Volume 01 - Properties and Selection Irons, Steels, and High-Performance Alloys Episode 8 pps

The carbon and low-alloy steels group comprises a large number of steels that differ in chemical composition, strength, heat treatment, corrosion resistance, and weldability.. Thus, weld

8740 65 184-217

(a) Ratings are for cold-finished bars

(b) Microstructure composed of ferrite and lamellar pearlite

(c) Microstructure composed mainly of acicular pearlite and bainite

(d) Microstructure composed primarily of spheroidite

Sample Hardness, Heat

HB treatment

A 206 Spheroidized

&fcirc; B 221 Annealed

C 321 Normalized

D 400 Hardened and tempered

E 500 Hardened and tempered

F 515 Hardened and tempered

Fig 15 Effect of hardness on tool life curves Workpiece: 4340 steel Tool material: C6 carbide Source: Ref 25

A comparison of the machinability ratings with the compositions of these steels indicates that all of the alloying elements that increase the hardenability of the steel decrease machinability; ferrite-strengthening elements such as nickel and silicon decrease the machinability more than equivalent amounts of carbide-forming elements such as chromium and molybdenum

It is not uncommon for heat-treating considerations to overshadow both machining and material costs in the selection of steel On occasion, heat-treating responses may dictate the selection of a less machinable or a more expensive steel so that the lowest total costs can be realized

The sulfur content of through-hardening alloy steels can significantly affect machining behavior Variations in residual sulfur level can account for unexplained differences in the machining behavior of different lots of the same material Many grades of hardenable alloy steels can be obtained in the resulfurized condition The differences in tool life and cutting speed between standard and high-sulfur 4150 steels are substantial Tests by Field and Zlatin (Ref 26) showed that raising the sulfur content from 0.04 to 0.09% increased the cutting speed for 60-min tool life by 25% Alloy steels containing lead are available and useful As indicated in Table 6, the machinability rating of the leaded grade 41L40 is 85, while the rating for 4140 is only 65 The performance of these two grades in several machining operations is indicated in Table 7 The data are from a case study described in Ref 26

Table 7 Effect of lead on cutting speed and tool life in machining alloy steels

In drilling standard 4140 steel, the 19 mm (3

4 in.) diam hole jammed with chips and the drill had to be removed frequently for cleaning When using leaded 4140 steel, the entire depth was drilled without removing the tool

Another important factor that can affect the choice of steel for a through-hardening application is the effect of alloying elements added for machinability on the mechanical properties of the steel These steels are often used at high-strength levels, where the deleterious effects of inclusions, particularly on transverse properties, might not be permissible The effect of sulfur, in the amounts usually specified for enhanced machinability, is generally considered to be more damaging than that of lead For some applications, neither machinability additive can be tolerated

References cited in this section

25 N Zlatin and J Christopher, Machining Characteristics of Difficult to Machine Materials, in Influence of Metallurgy on Machinability, V.A Tipnis, Ed., American Society for Metals, 1975, p 296-307

26 M Field and N Zlatin, Evaluation of Machinability of Rolled Steels, Forgings and Cast Irons, Theory and Practice, American Society for Metals, 1950, p 341-376

Machining Machinability of Steels

Francis W Boulger, Battelle-Columbus Laboratories (retired)

Cold-Drawn Steel

Cold drawing generally improves the machinability of steels containing less than about 0.2% C The improvement is most noticeable in plain carbon steels, as shown in Fig 16 The machinability of higher-carbon steels, or alloy steels, is less affected by cold work This improvement in machinability may be attributed to reduced cutting forces and/or the characteristics of chip removal Kopalinsky and Oxley (Ref 28) found that cold drawing lowered the cutting forces and improved the tool life and surface finish of low-carbon steels Screw machine tests by Yaguchi (Ref 29) showed that the workpiece surface finish improved continuously with increases in reduction in area up to 29% These effects were not characteristic of steels with high nitrogen contents (Ref 18) The improved machinability of cold-drawn steels can also be attributed to the decrease in ductility that results from cold working; thus, the chips are generally not long and stringy

Fig 16 Effect of cold drawing on tool life Workpiece: 1016 steel, 25 mm (1 in.) in diameter Machining

conditions: multiple-operation machined with a cutting speed of 0.73 m/s (144 sfm) Source:Ref 27

Cold-finished bars have closer dimensional tolerances, better surfaces, and usually, higher strength than hot-finished bars The first two factors may be significant in the selection of steels to be machined in multiple-operation machines or other high-production equipment These considerations are discussed in the article "Cold-Finished Steel Bars" in this Volume

