories assume that failure occurs at some critical cavity distribution or crack size. However, it has been shown that the extent of cavitation at failure or at any given fraction of the failure life is very sensitive to the test conditions. Thus cavitation damage at failure at a high stress may be comparable to damage in the very early stage of a test at low stress. For stress-change experiments, there is therefore a loading sequence effect on rupture life, which is discussed later in this article, for engineering alloys.

Embrittlement Phenomena. As pointed out previously, rupture life is primarily a measure of creep strength; fracture resistance would be identified better with a separate measure that reflects the concern with embrittlement phenomena that may lead to component failure. Most engineering alloys lose ductility during high-temperature service. This has been shown to be a function of temperature and strain rate so that there is a critical regime for maximum embrittlement. At a fixed strain rate, for example, ductility first decreases with increasing temperature. This is believed to be caused by grain boundaries playing an increasing role in the deformation process leading to the nucleation of intergranular cracks. At still higher temperatures, processes of recovery and relaxation at local stress concentrations lead to an improvement in ductility.

Environmental Effects
It has long been known that test environment may affect creep-rupture behavior. Until quite recently, however, the work has been largely empirical with creep tests being conducted in various atmospheres and differences noted in creep rates and rupture lives.

The effect on rupture life, in particular, was often less than a factor of ten in environments such as oxygen, hydrogen, nitrogen, carbon dioxide, and impure helium compared with vacuum. In many cases, it was not clear how inert the vacuum was, and little account was taken of specimen thickness. Often, effects on ductility were not reported, and there were very few studies of crack propagation.

Embrittling Effects of Oxygen. At about the same time that the ideas on environmental attack at an intergranular crack tip were being developed, it was also shown that short-term prior exposure in air at high temperature (greater than about 900°C, or 1650°F) could lead to profound ernbrittlement at intermediate temperatures (700 to 800°C). This was shown to be caused by intergranular diffusion of oxygen that penetrated on the order of millimeters in a few hours at 1000°C (1830°F).

Combined Effects of Oxygen and Carbon. Of special interest relative to the previous discussion of creep cavitation is the reaction between diffusing oxygen and carbon. In nickel, it was found that if this reaction were prevented, creep cavitation could not develop during creep tests. Prevention was achieved either by removing the carbon (decarburizing) or by applying an environmental protective diffusion or overlay coating.

Effect of Other Gaseous Elements. Hydrogen, chlorine, and sulfur may also cause embrittlement as a result of penetration. Sulfur is particularly aggressive in that it diffuses more rapidly and embrittles more severely than does oxygen. It is also frequently found in coal gasification and oil-refining processes as well as industrial gas turbines operating on impure fuel.

Creep Rupture Data Presentation
Laboratory creep tests are typically run between 100 and 10,000 h, although a few are run for shorter times (for example, for acceptance tests), and occasionally some testing is conducted for longer times. Since most high-temperature components are expected to last ten years or more, service stresses are obviously lower than those used in the longest creep tests to generate data for most of the alloys used.

Therefore, to provide data for creep rates and rupture lives that are appropriate for the setting of design stresses, it became necessary to develop methods for extrapolation. Over the years, a tremendous amount of effort has gone into optimizing methods of data extrapolation.

One of the major considerations in such procedures must be statistical issues, such as the best estimate of the stress associated with a given median life or creep rate, the use of stress or time as the dependent variable in the data fitting, the treatment of variability among heats of the same alloy, and the analysis of data with run-outs. All of these issues have been treated with considerable rigor and shown to be important relative not only to the proper interpretation of data, but to the proper design of experiments. In addition, there are different practices among testing laboratories that may have appreciable effects on results. These include specimen geometry, loading procedure, specimen alignment, furnace type, and temperature control.

Despite all these concerns regarding proper statistical treatment of data, a methodology has been developed based on time-temperature parameters that are now in widespread use. The approach may be used to achieve the following major design objectives:

* It allows the representation of creep rupture (or creep) data in a compact form, allowing interpolation of results that are not experimentally determined.
* It provides a simple basis for comparison and ranking of different alloys.
* Extrapolation to time ranges beyond those normally reached is straightforward.

Damage Accumulation and Life Prediction
Engineering procedures for life management of operating components assume that the material is progressively degraded or damaged as creep strain increases and operating time accumulates.

Damage may be in the form of precipitate changes that may result in softening (overaging) and reduced creep strength, or embrittlement and reduced resistance to fracture. The embrittlement may be due to segregation of harmful species, either from the interior or from the external environment, to interfaces, especially grain boundaries. Damage may also occur as a result of progressive intergranular cavitation and cracking, as previously described. Some of this damage may be reversed by suitable heat treatment or by hot isostatic pressing and may allow the possibility of component rejuvenation.

There are two basic approaches to using the concept of damage accumulation for life assessment:

* Based on a detailed knowledge of the operating conditions, including temperature and stress changes, the remaining life is estimated from the known original properties of the material of construction.
* Remaining life estimates are made using post-exposure measurements of microstructural changes, intergranular cavitation, or mechanical properties such as hardness, impact energy, or stress-rupture life.

Conclusions
Time-dependent deformation and fracture of structural materials are among the most challenging engineering problems faced by materials engineers. The critical role played by high-temperature energy-conversion machines in modern society attests to the remarkable success of the design methodology developed during last decades.

However, modern design needs, including accelerated evaluation and development of advanced materials, and improved remaining life assessment methods for operating equipment have identified some ways in which the methodology might be improved. It is desirable to decouple the creep strength and fracture resistance criteria. This could lead to new accelerated short-time testing in which the objective is not to attempt to incorporate microstructural evolution and damage-development in the test, as in the traditional long-time creep-to-rupture test. Rather, the accelerated test may be used to measure separately the consequences of these changes on creep strength and fracture resistance.

The generally neglected anelastic or time-dependent recoverable component of creep may be the dominant strain component in many service situations at low stresses and needs to be incorporated in design analysis. This is true for ceramics and metals as well as polymers. It may also provide, in some cases, a critical link between deformation and fracture.

Apart from nineteenth-century steam boilers, machines and equipment for high-temperature operation have been developed principally in the 20th century. Energy conversion systems based on steam turbines, gas turbines, high-performance automobile engines, and jet engines provide the technological foundation for modern society.

All of these machines have in common the use of metallic materials at temperatures where time-dependent deformation and fracture processes must be considered in their design. The single valued time-invariant strain associated with elastic or plastic design analysis in low-temperature applications is not applicable, nor is there in most situations a unique value of fracture toughness that may be used as a limiting condition for part failure. In addition to the phenomenological complexities of time-dependent behavior, there is now convincing evidence that the synergism associated with gaseous environmental interactions may have a major effect, in particular on high-temperature fracture.

Basic Concepts of Elevated-Temperature Design. Time-dependent deformation and fracture of structural materials at elevated temperatures are among the most challenging engineering problems faced by materials engineers. In order to develop an improved design methodology for machines and equipment operating at high temperatures, several key concepts and their synergism must be understood:

* Plastic instability at elevated temperatures
* Deformation mechanisms and strain components associated with creep processes
* Stress and temperature dependence
* Fracture at elevated temperatures
* Environmental effects.

The issues of interest from a design basis are the nature of primary creep, the validity of the concept of viscous steady-state creep, and the dependence of deformation on both temperature and stress. The simplest and most pervasive idea in creep of metals is an approach to an equilibrium microstructural and mechanical state. Thus a hardening associated with dislocation generation and interaction is countered by a dynamic microstructural recovery or softening. This process proceeds during primary creep and culminates in a steady-state situation.

Plastic Instability
A major issue in the tensile creep test is the role of plastic instability in leading to tertiary creep. Understanding of the nature of plastic instability for time-dependent flow has depended on the theory of Hart. He showed that the condition for stable deformation is:

γ + m ≥ 1

where:
m is the strain-rate sensitivity, and
γ, is a measure of the strain-hardening rate.

For steady-stale flow, γ is equal to 0. For constant stress tests, Burke and Nix concluded that flow must be unstable when steady state is reached according to Hart’s criterion but that macroscopic necking is insignificant and that the flow remains essentially homogeneous. They concluded that a true steady state does exist. Hart himself questioned the conclusions based on their analysis but did not rule out the possibility of a steady state for pure metals.

In a very careful experimental analysis, Wray and Richmond later concluded that the concept of a family of steady states is valid. Tests were performed in which two of the basic parameters (stress, strain rate, and temperature) are held constant. However, they reported the intrusion of nonuniform deformation before the steady state was reached. They also pointed out the complexities associated with uncontrolled and often unmeasured loading paths, which produce different structures at the beginning of the constant stress or constant strain rate portions of the test. For constant stress tests in pure metals, although the concept of steady state is appealing, it appears not yet to have been rigorously demonstrated.

In constant load tests, steady-state behavior would of course result in an increasing creep rate after the minimum, as the true stress increases. As such, the test is inappropriate to evaluate the concept. However, it is by far the most common type of creep test and can be analyzed for instability.

Creep Processes
Creep behavior can be characterized either in terms of deformation mechanisms or in terms of strain constituents.

Deformation Mechanisms. Creep of metals is primarily a result of the motion of dislocations, but is distinct from time-independent behavior in that flow continues as obstacles, which may be dislocation tangles or precipitate particles, are progressively overcome. The rate-controlling step involves diffusion to allow climb of edge dislocations or cross slip of screw dislocations around obstacles. In steady-state theory, there is a balance between the hardening associated with this dislocation motion and interaction, and a dynamic recovery associated with the development of a dislocation substructure.

Theory for such a process predicts a power-law dependence of creep rate on applied stress. At very high homologous temperatures (T/Tm) and low stresses, creep may occur in both metals and ceramics by mass transport involving stress-directed flow of atoms from regions under compression to regions under tension. In this case, theory i ndicates that there is a stress dependence of unity and that the process is controlled either by bulk diffusion or by grain-boundary diffusion. These various processes of creep (dislocation controlled as well as diffusion controlled) may be represented on a deformation mechanism map to highlight regimes of stress and temperature where each mechanism, based on current theories, may be operating. However, such maps are only as good as the theories on which they are based and give no guidance on deformation path dependence.

Another important deformation process in metallic and ionic polycrystals at high temperature and low stresses is grain-boundary sliding. The resistance to sliding is determined by the mobility of grain-boundary dislocations and by the presence of hard particles at the boundary. This sliding leads to stress concentrations at grain junctions, which are important in nucleating cracks. In ductile materials, these stress concentrations may be relieved by creep and stress relaxation in the matrix or by grain-boundary migration.

Strain Components. There are several different sources of strain at high temperature in response to an applied stress. The elastic strain is directly proportional to stress, and a modulus that is temperature dependent can be determined. For metallic materials and ceramics, although there is a strain-rate dependence of elastic modulus, it is small and often ignored. Plastic strain for all materials may be treated as three separate constituents:

* Time-independent nonrecoverable, which may be thought of as an instantaneous deformation
* Time-dependent nonrecoverable, which may involve any or all of the micromechanisms described above
* Time-dependent recoverable.

The first of these is unlikely to be significant in practical applications except in the region of stress concentrations since loading is normally well below the macroscopic yield stress. The second is the major source of creep in normal laboratory testing. The third constituent is not widely studied or analyzed, but may become very important at low stresses and under nonsteady conditions, that is, high-temperature service. It leads to what has been termed creep recovery and anelasticity.

At high temperatures, the application of a stress leads to creep deformation resulting from the motion of dislocations, mass transport by diffusion, or grain-boundary sliding. These processes in turn lead to a distribution of internal stresses that may relax on removal of the stress. In metals it is associated with the unbowing of pinned dislocations, rearrangement of dislocation networks, and local grain-boundary motion.

Whereas the importance of creep recovery is well recognized in polymer design, it has often been ignored in design of metallic and ceramic materials. A few extensive studies have been reported on metals that have led to several broad conclusions:

* Creep-recovery strain increases linearly with stress for a fixed time at a given temperature, but is dependent on prestrain.
* The rate of creep recovery increases with increasing temperature.
* When the stress is low enough, essentially all transient creep is linear with stress and recoverable.
* Mathematically, the recovery may be described by a spectrum of spring dashpot combinations with a wide range of relaxation times.

Stress and Temperature Dependence
The minimum creep rate in both constant load and constant stress tests is normally represented by a power function of stress, and the temperature by an Arrhenius e xpression including an activation energy term (Q) derived from chemical reaction rate theory:

where S, which is a constant, depends on structure. Although an exponential or hyperbolic sine stress function may provide a better fit in some cases, the power function has generally prevailed and has become strongly linked with mechanistic treatments.

Shapes and bars are normally available as Grades A, AH32, or AH36. Other grades may be furnished by agreement between the purchaser and the manufacturer.

When the steel is to be welded, it is presupposed that a welding procedure suitable for the grade of steel and intended use or service will be utilized.
Applicable Documents

* ASTM A6M Specification for General Requirements for Delivery of Rolled Steel Plates, Shapes, Sheet Piling, and Bars for Structural Use
* ASTM A 370 Methods and Definitions for Mechanical Testing of Steel Products
* ASTM E 112 Method of Determining the Average Grain Size

Manufacture
The steel may be made by any of the following processes: open-hearth, basic-oxygen, electric-furnace, vacuum arc remelt (VAR), or electroslag remelt (ESR).