The machining characteristics of cold-drawn steels are only rarely a decisive criterion for selection The extra strength obtained with cold-drawn steel may be more important from a cost standpoint, because it is often high enough to eliminate the need for heat treatment

References cited in this section

18 J.D Watson and R.H Davies, The Effects of Nitrogen on the Machinability of Low-Carbon

Free-Machining Steels, J Appl Metalwork., Vol 3 (No 2), 1984, p 110-119

27 J.D Armour, Metallurgy and Machinability of Steels, Machining Theory and Practice, American Society

for Metals, 1950, p 123-168

28 E.M Kopalinsky and P.L.B Oxley, Predicting Effects of Cold Working on Machining Characteristics of

Low-Carbon Steels, J Eng Ind (Trans ASME), Vol 109 (No 3), 1987, p 257-264

29 H Yaguchi and N Onodera, Effect of Cold Working on the Machinability of AISI 12L14 Steel, in

Strategies for Automation of Machining: Materials and Processes, Proceedings of an International

Conference (Orlando, FL), ASM INTERNATIONAL, 1987, p 15-26

Machinability of Steels

Francis W Boulger, Battelle-Columbus Laboratories (retired)

References

1 Machining Data Handbook, 3rd ed., Metcut Research Associates Inc., 1980

2 "Life Tests for Single-Point Tools of Sintered Carbide," B94.36-1956 (R 1971), American National Standards Institute

3 "Tool Life Testing With Single-Point Turning Tools," ANSI/ASME B94.55M-1985, American National Standards Institute

4 J.F Kahles, Elements of the Machining Process, in Metals Handbook: Desk Edition, American Society for

Metals, 1985, p 27.10

5 "Machining Performance of Ferrous Metals Using an Automatic Screw/Bar Machine," E 618-81-03.01,

Annual Book of ASTM Standards, American Society for Testing and Materials

6 F.W Boulger, Influence of Metallurgical Properties on Metal-Cutting Operations, Society of

Manufacturing Engineers, 1958

7 F.W Boulger and H.J Grover, Machinability Can Be Related to Composition, Tool Eng., Vol 40, March

1958

8 V.C Venkatesh and V Narayanan, Machinability Correlations Among Turning, Milling and Drilling

Processes, Ann CIRP, Vol 35 (No 1), 1986, p 59-62

9 T Araki et al., Some Results of Cooperative Research on the Effect of Heat Treated Structure on the Machinability of a Low Alloy Steel in Influence of Metallurgy on Machinability, V.A Tipnis, Ed.,

American Society for Metals, 1975, p 381-395

10 E.J.A Armarego and R.H Brown, The Machining of Metals, Prentice-Hall, 1969

11 J.E Mayer, Jr., and D.G Lee, Influence of Machinability on Productivity and Machining Cost, in

Influence of Metallurgy on Machinability, V.A Tipnis, Ed., American Society for Metals, 1975, p 31-54

12 F.W Boulger et al., Superior Machinability of MX Steel Explained, Iron Age, Vol 167, 17 May 1951, p

90-95

13 W.E Royer, Making Stainless More Machinable 303 Super X, Autom Mach., Vol 47 (No 5), May 1986,

p 47-49

14 H Yaguchi, Effect of MnS Inclusion Size on Machinability of Low-Carbon, Leaded, Resulfurized

Free-Machining Steel, J Appl Metalwork., Vol 3 (No 3), July 1986, p 214-225

15 S Abeyama et al., Development of Free Machining Steel With Controlled-Shape Sulfides, Bull Jpn Inst Met., Vol 24 (No 6), 1985, p 518-520

16 S Katayama et al., Improvements in Machinability of Continuously-Cast, Low-Carbon, Free-Cutting Steels, Trans ISI, Vol 25 (No 9), Sept 1985, p B229

17 R.M Welburn and D.J Naylor, Production and Machinability of Billet-Cast Medium Carbon High Sulfur

(Over 0.08%) Free-Machining Steels, in Proceedings of the Conference on Continuous Casting, Institute

of Metals, 1985

18 J.D Watson and R.H Davies, The Effects of Nitrogen on the Machinability of Low-Carbon

Free-Machining Steels, J Appl Metalwork., Vol 3 (No 2), 1984, p 110-119

19 H.J Tata and R.E Sampsell, Effects of Additions on Machinability and Properties of Alloy-Steels Bars,

Paper 730114, Trans SAE, Vol 82, 1973

20 R.A Joseph and V.A Tipnis, The Influence of Non-Metallic Inclusions on the Machinability of

Free-Machining Steels, in Influence of Metallurgy on Machinability, V.A Tipnis Ed., American Society for