Except for Grade A steel up to and including 12.5 mm in thickness, rimming-type steels shall not be applied.

Grades AH32 and AH36 shapes through 426 lb/ft, and plates up to 12.5 mm in thickness may be semi-killed, in which case the 0.10 % minimum silicon does not apply.

Besides few exceptions Grades D, DS, CS, E, DH32, DH36, EH32, and EH36 shall be made using a fine grain practice. For ordinary strength grades, aluminum shall be used to obtain grain refinement. For high strength grades, aluminum, vanadium, or columbium (niobium) may be used for grain refinement.

Grade D material 35 mm and under in thickness, at the option of the manufacturer, may be semi-killed and exempt from the fine austenitic grain size.
Heat Treatment
Plates in all thicknesses ordered to Grades CS and E shall be normalized. Plates over 35 mm in thickness ordered to Grade D shall be normalized. When Grade D steel is furnished semi-killed, it shall be normalized over 25 mm in thickness. Upon agreement between the purchaser and the manufacturer, control rolling of Grade D steel may be substituted for normalizing, in which case impact tests are required for each 25 tons [25 Mg] of material in the heat.

Plates in all thicknesses ordered to Grades EH32 and EH36 shall be normalized. Grades AH32, AH36, DH32, and DH36 shall be normalized when so specified. Upon agreement between the purchaser and the manufacturer, control rolling of Grade DH may be substituted for normalizing, in which case impact tests are required on each plate.

In the case of shapes, the thicknesses referred to are those of the flange.
Metallurgical Structure
Fine grain practice for ordinary strength grades shall be met using aluminum. For higher strength grades, aluminum, vanadium, or columbium may be used as grain refining elements.

Grain size shall be determined on each heat by the Mc-Quaid-Ehn Method of Method E 112. The grain size so determined shall be No. 5 or finer in 70 % of the area examined.

As an alternative to the McQuaid-Ehn test, a fine grain practice requirement may be met by a minimum acid-soluble aluminum content of 0.015 % or minimum total aluminum content of 0.020 % for each heat.

For Grades DH32, DH36, EH32, and EH36 the fine grain practice requirement may also be met by the following:

* Minimum columbium (niobium) content of 0.020 % or minimum vanadium content of 0.050 % for each heat, or
* When vanadium and aluminum are used in combination, minimum vanadium content of 0.030 % and minimum acid-soluble aluminum content of 0.010 % or minimum total aluminum content of 0.015%.

Mechanical Requirements
Tension Tests. Except as specified in the following paragraphs the material as represented by the test specimens shall conform to the prescribed tensile requirements.

Unless a specific orientation is called for on the purchase order, tension test specimens may be taken parallel or transverse to the final direction of rolling at the option of the steel manufacturer.

Shapes less than 645 mm2 in cross section, and bars, other than flats, less than 12.5mm in thickness or diameter need not be subjected to tension tests by the manufacturer. For material under 8 mm in thickness or diameter, a deduction from the percentage of elongation in 200 mm of 1.25 percentage points shall be made for each decrease of 0.8 mm of the specified thickness or diameter below 8 mm.

Toughness Tests (material 50 mm and less in thickness). Except as permitted bellow, Charpy V-notch tests shall be made on Grade B material over 25 mm in thickness and on material of Grades D, E, AH32, AH36, DH32, DH36, EH32, and EH36.

Toughness tests are not required: (a) on Grade D normalized material made fully killed and having a fine austenitic grain size, (b) on Grades AH32 and AH36 when normalized, or when 12.5 mm or less in thickness when treated with vanadium or columbium (niobium) or 35 mm or less in thickness when treated with aluminum, and (c) on Grades DH32 and DH36 material when normalized or when 12.5 mm or less in thickness when treated with vanadium or columbium (niobium) or less in thickness when treated with aluminum, and on Grades DH32 and DH36 material when normalized.

For plate material, when required, one set of three impact specimens shall be made from the thickest material in each 50 tons [45 Mg] of each heat of Grades B, D, AH32, AH36, DH32, and DH36 steels and from each rolled product of normalized Grades E, EH32, and EH36 steels. When heat testing is called for, a set of three specimens shall be tested for each 50 tons [45 Mg] of the same type of product produced on the same mill from each heat of steel. The set of impact specimens shall be taken from different as-rolled or heat-treated pieces of the heaviest gage produced. An as-rolled piece refers to the product rolled from a slab, billet, bloom, or directly from an ingot.

For flats, rounds, and shapes, one set of three impact tests shall be taken from each 25 tons [25 Mg] of each heat for Grade E, EH32, or EH36 and, when required, from each 50 tons [45 Mg] of each heat of Grade B, D, AH32, AH36, DH32, or DH36 material. Where the maximum thickness or diameter of various sections differs by 10 mm or more, one set of impacts shall be made from both the thickest and the thinnest material rolled regardless of the weight represented.

The specimens for plates shall be taken from a corner of the material and the specimens from shapes shall be taken from the end of a shape at a point one-third the distance from the outer edge of the flange or leg to the web or heel of the shape.

Specimens for bars shall be in accordance with Specification A 6M. The center longitudinal axis of the specimens shall be located as near as practical midway between the surface and the center of the material and the length of the notch shall be perpendicular to the rolled surface. Unless a specific orientation is called for on the purchase order, the longitudinal axis of the specimens may be parallel or transverse to the final direction of rolling of the material at the option of the steel manufacturer.

Each impact test shall constitute the average value of three specimens taken from a single test location.

After heat treatment or reheat treatment a set of three specimens shall be tested and evaluated in the same manner as for the original material.

Toughness Tests (material over 50 mm thick). Charpy V-notch tests are required for all grades of steel over 50 mm thick, except for Grade A that is produced killed, using a fine grain practice and normalized. For plate material one set of three impact specimens shall be made from the thickest material in each 50 tons [45 Mg] of each heat of Grades A, B, D, DS, AH32, AH36, DH32, and DH36, and from each rolled product of Grades CS, E, EH32, and EH36. For flats, rounds, and shapes, one set of three impact tests shall be taken from each 25 tons [25 Mg] of each heat for Grades CS, E, EH32, and EH36, and from each 50 tons of each heat of Grades A, B, D, DS, AH32, AH36, DH32, and DH36 material.

Rivet Steel and Rivets. For rivet steel a sulfur print requirement shall be met when other than killed or semi-killed steel is applied, in order to confirm that its core is free of concentrations of sulfur segregates and other nonmetallic substances. Test specimens for rivet bars that have been cold drawn shall be normalized before testing. Finished rivets are to be selected as sample specimens from each diameter and tested hot and cold by bending and crushing in the following manner: the shank must stand being doubled together cold, and the head being flattened hot to a diameter 2.5 times the diameter of the shank, both without fracture.

This group of ASTM standard specifications covers a common requirements that, unless otherwise specified in the material specification, apply to rolled steel plates, shapes, sheet piling, and bars under each of the following specifications issued by ASTM.

ASTM Designation	Title of Specification
A 36/A 36M	Structural Steel
A 131/A 131 M	Structural Steel for Ships
A 242/A 242 M	High-Strength Low-Alloy Structural Steel
A 283/A 283 M	Low and Intermediate Tensile Strength Carbon Steel Plates, Shapes, and Bars
A 284/A 284 M	Low and Intermediate Tensile Strength Carbon-Silicon Steel Plates for Machine Parts and General Construction
A 328/A 328 M	Steel Sheet Piling
A 441/A 443 M	High-Strength Low-Alloy Structural Manganese Vanadium Steel
A 514/A 514 M	High-Yield Strength, Quenched and Tempered Alloy Steel Plate Suitable for Welding
A 529/A 529 M	Structural Steel with 42 000 psi {290 MPa) Minimum Yield Point (12.7 mm Maximum Thickness)
A 572/A 572 M	High-Strength Low-Alloy Columbium-Vanadium Steels of Structural Quality
A 573/A 573 M	Structural Carbon Steel Plates of Improved Toughness
A 588/A 588 M	High-Strength Low-Alloy Structural Steel with 50 ksi (345 MPa) Minimum Yield Point to 4 in. Thick
A 633/A 633 M	Normalized High-Strength Low-Alloy Structural Steel
A 656/A 656 M	Hot-Rolled Structural Steel, High-Strength Low-Alloy Plate with Improved Formability
A 678/A 678 M	Quenched and Tempered Carbon Steel Plates for Structural Applications
A 690/A 690 M	High-Strength Low-Alloy Steel H-Piles and Sheet Piling for Use in Marine Environments
A 699	Low-Carbon Manganese-Molybdenum-Columbium Alloy Steel Plates, Shapes, and Bars
A 709	Structural Steel for Bridges
A710/A710 M	Low-Carbon Age-Hardening Nickel-Copper-Chromium-Molybdenum-Columbium and Nickel-Copper-Columbium Alloy Steels
A 769	Electric Resistance Welded Steel Shapes
A 786/A 786 M	Rolled Steel Floor Plates
A 808/A 808 M	High-Strength Low-Alloy Carbon, Manganese, Columbium, Vanadium Steel of Structural Quality with Improved Notch Toughness
A 827	Plates, Carbon Steel, for Forging and Similar Applications
A 829	Plates, Alloy Steel, Structural Quality
A 830	Plates, Carbon Steel, Structural Quality, Furnished to Chemical Composition Requirements

Descriptions of Terms Specific to This Standard
Plates (other than floor plates or coiled product) - Flat hot-rolled steel classified as follows:

When ordered to thickness:

* Over 200 mm in width and over 6 mm or over in thickness.
* Over 1200 mm in width and over 4.5 mm or over in thickness.

When ordered to weight:

* Over 200 mm in width and 47.1 kg/m2 or heavier.
* Over 1200 mm in width and 35.3 kg/m2 or heavier.

Slabs, sheet bars, and skelp, though frequently falling in the foregoing size ranges, are not classed as plates. Coiled product is excluded from qualification to this specification until cut to length.

Structural-Size Shapes - Rolled flanged sections having at least one dimension of the cross section 75 mm or greater.
Bar Size Shapes - Rolled flanged sections having a maximum dimension of the cross section less than 75 mm.

"W"Shapes are doubly-symmetric wide-flange shapes used as beams or columns whose inside flange surfaces are substantially parallel. A shape having essentially the same nominal weight and dimensions as a "W" shape but whose inside flange surfaces are not parallel may also be considered a "W" shape having the same nomenclature, provided its average flange thickness is essentially the same as the flange thickness of the "W" shape.

"HP"Shapes are wide-flange shapes generally used as bearing piles whose flanges and. webs are of the same nominal thickness and whose depth and width are essentially the same.

"S"Shapes are doubly-symmetric shapes produced in accordance with dimensional standards adopted in 1896 by the Association of American Steel Manufacturers for American Standard beam shapes.

"M" Shapes are doubly-symmetric shapes that cannot be classified as "W," "S," or "HP" shapes.

"C" Shapes are channels produced in accordance with dimensional standards adopted in 1896 by the Association of American Steel Manufacturers for American Standard channels.

"MC" Shapes are channels that cannot be classified as "C" shapes.

"L" Shapes are equal-leg and unequal-leg angles.

Sheet Piling - Steel sheet piling consists of rolled sections that can be interlocked, forming a continuous wall when individual pieces are driven side by side.

Bars - Rounds, squares, and hexagons, of all sizes; flats over 5 mm and over in specified thickness, not over 150 mm in specified width; and flats over 6 mm in specified thickness, over 150 to 200 mm incl, in specified width.

Exclusive - When used in relation to ranges, as for ranges of thickness in the tables of permissible variations in dimensions, the term is intended to exclude only the greater value of the range.

Rimmed Steel - Steel containing sufficient oxygen to give a continuous evolution of carbon monoxide while the ingot is solidifying, resulting in a case or rim of metal virtually free of voids.

Semi-killed Steel - Incompletely deoxidized steel containing sufficient oxygen to form enough carbon monoxide during solidification to offset solidification shrinkage.

Capped Steel - Rimmed steel in which the rimming action is limited by an early capping operation. Capping may be carried out mechanically by using a heavy metal cap on a bottle-top mold or it may be carried out chemically by an addition of aluminum or ferrosilicon to the top of the molten steel in an open-top mold.

Killed Steel - Steel deoxidized, either by addition of strong deoxidizing agents or by vacuum treatment, to reduce the oxygen content to such a level that no reaction occurs between carbon and oxygen during solidification.

Groupings for Tensile Properly Classification - In some of the material specifications, the tensile property requirements vary for different sizes of shapes due to mass effect, etc. For the convenience of those using the specifications, the various sizes of shapes have been divided into groups based on section thickness at the standard tension test location (webs of beams, channels, and zees; legs of angles; and stems of tees).

Mill Edge - The normal edge produced by rolling between horizontal finishing rolls. A mill edge does not conform to any definite contour. Mill edge plates have two mill edges and two trimmed edges.

Universal Mill Edge - The normal edge produced by rolling between horizontal and vertical finishing rolls. Universal mill plates, sometimes designated UM Plates, have two universal mill edges and two trimmed edges.

Sheared Edge - The normal edge produced by shearing. Sheared edge plates are trimmed on all edges.

Gas Cut Edge - The edge produced by gas flame cutting.