Metals, 1975, p 55-72

21 S.V Subramanian and D.A.R Kay, Inclusions and Matrix Effects on the Machinability of Medium Carbon Steels, in Conference Proceedings, Ottawa, Ontario, Canada, Canadian Government Publishing Centre, 1985

22 T Kato, S Abeyama, A Kimura, and S Nakamura, The Effect of Ca Oxide Inclusions on the

Machinability of Heavy Duty Steels, in The Machinability of Engineering Materials, R.W Thompson,

Ed., Conference Proceedings, 13-15 Sept (Rosemont, IL), American Society for Metals, 1983, p 323-337

23 J Fombarlet, Improvement in the Machinability of Engineering Steels Through Modification of Oxide

Inclusions, in The Machinability of Engineering Materials, 13-15 Sept (Rosemont, IL), R.W Thompson,

Ed., Conference Proceedings, American Society for Metals, 1983, p 366-382

24 B Reh, U Finger et al., Development of Bismuth-Alloyed High Performance Easy Machining Steel, Neue

Hütte, Vol 31 (No 9), Sept 1986, p 327-330

25 N Zlatin and J Christopher, Machining Characteristics of Difficult to Machine Materials, in Influence of Metallurgy on Machinability, V.A Tipnis, Ed., American Society for Metals, 1975, p 296-307

26 M Field and N Zlatin, Evaluation of Machinability of Rolled Steels, Forgings and Cast Irons, -Theory and Practice, American Society for Metals, 1950, p 341-376

Machining-27 J.D Armour, Metallurgy and Machinability of Steels, Machining Theory and Practice, American Society

for Metals, 1950, p 123-168

28 E.M Kopalinsky and P.L.B Oxley, Predicting Effects of Cold Working on Machining Characteristics of

Low-Carbon Steels, J Eng Ind (Trans ASME), Vol 109 (No 3), 1987, p 257-264

29 H Yaguchi and N Onodera, Effect of Cold Working on the Machinability of AISI 12L14 Steel, in

Strategies for Automation of Machining: Materials and Processes, Proceedings of an International

Conference (Orlando, FL), ASM INTERNATIONAL, 1987, p 15-26

Weldability of Steels

S Liu, Center for Welding and Joining Research, Colorado School of Mines; J.E Indacochea, Department of Civil Engineering, Mechanics, and Metallurgy, University of Illinois at Chicago

Introduction

THE MAIN OBJECTIVE of this article is to survey the factors controlling the weldability of carbon and low-alloy steels

in arc welding A good understanding of the chemical and physical phenomena that occur in the weldment is necessary for welding modern steels Therefore, the influence of operational parameters, thermal cycles, and metallurgical factors

on weld metal transformations and the susceptibility to hot and cold cracking are discussed Common tests to determine steel weldability are also described

The carbon and low-alloy steels group comprises a large number of steels that differ in chemical composition, strength, heat treatment, corrosion resistance, and weldability These steels can be further divided into subgroups:

• Carbon steels

• High-strength low-alloy (HSLA) steels

• Quenched and tempered (QT) steels

• Heat-treatable low-alloy (HTLA) steels

S Liu, Center for Welding and Joining Research, Colorado School of Mines; J.E Indacochea, Department of Civil Engineering, Mechanics, and Metallurgy, University of Illinois at Chicago

Characteristic Features of Welds

Single-Pass Weldments. To understand weldability, it is necessary to recognize the various weld regions In the case

of a single-pass bead, the weldment is generally divided into two main regions: the fusion zone, or weld metal, and the heat-affected zone (HAZ), as shown in Fig 1 Within the fusion zone, the peak temperature exceeds the melting point of the base metal, and the chemical composition of the weld metal will depend on the choice of welding consumables, the base metal dilution ration, and the operating conditions

Fig 1 Various regions of a bead-on-plate weld

Under conditions of rapid cooling and solidification in the weld metal, alloying and impurity elements segregate extensively to the center of the interdendritic or intercellular regions and to the center parts of the weld, resulting in significant local chemical inhomogeneities Accordingly, the transformation behavior of the weld metal may be quite different from that of the base metal, even when the bulk chemical composition is not significantly changed by the welding process The typical anisotropic nature of the solidified weld and structure is also shown in Fig 1

The chemical composition remains largely unchanged in the HAZ because the peak temperature remains below the melting point of the parent plate Nevertheless, considerable microstructural change takes place within the HAZ during welding as a result of the extremely harsh thermal cycles The material immediately adjacent to the fusion zone is heated high into the austenitic temperature range The microalloy precipitates that development in the previous stages of processing will generally dissolve, and unpinning of austenite grain boundaries occurs with substantial growth of the grains, forming the coarse-grain HAZ The average size of the austenite grains, which is a function of the peak temperature attained, decrease with increasing distance from the fusion zone The cooling rate also varies from point to point in the HAZ; it increases with increasing peak temperature at constant heat input and decreases with increasing heat input at constant peak temperature Because of varying thermal conditions as a function of distance from the fusion line, the HAZ is actually composed of coarse-grain zones (CGHAZ), fine-grain zones (FGHAZ), intercritical zones (ICHAZ), and subcritical zones (SCHAZ) The various HAZ regions of a single-pass low-carbon steel butt weld are shown in Fig 2

Fig 2 Various regions of the HAZ of a single-pass low-carbon steel weld metal with 0.15 wt% C

In multipass weldments, the situation is much more complex because of the presence of reheated zones within the fusion zone, as shown in Fig 3 The partial refinement of the microstructure by subsequent weld passes increases the inhomogeneity of the various regions with respect to microstructure and mechanical properties Reaustenitization and subcritical heating can have a profound effect on the subsequent structures and properties of the HAZ Toughness property deterioration is related to small regions of limited ductility and low cleavage resistance within the CGHAZ that are known as the localized brittle zones (LBZ) Localized brittle zones consist of unaltered CGHAZ, intercritically reheated coarse-grain (IRCG) heat-affected zone, and subcritically reheated coarse-grain (SRCG) heat-affected zone At

an adjacent fusion line, that LBZs may be aligned, as shown in Fig 3 The aligned LBZs offer short and easy paths for crack propagation Fracture occurs along the fusion line

Fig 3 Overlapping of HAZ to form localized brittle zones aligned along the fusion line

Weldability of Steels

S Liu, Center for Welding and Joining Research, Colorado School of Mines; J.E Indacochea, Department of Civil Engineering, Mechanics, and Metallurgy, University of Illinois at Chicago

Metallurgical Factors That Affect Weldability

Hardenability and Weldability. Hardenability in steels is generally used to indicate austenite stability with alloy additions However, it has also been used as an indicator of weldability and as a guide for selection a material and welding process to avoid excessive hardness and cracking in the HAZ Steels with high hardness often contain a high

volume fraction of martensite, which is extremely susceptible to cracking during processing Hardenability is also used to indicate the susceptibility of a steel to hydrogen-induced cracking

Traditionally, empirical equations have been developed experimentally to express weldability Carbon equivalent (CE) is one such expression; it was developed to estimate the cracking susceptibility of a steel during welding and to determine whether the steel needs pre- and postweld heat treatment to avoid cracking Carbon equivalent equations do include the hardenability effect of the alloying elements by expressing the chemical composition of the steel as a sum of weighted alloy contents To date, several CE expressions with different coefficients for the alloying elements have been reported The International Institute of Welding (IIW) carbon equivalent equation is:

(Eq 1)

where the concentration of the alloying elements is given in weight percent It can be seen in Eq 1 that carbon is the element that most affects weldability Together with other chemical elements, carbon may affect the solidification temperature range, hot tear susceptibility, hardenability, and cold-cracking behavior of a steel weldment Figure 4 summarizes the CE and weldability description of some steel families Because of the simplification and generalization involved in Fig 4, it should be used cautiously for actual welding situations

Fig 4 Weldability of several families of steels as a function of carbon equivalent 1, Mo; 2, Cr + Ni + Mo + Si,

and so on; 3, Cr or V or Ni + Si, and so on

The application of CE expressions is also empirical For example, the IIW carbon equivalent equation has been used successfully with traditional medium-carbon low-alloy steels Steels with lower CE values generally exhibit good weldability When the CE of a steel is less than 0.45 wt%, weld cracking is unlikely, and no heat treatment is required When the CE is between 0.45 and 0.60 wt%, weld cracking is likely, and preheat in the range of approximately 95 to 400