Special Cut Edge - Usually the edge produced by gas flame cutting involving special practices such as pre-heating or post-heating, or both, in order to minimize stresses, avoid thermal cracking and reduce the hardness of the gas cut edge. In special instances, special cut edge may be used to designate an edge produced by machining.

Sketch - When used to describe a form of plate, denotes a plate other than rectangular, circular, or semi-circular. Sketch plates may be furnished to a radius or with four or more straight sides.

Manufacture
Unless otherwise specified in the material specification, the steel shall be made by the open-hearth, basic-oxygen, or electric-furnace process. Additional refining by vacuum-arc-remelt (VAR) or electroslag-remelt (ESR) is permitted.

Plates are produced in either discrete cut lengths of flat product or from coils.

Plates produced from coil means plates that have been cut to individual lengths from a coiled product and are furnished without heat treatment. For the purposes of this paragraph, stress relieving is not considered to be a heat treatment.
Heat Treatment
When material is required to be heat treated, the heat treatment may be performed either by the manufacturer, processor, or fabricator unless otherwise specified in the material specification.

When heat treatment is to be performed by the manufacturer or processor, the material shall be heat treated as specified in the material specification. The purchaser may specify the heat treatment to be used provided it is not in conflict with the requirements of the material specification.

When normalizing is to be performed by the fabricator, it may be accomplished by heating uniformly for hot forming. The temperature to which the plates are heated for hot forming shall not significantly exceed the normalizing temperature.

When no heat treatment is required, the manufacturer or processor may, at his option, heat treat the plates by normalizing, stress relieving, or normalizing and then stress relieving to meet the material specification.

If approved by the purchaser, cooling rates faster than those obtained by cooling in air are permissible for improvement of the toughness, provided the plates are subsequently tempered in the temperature range from 595 to 705°C.
Chemical Analysis
An analysis of each heat shall be made by the manufacturer to determine the percentage of carbon, manganese, phosphorus, sulfur, and any other elements specified or restricted by the applicable specification. This analysis shall be made from a test sample preferably taken during the pouring of the heat.

When vacuum-arc-remelting or electroslag remelting is used, a heat is defined as all the ingots remelted from a single primary melt. The heat analysis shall be obtained from one remelted ingot, or the product of one remelted ingot, of each primary melt providing the heat analysis of the primary melt meets the heat analysis requirements of the material specification. If the heat analysis of the primary melt does not meet the heat analysis requirements of the material specification, one test sample shall be taken from the product of each remelted ingot. In either case, the analyses so obtained from the remelted material shall conform to the heat analysis requirements of the applicable specification.
Metallurgical Structure
When a fine austenitic grain size is specified, the steel shall have a grain size number of 5 or finer as determined by the McQuaid-Ehn test. Determination shall be in accordance with Plate IV of Methods E 112, by carburizing for 8 h at 925°C. Conformance to this grain size of 70 % of the grains in the area examined shall constitute the basis of acceptance. One test per heat shall be made.

Abrasion-Resisting Steel
Abrasion-resisting steel (AR) is carbon steel usually with a high-carbon analysis used as liners in material-moving systems and for construction equipment where severe abrasion and sharp hard materials are encountered. Abrasion-resisting steels are often used to line dump truck bodies for quarry service, for lining conveyors, chutes, bins, etc.

Normally the abrasion-resisting steel is not used for structural strength purposes, but only to provide lining materials for wear resistance. Various steel companies make different proprietary alloys that all have similar properties and, in general, similar compositions. Most AR steels are high-carbon steel in the 0.80-0.90% carbon range; however some are low carbon with multiple alloying elements.

These steels are strong and have hardness up to 40 HRc or 375 BHN. Abrasion-resisting bars or plates are welded to the structures and as they wear they are removed by oxygen cutting or air carbon arc and new plates installed by welding.

Low-hydrogen welding processes are required for welding abrasion-resisting steels. Local preheat of 400°F (204°C) is advisable to avoid underbead cracking of the base metal or cracking of the weld. In some cases this can be avoided by using a preheat weld bead on the carbon steel structure and filling in between the bead and the abrasion-resisting steel with a second bead in the groove provided. The first bead tends to locally preheat the abrasion-resisting steel to avoid cracking and the second bead is made having a full throat. Intermittent welds are usually made since continuous or full-length welds are usually not required. Efforts should be made to avoid deep weld penetration into the abrasion-resisting steel so as not to pick up too much carbon in the weld metal deposit. If too much carbon is picked up the weld bead will have a tendency to crack.

When using the shielded metal arc welding process the E-XX15, E-XX16, or E-XXX8 type electrodes are used. When using gas metal arc welding the low penetrating type shielding gases such as the 15% argon-25% CO2 mixture should be used. The flux-cored arc welding process is used and the self-shielding version is preferred since it does not have the deep penetrating quality as the CO2 shielded version.

During cold weather applications, it is recommended that the abrasion-resistant steel be brought up to 100°F (38°C) temperature prior to welding.

Free Machining Steels
The term free machining can apply to many metals but it is normally associated with steel and brass. Free machining is the property that makes machining easy because small cutting chips are formed. This characteristic is given to steel by sulfur and in some cases by lead. It is given to brass by lead.

Sulfur and lead are not considered alloying elements. In general, they are considered impurities in the steel. The specifications for steel show a maximum amount of sulfur as 0.040% with the actual sulfur content running lower, in the neighborhood of 0.030%. Lead is usually not mentioned in steel specifications since it is not expected and is considered a "tramp" element. Lead is sometimes purposely added to steel to give it free-machining properties.

Free-machining steels are usually specified for parts that require a considerable amount of machine tool work. The addition of the sulfur makes the steel easier to turn, drill, mill, etc., even though the hardness is the same as a steel of the same composition without the sulfur.

The sulfur content of free-machining steels will range from 0.07-0.12% as high as 0.24-0.33%. The amount of sulfur is specified in the AISI or other specifications for carbon steels. Sulfur is not added to any of the alloy steels. Leaded grades comparable to 12L14 and 11L18 are available.

Unless the correct welding procedure is used, the weld deposits on free-machining steel will always be porous and will not provide properties normally expected of a steel of the analysis but without the sulfur or lead.

The basis for establishing a welding procedure for free-machining steels is the same as that required for carbon steels of the same analysis. These steels usually run from 0.010% carbon to as high as 1.0% carbon. They may also contain manganese ranging from 0.30% to as high as 1.65%. Therefore, the procedure is based on these elements. If the steels are free-machining and contain a high percentage of sulfur the only change in procedure is to change to a low-hydrogen type weld deposit.

In the case of shielded metal arc welding this means the use of low-hydrogen type electrode of the E-XX15, E-XX16, or the E-XXX8 classification. In the case of gas metal arc welding or flux-cored arc welding the same type of filler metal is specified as is normally used since these are no-hydrogen welding processes.