°C, (200 to 750 °F) is generally recommended When the CE of a steel is greater than 0.60 wt%, there is a high

probability that the weld will crack and that both preheat and postweld heat treatments will be required to obtain a sound weld

However, Eq 1 does not accurately correlate with the microstructures and properties of newly developed low-carbon microalloyed steels over extended alloy ranges Thus, new expressions based on solution thermodynamics and kinetic considerations were developed to obtain better predictions of the alloy behavior and weldability of low-carbon low-alloy steels Complex interactive terms, rather than simple additive forms, are included in these equations An example of one such expression is:

CE = k1C[1 + k2C + k3Mn +

+ k11 ln C + k22C ln C + k33Mn ln Mn

+ + k111CMn + ]

(Eq 2)

where k1, k2, , and so on, are the weighted coefficients multiplied to the concentration of the alloying elements

Nonlinear terms such as ln Xi, Xi ln Xi, and XiXj represent the interaction effect among the alloying elements Xi and Xj Equations with these nonlinear terms are more useful in predicting arc welding behavior

Several expressions are also available for other steel groups with a wider range of alloying elements and with different prior heat treatments, hydrogen contents, and weld hardnesses Recently, expressions that include fabrication conditions such as heat input, cooling rate, joint design and restraint conditions have also been proposed An example of this type of equation is:

(Eq 3)

where PH is the cracking susceptibility parameter, H is the concentration of hydrogen (in parts per million), Rf is the restraint stress (in megapascals), and:

(Eq 4)

The thickness of the part being welded can also be related to CE as a compensated carbon equivalent (CCE) as follows:

where e is the thickness of the part (in millimeters) Equations 3, 4, and 5 are valid only for specific ranges of chemical

composition and welding conditions Nevertheless, despite the different forms and terms included in the predictive equations, the main objective remains that of estimating the weldability and cracking susceptibility of the material

Weld Metal Microstructure. Inherent in the welding process is the formation of a pool of molten metal directly below

a moving heat source The shape of this molten pool is determined by the flow of both heat and metal, with melting occurring ahead of the heat source and solidification occurring behind it Heat input determines the volume of molten metal and therefore the dilution and weld metal composition, as well as the thermal conditions under which solidification takes place Also important to solidification is the crystalline growth rate, which is geometrically related to weld travel speed and weld pool shape Thus, weld pool shape, weld metal composition, cooling rate, and growth rate are all factors that are interrelated with heat input, which in turn will affect the solidification microstructure and the tolerance of the weldment to hot cracking

Incipient melting at base metal grain boundaries immediately adjacent to the fusion zone allows these grains to serve as seed crystals for epitaxial grain growth during weld metal solidification The continuous growth of the epitaxial grains results in large columnar grains whose boundaries provide easy paths for crack propagation An elongated weld pool will

yield straight and broad columnar grains, which promote the formation of centerline cracking because of impurity segregation, mechanical entrapment of inclusions, and the shrinkage stresses that develop during solidification Epitaxial columnar growth is particularly deleterious in multipass welds where grains can extend continuously from one weld bead

to another

Hot tears originate near the liquid/solid interface when strains from solidification shrinkage and thermal contraction cause rupture of the liquid films of low melting point located at grain boundaries The susceptibility of an alloy to hot tearing is related to its inability to accommodate strain through dendrite interlocking as well as the tendency of tears to backfill with the remaining liquid The time interval during which liquid films can exist in relation to the rate of strain generation may also play a role in hot tear susceptibility Ferrous alloys can be hot tear sensitive depending on the amount of phosphorus and sulfur impurities they contain Carbon and nickel are also known to influence hot cracking in steel welding

When the solidified steel weld metal cools down, solid-state transformation reactions may occur As in solidification, the two main factors that determine the final microstructure are the chemical composition and thermal cycle of the weld metal In most structural steels, weld metal will solidify as δ-ferrite At the peritectic temperature, austenite will form from the reaction between liquid weld metal and δ-ferrite, and subsequent cooling will lead to the formation of α-ferrite During the austenite-to-ferrite transformation, proeutectoid ferrite forms first along the austenite grain boundaries; this is known as grain-boundary ferrite Subsequent to grain-boundary ferrite formation, ferrite sideplates develop in the form of long needlelike ferrite laths that protrude from the allotriomorphs A coarse austenite grain size and a low carbon content,

in combination with a relatively high degree of supercooling, are found to promote ferrite sideplate formation These laths can be properly characterized by their length-to-width aspect ratios; values above 10:1 are common

As the temperature continues to drop, intragranular acicular ferrite will nucleate and grow in the form of short laths separated by high angle boundaries The inclination between orientations of adjacent acicular ferrite laths is usually larger than 20° The random orientation of these laths provides good resistance to crack propagation Acicular ferrite laths have aspect ratios ranging from 3:1 to 10:1

During proeutectoid ferrite formation, carbon is rejected continuously from the ferrite phase, enriching the remaining austenite, which later transforms into a variety of constituents, such as martensite (both lath and twinned), bainite, pearlite, and retained austenite Because of the acicular nature of the bainite laths, they can also be described by their aspect ratio, with values similar to those of Widmanstätten side-plates More frequently, however, bainite laths occur in the form of packets associated with grain boundaries Figure 5 illustrates the microstructure of a low-carbon steel weld metal

Fig 5 Weld metal microstructure of HSLA steel A, grain-boundary ferrite; B, acicular ferrite; C, bainite; D,

sideplate ferrite

Heat-Affected Zone Microstructure. In terms of microstructure, long bainite laths with alternate layers of connected martensite islands are generally found in the CGHAZ of high-strength low-alloy steel weldments Martensite islands (martensite-retained austenite constituents) are formed because of the enrichment of carbon in austenite in the intercritical zone Coarse austenite grain size in the near-fusion region of the HAZ can suppress high-temperature transformation products in favor of martensite and bainite upon cooling Upper bainite has a relatively high transformation temperature and is stable relative to the thermal cycles subsequent to those of the first pass Fluctuation of the chemical composition of the microalloying elements could also contribute to carbon equivalent change and to the amount of hard martensite present in the CGHAZ

In the FGHAZ, even though the peak temperature attained is above thermal cycle Ac3, it is still well below the coarsening temperature The smaller prior-austenite grain size and subsequent ferrite transformation produce a refined microstructure having grains smaller than those of the parent material The microstructure is similar to that of a normalized steel, with considerable toughness

grain-Only partial transformation takes place in the ICHAZ, resulting in a mixture of austenite and ferrite at the peak temperature of the thermal cycle Upon cooling, the austenite in a matrix of soft ferrite decomposes, and the final microstructure depends on the bulk and local composition of the alloying elements The cooling rate is also an important factor in determining the amount of martensite and bainite in the ferrite matrix In the SCHAZ, no observable microstructural changes are observed Some spheroidization of carbides may occur

Upon reheating by subsequent weld passes, precipitates or preprecipitate clusters may form, reducing the toughness Irregularly shaped particles may also coalesce and strain the surrounding matrix, further lowering the toughness During HAZ thermal cycles between Ac1 and Ac3, the austenite becomes enriched with carbon, which, upon cooling, transforms

to martensite islands In the as-welded condition, this transformation affects the IRCG region more than the other reheated zone Figure 6 illustrates the different phases that can be found in a low-carbon steel HAZ

Fig 6 Heat-affected zone exhibiting a wide variety of microstructures in the intercritical and subcritical regions

A, spheroidized carbides; B, bainite and martensite

Chemical Composition Effect. The presence of a certain phase in the final microstructure of a weldment can be explained by means of a continuous cooling transformation (CCT) diagram, which is formed by two sets of curves: the percent transformation curves and the cooling curves The percent transformation curves define the regions of stability of the different phases The cooling curves represent the actual thermal conditions that the weld experienced The intersection of these two sets of curves determines the final microstructure of the different weld zones Figure 7 illustrates the use of a CCT diagram to determine the microstructure of a low-carbon low-alloy steel weld metal

Fig 7 Continuous cooling transformation diagram for an HSLA steel weld metal showing the effect of cooling

rate and chemical composition on microstructure CR, cooling rate

Hardenability elements, such as carbon, manganese, chromium, and molybdenum, suppress the start of austenite decomposition to lower temperatures This is equivalent to pushing the transformation curves to the right side of the CCT diagram, resulting in a refined microstructure Inclusion formers, such as oxygen and sulfur, accelerate the austenite-to-ferrite transformation by providing more nucleation sites for the reaction to initiate at higher temperatures