Submerged arc welding would not normally be used on free-machining steels. Gas tungsten arc welding is not normally used since free-machining steels are used in thicker sections which are not usually welded with the GTAW process.

Manganese Steel
Manganese steel is sometimes called austenitic manganese steel because of its metallurgical structure. It is also called Hadfield manganese steel after its inventor. It is an extremely tough, nonmagnetic alloy. It has an extremely high tensile strength, a high percentage of ductility, and excellent wear resistance. It also has a high resistance to impact and is practically impossible to machine.

Hadfield manganese steel is probably more widely used as castings but is also available as rolled shapes. Manganese steel is popular for impact wear resistance. It is used for railroad frogs, for steel mill coupling housings, pinions, spindles, and for dipper lips of power shovels operating in quarries. It is also used for power shovel track pads, drive tumblers, and dipper racks and pinions.

The composition of austenitic manganese is from 12-14% manganese and 1-1.4% carbon. The composition of cast manganese steel would be 12% manganese and 1.2% carbon. Nickel is oftentimes added to the composition of the rolled manganese steel.

A special heat treatment is required to provide the superior properties of manganese steel. This involves heating to 1850°F (1008°C) followed by quenching in water. In view of this type of heat treatment and the material toughness, special attention must be given to welding and to any reheating of manganese steel.

Manganese steel can be welded to itself and defects can be weld repaired in manganese castings. Manganese steel can also be welded to carbon and alloy steels and weld surfacing deposits can be made on manganese steels.

Manganese steel can be prepared for welding by flame cutting; however, every effort should be made to keep the base metal as cool as possible. If the mass of the part to be cut is sufficiently large it is doubtful if much heat will build up in the part sufficient to cause embrittlement. However, if the part is small it is recommended that it be frequently cooled in water or, if possible, partially submerged in water during the flame cutting operation. For removal of cracks the air carbon arc process can be used. The base metal must be kept cool. Cracks should be completely removed to sound metal prior to rewelding. Grinding can be employed to smooth up these surfaces.

There are two types of manganese steel electrodes available. Both are similar in analysis to the base metal but with the addition of elements which maintain the toughness of the weld deposit without quenching. The EFeMn-A electrode is known as the nickel manganese electrode and contains from 3-5% nickel in addition to the 12-14% manganese. The carbon is lower than normal manganese ranging from 0.50 to 0.90%. The weld deposits of this electrode on large manganese castings will result in a tough deposit due to the rapid cooling of the weld metal.

The other electrode used is a molybdenum-manganese steel type EFeMn-B. This electrode contains 0.6-1.4% molybdenum instead of the nickel. This electrode is less often used for repair welding of manganese steel or for joining manganese steel itself or to carbon steel. The manganese nickel steel is more often used as a buildup deposit to maintain the characteristics of manganese steel when surfacing is required.

Stainless steel electrodes can also be used for welding manganese steels and for welding them to carbon and low-alloy steels. The 18-8 chrome-nickel types are popular; however, in some cases when welding to alloy steels the 29-9 nickel is sometimes used. These electrodes are considerably more expensive than the manganese steel electrode and for this reason are not popular.

Silicon Steels
Silicon steels or, as they are sometimes called, electrical steels, are steels that contain from 0.5% to almost 5% silicon but with low carbon and low sulfur and phosphorous. Silicon steel is primarily provided as sheet or strip so that it can be punched or stamped to make laminations for electrical machinery.

The silicon steels are designed to have lower hysteresis and eddy current losses than plain steel when used in magnetic circuits. This is a particular advantage when used in alternating magnetic fields. Their magnetic properties make silicon steels useful in direct current fields for most applications.

Silicon steel stampings are used in the laminations of electric motor armatures, rotors, and generators. They are widely used in transformers for the electrical power industry and for transformers, chokes, and other components in the electronics industry.

Welding is important to silicon steels since many of which are welded together. Welds are made on the edge of each sheet to hold the stack together. Welding is done instead of punching holes and riveting the laminations in order to reduce manufacturing costs. Almost all of the arc welding processes are used, submerged arc, shielded metal arc, gas metal arc, gas tungsten arc and plasma arc welding. The more popular processes are gas metal arc using CO2 for gas shielding and the gas tungsten arc process. The plasma arc process is used for some of the smaller assemblies.

When the consumable electrode processes are used the stampings are usually indented to allow for deposition of filler metal. For gas tungsten arc and plasma arc the filler metals are not used and the edges are fused. The size of the weld bead should be kept at minimum so that eddy currents are not conducted between laminations in the electrical stack.

One precaution that should be taken in welding silicon steel laminations is to make sure that the laminations are tightly pressed together and that all of the oil used for protection and used in manufacturing is at a minimum. Oil can cause porosity in the welds which might be detrimental to the lamination assembly.

Most clad metals are composites of a cladding metal such as stainless steel, nickel and nickel alloys, and copper and copper alloys welded to a backing material of either carbon or alloy steel. The two metals are welded together at a mill in a roll under heat and pressure. The clad composite plates are usually specified in a thickness of the cladding which ranges from 5% to 20% of the total composite thickness. The advantage of composite material is to provide at relatively low cost the benefits of an expensive material which can provide corrosion resistance, abrasion resistance, and other benefits with the strength of the backing metal.

Clad metals were developed in the early 1930s and one of the first to be used was nickel bonded to carbon steel. This composite was used in the construction of tank cars. Other products made of clad steels are heat exchangers, tanks, processing vessels, materials-handling equipment, storage equipment, etc.

Clad or composites can be made by several different welding manufacturing methods. The most widely used process is roll welding which employs heat and roll pressure to weld the clad to the backing steel. Explosive welding is also used and weld surfacing or overlay is another method of producing a composite material.

Clad steels can have as the cladding material chromium steel in the 12-15% range, stainless steels primarily of the 18/8 and 25/12 analysis, nickel base alloys such as Monel and Inconel, copper-nickel, and copper. The backing material is usually high-quality steel of the ASTM-A285, A212, or similar grade. The tensile strength of clad material depends on the tensile strength of its components and their ratio to its thickness. The clad thickness is uniform throughout the cross section, and the weld between the two metals is continuous throughout.

A slightly different procedure is used for oxygen cutting of clad steel. All of the clad metals mentioned above can be oxygen flame cut with the exception of the copper-clad composite material. The normal limit of clad plate cutting is when the clad material does not exceed 30% of the total thickness. However, higher percentage of cladding may be cut in thicknesses of 12 mm and over. The oxygen pressure is lower when cutting clad steel; however, larger cutting tips are used.

The quality of the cut is very similar to the quality of the cut of carbon steel. When flame cutting clad material the cladding material must be on the underside so that the flame will first cut the carbon steel. The addition of iron powder to the flame will assist the cutting operation.

Schedules of flame cutting are provided by clad steel producers as well as flame-cutting equipment producers. For oxygen flame cutting copper and copper-nickel clad steels the copper clad surface must be removed and the backing steel cut in the same fashion as bare carbon steel.

Copper and brass clad plate can be cut using iron powder cutting. Clad steels can be fabricated by bending and rolling, shearing, punching, and machining in the same manner as the equivalent carbon steels. Clad materials can be preheated and given stress relief heat treatment in the same manner as carbon steels. However, stress relieving temperatures should be verified by consulting with the manufacturer of the clad material.

Clad materials can be successfully welded by adopting special joint details and following specific welding procedures. Special joint details and welding procedures are established in order to maintain uniform characteristic of the clad material. Inasmuch as the clad material is utilized to provide special properties it is important that the weld joint retain these same properties. It is also important that the structural strength of the joint be obtained with the quality welds of the backing metal.

The normal procedure for making a butt joint in clad plate is to weld the backing or steel side first with a welding procedure suitable for the carbon steel base material being welded. Then the clad side is welded with the suitable procedure for the material being joined. This sequence is preferable in order to avoid the possibility of producing hard brittle deposits, which might occur if carbon steel weld metal is deposited on the clad material.

Different joint preparations can be used to avoid the possible pickup of carbon steel in the clad alloy weld. Any weld joint made on clad material should be a full-penetration joint. When designing the joint details it is wise to make the root of the weld the clad side of the composite plate. This may not always be possible; however, it is more economical since most of the weld metal can be of the less expensive carbon steel rather than the expensive alloy clad metal.

The selection of the welding process or processes to be used would be based on that normally used for welding the material in question in the thickness and position required. Shielded metal arc welding is probably used more often; however, submerged arc welding is used for fabricating large thick vessels and the gas metal arc welding process is used for medium thicknesses; the flux-cored arc welding process is used for the steel side, and gas tungsten arc welding is sometimes used for the thinner materials, particularly the clad side.

The selection of process should be based on all factors normally considered. It is important to select a process that will avoid penetrating from one material into the other. The welding procedure should be designed so that the clad side is joined using the appropriate process and filler metal to be used with the clad metal and the backing side should be welded with the appropriate process and filler metal recommended for the backing metal. For code work the welding procedure must be qualified in accordance with the specification requirements.

The normal procedure, assuming that the material is properly prepared and fitted is as follows. The backing side or steel side would be welded first. The depth of penetration of the root pass must be closely controlled by selecting the proper procedure and filler metal. It is desirable to produce a root pass which will penetrate through the root of the backing metal weld joint into the root face area yet not come in contact with the clad metal.

A low-hydrogen deposit is recommended. If penetration is excessive and the root bead melts into the clad material because of poor fitup or any other reason the deposit will be brittle. If this occurs the weld will have to be removed and remade. However, if the penetration of the backing steel root bead is insufficient the amount of back gouging will be excessive and larger amounts of the clad material weld metal will be required. The steel side of the joint should be welded at least half way prior to making any of the weld on the clad side. If warpage is not a factor, the steel side weld can be completed before welding is started on the clad side.

The clad side of the joint is prepared by gouging to sound metal or into the root pass made from the backing steel side. This can be done by air carbon arc gouging or by chipping. The gouging should be sufficient to penetrate into the root pass so that a full penetration of the joint will result. This will determine the depth of the gouging operation. It is also a measure of the depth of penetration of the root pass. Grinding is not recommended since it tends to wander from the root of the joint and may also cover up an unfused root by smearing the metal. If the depth of gouging is excessive, weld passes made with the steel electrode may be required to avoid using an excessive amount of clad metal electrode.

On thin materials the gas tungsten arc welding process may be used, on thicker materials the shielded metal arc welding process or the gas metal arc process may be used. The filler metal must be selected to be compatible with the clad metal analysis.

There is always the likelihood of diluting the clad metal deposit by too much penetration into the steel backing metal. Special technique should be used to minimize penetration into the steel backing material. This is done by directing the arc on the molten puddle instead of on the base metal.

When welding copper or copper nickel clad steels a high nickel electrode is recommended for the first pass (ECuNi or ENi-1). The remaining passes of the joint in the clad metal should be welded so that the copper or copper nickel electrode matches the composition of the clad metal.

When the clad metal is stainless steel the initial pass which might fuse into the carbon steel backing should be of a richer analysis of alloying elements than necessary to match the stainless cladding. This same principle is used when the clad material is Inconel or Monel. The remaining portion of the clad side weld should be made with the electrode compatible with or having the same analysis as the clad metal. The procedure should be designed so that the final weld layer will have the same composition as the clad metal.

On heavier thicknesses where the weld of the backing steel is made from both sides it is important to avoid allowing the steel weld metal to come in contact or to fuse with the clad metal. This will cause a contamination of the deposit which may result in a brittle weld.

When welding thinner gauge clad plate and inside clad pipe it may be more economical to make the complete weld using the alloy weld metal compatible with the clad metal instead of using two types of filler metal. The alloy filler metal must be compatible with the steel backing metal. The expense of the welding filler metal may be higher, but the total weld joint may be less expensive because of the more straightforward procedure. Joint preparation may also be less extensive using this procedure.

For medium thickness, the joint preparation is a single vee or bevel without a large root face. The root face is obtained by grinding the feather edge to provide a small root face. If possible the face of the weld will be the steel or backing side of the joint. The backing side or steel side is welded first using the small diameter electrode for the root pass to insure complete penetration.

If the composite is a pipe or if it must be welded from one side, the buttering technique should be used. In this case the filler metal must provide an analysis equal to the clad metal and be compatible with the backing steel. Weld passes are made on the edge of the composite to butter the clad and backing metal. The buttering pass must be smoothed to the design dimensions prior to fitup. The same electrode can be used to make the joint.

When welding heavy, thick composite plate the U groove weld joint design is recommended instead of the vee groove in order to minimize the amount of weld metal. The same principles mentioned previously are used.

When the submerged arc welding process is used for the steel side of the clad plate caution must be exercised to avoid penetrating into the clad metal. This same caution applies to automatic flux-cored arc welding or gas metal arc welding. A larger root face is required and fitup must be very accurate in order to control root bead penetration.

The submerged arc process can also be used on the clad side when welding stainless alloys. However, caution must be exercised to minimize dilution of a high-alloy material with the carbon steel backing metal. The proper filler metal and flux must be utilized. To minimize admixture of the final pass it is recommended that the clad side be welded with at least two passes so that dilution would be minimized in the final pass.