Faster cooling has the same effect as an increase in hardenability elements, while a slower cooling rate acts in the same direction as a decrease in hardenability agents or an increase in nucleation site providers Because the cooling rate varies from point to point in the HAZ, the microstructure also changes accordingly, with martensite and bainite in regions close

to the fusion line

Preweld and Postweld Heat Treatments. In the welding of carbon and low-alloy steels, the final microstructure of the weldment is primarily determined by the cooling rate from the peak temperature Because the alloy level in carbon and low-alloy steel is low, the major physical properties of the steel are not affected Thus, temperature gradient and heat input become the important parameters in weld metal microstructural evolution A slower cooling rate decreases shrinkage stress, prevents excessive hardening, and allows time for hydrogen diffusion Cooling rate (CR) is of particular

importance and is a function of the difference in temperature, ∆T, as well as the thermal conductivity of the material, k

The cooling rate can be expressed for thinplate and thick-plate welding, respectively, as:

CR k ∆T3

During preheating, the initial temperature of the plate increases, decreasing the cooling rate and the amount of the hard phases, such as martensite and bainite, in the weld microstructure For the welding of hardenable steels, it is important to determine the critical cooling rate (CCR) that the base metal can tolerate without cracking:

(Eq 7)

The higher the carbon equivalent of an alloy, the lower the critical or allowable cooling rate The use of a low-hydrogen welding electrode also becomes more important Preheating should be applied to adjust the cooling rate accordingly

in the regions next to the SCHAZ and the lowest at or next to the fusion line As a result, there is a slight increase in microalloying element in solution in the CGHAZ, which increases the hardenability of this region

Hydrogen-Induced Cracking. The effect of hydrogen on weld cracking should also be mentioned Moisture pickup from the atmosphere that is incorporated into the molten puddle, either directly or via the welding consumables, is the main source of hydrogen The presence of hydrogen increases the HAZ cracking susceptibility of high-strength steel weldments Also known as underbead, cold, or delayed cracking, it is perhaps the most serious and least understood of all weld-cracking problems It generally occurs at the temperature below approximately 95 °C (200 °F) either immediately upon cooling or after a period of several hours The crack can be both transgranular and intercrystalline in character, but mainly follows prior-austenite grain boundaries The initiation of cold cracking is particularly associated with notches, such as the toe of the weld, or with inhomogeneities in microstructure that exhibit sudden changes in hardness, such as slag inclusions, martensite/ferrite interfaces, or even grain boundaries Like most other crack growth phenomena, hydrogen-induced cracking is accentuated in the presence of high-restraint weld geometries and matrix hardening Such cracking is associated with the combined presence of three factors:

• The presence of hydrogen in the steel (even very small amounts, measured in parts per million)

• A microstructure that is partly or wholly martensitic

• High residual stresses (generally as a result of thick material)

If one or more of these conditions is absent or at a low level, hydrogen-induced cracking will not occur However, high cooling rates such as those found in manual processes further enhance the probability of weld HAZ cold cracking

The tolerance of steels for hydrogen decreases with increasing carbon or alloy content Hydrogen-induced cracking can be controlled by choosing a welding process or an electrode that produces little or no hydrogen Postweld heat treatments can

be used to decrease or eliminate the residual hydrogen or to produce a microstructure that is insensitive to hydrogen cracking Finally, welding procedures that result in low restraining stresses will also reduce the risk of weld cracking

Stress-relief cracking due to reheating is of concern when welding quenched and tempered grades and heat-resistant steels containing significant levels of carbide formers, such as chromium, molybdenum, and vanadium When weldments

of these steels are heated above approximately 510 °C (950 °F), intergranular cracking along the prior-austenite grain boundaries may take place in the CGHAZ Also known as reheat cracking and stress-rupture cracking, stress-relief cracking is thought to be closely related to the phenomenon of creep rupture Furthermore, during reheating, the reprecipitation of carbides is likely to occur, further increasing the hardness The precipitation of carbides during stress relaxation alters the delicate balance between resistance to grain-boundary sliding and resistance to deformation within the coarse grains of the heat-affected zone

Some procedures that can be used singly or in combination to decrease stress-relief cracking in steels include the selection

of a more appropriate weld joint design, weld location, and sequence of assembly to minimize restraint and stress concentrations Selecting a filler metal that will provide a weld metal that has significantly lower strength than that of the HAZ at the heat-treating temperature is another way to minimize stress-relief cracking Peening each layer of weld metal

to generate a surface compressive stress state that counteracts shrinkage stresses is also very effective