Special quality control precautions must be established when welding clad metals so that undercut, incomplete penetration, lack of fusion, etc., are not allowed. In addition, special inspection techniques must be incorporated to detect cracks or other defects in the weld joints. This is particularly important with respect to the clad side which may be exposed to a corrosive environment.

Selection of a welding process is determined primarily by the characteristics of the joint, the materials involved, their shape and thickness, and joint design. Additionally, production requirements, such as rate and quality, must also be considered. Only after the process has been determined can the proper power supply and accessory equipment be chosen. The process is the primary factor in their selection.

It is our purpose in this article to provide a guide to the selection of power supplies for some of the welding processes, which have come into maturity during the past decade. Once thought of primarily as special processes or processes designed primarily for mass production operations, they are now found in virtually every area of metal fabrication. The increase in the fabrication of once difficult-to-weld metals and the economic and qualitative advantages of the more advanced welding processes are today recognized by even the smallest metal fabricating shops.

Power and control requirements for these processes are somewhat more sophisticated than for conventional shielded metal arc or stick electrode welding. But since the average metallurgist or welding engineer isn’t an electrical expert, the selection of proper power supply and the reasons for it can be confusing. The purpose of this article is to eliminate some of this mystery through the examination of the power requirements of production welding processes.

The basic types of power supplies are:

* constant current and
* constant voltage.

Constant Current Power Supply
The conventional stick electrode welder is sometimes called a ’constant current machine’. It is also called a ’dropper’ because its voltage drops as welding current increases, thus its volt- ampere output curve ’droops’.

With the machine turned on but with no arc, and hence no current flowing, it has a relatively high open circuit voltage of 70-80 volts. Generally speaking welding is done at the steeper portions of the curve and this is ideal for manual stick electrode welding.

Arc voltage depends upon the physical length of the arc between the electrode and the work and this can never be held completely constant in manual welding. But, since rate of burn off of filler metal is determined by the amount of current, burn off stays substantially constant if current doesn’t vary.

There are many variations of this type of machine based upon power input (single or three phase), output (ac, dc, or ac/dc), and the type of output control (mechanical or electrical).

Constant Voltage Power Supply
The other basic type of arc welding power supply produces a constant voltage. Thus at any voltage setting current may vary from zero to an extremely high short circuit current. Such a machine is designed specifically for gas shielded metal arc welding and is not generally suitable for stick electrode welding.

Actually, no welding machine can produce a truly constant voltage. In practice, voltage drops at least 1 volt for each 100 amp output. Nevertheless, short circuit currents may be as high as several thousand amperes.

Normally, constant voltage machines have lower open circuit voltages than the constant current machines, about 50 volts maximum as compared to 80 volts. As a means of obtaining the desired arc voltage, the operator sets open circuit voltage, rather than current, at the machine. Settings may range from 10 to 48 volts.

Welding current can reach several thousand amperes at short circuit. Current adjusts itself to burn off filler metal at a rate sufficient to maintain the arc length required by the present voltage and is thus determined by the rate of electrode feed.

Duty Cycle
All welding machines are rated according to their duty cycle. Understanding this term is of utmost importance and it is often misunderstood. Duty cycle is based upon a ten minute time period. At rated voltage, a power supply with a 100% duty cycle rating can operate continuously at or below its rated current.

A 60% duty cycle does not mean that it can operate 60% of an indeterminate time at rated current and voltage. It means that the welder should operate only 6 out of every 10 minutes at that current and voltage. It should be allowed to idle 4 out of every 10 minutes for cooling. Machines rated for less than 100% duty cycle can be used continuously by decreasing their current rating.

Tungsten Arc Welding
Tungsten arc welding, using inert shielding gases, is hardly a new process. Nevertheless, it should be covered in any survey of advanced welding processes because of its particular applicability to difficult welding problems. This includes joining hard-to-weld and exotic materials such as the stainless steels, aluminum, magnesium, copper, beryllium copper, Hastelloy, Inconel, Invar and Kovar, especially in very thin cross sections.

Essentially, TIG welding calls for the same type of power source as shielded metal arc, or stick electrode welding, that is one with a drooping volt-ampere output curve. However, the process does present some problems which make a machine especially designed for the process much more suitable.

It should be noted that a conventional AC power source not specifically designed for TIG welding must be derated for AC TIG service. This is because partial rectification occurring at the arc introduces a DC component, which causes overheating of the main transformer. Other problems associated with TIG welding which are more acute than in conventional metallic arc work include arc starting, arc stabilization, and, of course, as the work becomes more delicate, control of all welding variables.

Gas Metal Arc Processes
Gas metal arc welding, in which a consumable wire electrode is fed continuously to a gas shielded arc zone, has replaced non-consumable tungsten inert gas welding and conventional shielded metal arc (stick electrode) welding in many applications. The main reason is its speed. Where it can be used, it is usually several times faster than other processes.

Other advantages include: cleaner welds, because the shielding gas greatly reduces and often prevents oxide formation; electrode savings through the virtual elimination of stub losses; excellent weld metallurgical and physical characteristics: and simplicity of operation which increases weld quality and reproducibility and generally reduces the human variable.

Arc Spot Welding
Arc spot welding has become very popular in recent decades. Its chief advantage is the ability to spot weld from one side of the work. In addition, it is a fast method of producing multiple spot welds with a high degree of reproducibility.

TIG Spot Welding
Arc spot welding may be performed either by tungsten inert gas or gas shielded metal arc processes. For tungsten spot welding the same type of power supply used for regular TIG welding may be employed. However, it requires a special spot welding gun and controls.

TIG spot welding involves the fusion of the parent metal only. Filler wire is not used. Generally TIG spot welding is performed on cold rolled and stainless steels.

MIG Spot Welding
Gas shielded metal arc spot welding is characterized by high amperages. Machines designed for this purpose normally have higher volts. Welding currents with smaller diameter wires up to about 1/16 in. may reach 500 amp. Wires 1/16 in. diameter and larger call for welding currents up to 750 amp. As in TIG spot welding, special controls are required. These are normally incorporated in a special wire drive and control unit.

Plasma Arc Welding
An extension of the commercially accepted plasma cutting technique and similar in some respects to the tungsten inert gas process, it has fewer limitations than such methods as electron beam, laser and ultrasonic welding and at far less initial cost.

Like tungsten inert gas welding, plasma welding is normally a fusion process although cold filler wire may also be employed, depending on the job. The electrode is tungsten coated and water cooled because of the high temperatures involved. The process differs from TIG welding in that, in addition to a shielding gas, a plasma forming gas is also involved. The plasma ’flame’, sometimes in combination with an electric arc can produce extremely high temperatures. The flow of the plasma can be focused by the design of the torch and as a result can be highly concentrated to produce deep, narrow penetration.