Lamellar cracking, better known as lamelar tearing, is characterized by a steplike crack parallel to the rolling plane Figure 8 shows a typical feature of lamellar tearing, the horizontal and vertical cracking of the base plate The problem occurs particularly when making tee and corner joints in thick plates such that the fusion boundary of the weld runs parallel to the plate surface High tensile stresses can develop perpendicular to the midplane of the steel plate, as well as parallel to it This tearing, usually associated with inclusions in the steel, progresses from one inclusion to another

Fig 8 Lamellar tear caused by thermal contraction strain

There is some evidence that sensitivity to lamellar tearing is increased by the presence of hydrogen in the steel Inclusions that contain low-melting compounds, such as those of sulfur and phosphorus, also increase the sensitivity of steel to lamellar tearing by wetting the prior-austenite grain boundaries; this makes them too weak and fragile to withstand the thermal stresses during cooling Some approaches that can minimize lamellar tearing are:

• Changing the location and design of the welded joint to minimize through-thickness strains

• Using a lower-strength weld metal

• Reducing available hydrogen

• Buttering the surface of the plate with weld metal prior to making the weld

• Using preheat and interpass temperatures of at least 95 °C (200 °F)

• Using base plates with inclusion shape control

Hot cracking, or solidification cracking, occurs at elevated temperatures and is usually located in the weld metal Hot cracking also can be found in the HAZ, where it is known as liquation cracking Solidification cracking in weld deposits during cooling occurs predominately at the weld centerline or between columnar grains The fracture path of a hot crack

is intergranular The causes of solidification cracking are well understood The partition and rejection of alloying elements at columnar grain boundaries and ahead of the advancing solid/liquid interface produce significant segregation The elements of segregation form low-melting phases or eutectic structures to produce highly wetting films at grain

boundaries They weaken the structure to the extent that cracks form at the boundaries under the influence of the tensile residual stresses during cooling Liquation cracking is also associated with grain-boundary segregation and is aggravated

by the melting of these boundaries near the fusion line These impurity-weakened boundaries tend to rupture as the weld cools because of the high residual stresses

Inclusions. Large amounts of sulfur and phosphorus are added to some steels to provide free-machining characteristics These steels have relatively poor weldability because of hot tearing in the weld metal caused by low-melting compounds

of phosphorus and sulfur at the grain boundaries Iron oxide and iron sulfide inclusions, if present, are also harmful because of their solubility change with temperature and their propensity to precipitate at grain boundaries, contributing to low ductility, cracking, and porosity Laminations, which are flat separations or weaknesses that sometimes occur beneath and parallel to the surface of rolled products, have a slight tendency to open up if they extend to the weld joint

For low-carbon steels, a low-alloy filler metal is generally recommended for meeting mechanical property requirements The general procedure is to match the filler with the base metal in terms of strength or, for dissimilar welds, to match the lower-strength material Often, however, higher-strength weld metal may actually require a softer HAZ to undergo a relatively large amount of strain when the joint is subjected to deformation near room temperature Nevertheless, a low-strength filler metal should not be used indiscriminately as a remedy for cracking difficulties

Medium-Carbon Steels. If steel containing about 0.5 wt% C is welded by a procedure commonly used for low-carbon steel, the heat-affected zone is likely to be hard, low in toughness, and susceptible to cold cracking As indicated previously, preheating the base metal can greatly reduce the rate at which the weld area cools, thus reducing the likelihood of martensite formation Postheating can further retard the cooling of the weld or can temper any martensite that might have formed

The appropriate preheat temperature depends on the carbon equivalent of the steel, the joint thicknesses, and the welding procedure With a carbon equivalent in the 0.45 to 0.60 wt% range, a preheat temperature in the range of approximately

95 to 100 °C (200 to 400 °F) is generally recommended The minimum interpass temperature should be the same as the preheat temperature A low-hydrogen welding procedure is mandatory with these steels Modifications in welding procedure, such as the use of a larger V-groove or of multiple passes, also decrease the cooling rate and the probability of weld cracking

Dilution can be minimized by depositing small weld beads or by using a welding procedure that provides shallow penetration This is done to minimize carbon pickup from the base metal and the amount of hard transformation products

in the fusion zone Low heat input to limit dilution is also recommended for the first few layers in a multipass weld

High-carbon steels generally contain over 0.60 wt% C and exhibit a very high elastic limit They are often used in applications where high wear resistance is required These steels have high hardenability and sensitivity to cracking in both the weld metal and the HAZ A low-hydrogen welding procedure must be used for arc welding Preheat and postheat will not actually retard the formation of brittle high-carbon martensite in the weld However, preheating can minimize shrinkage stresses, and postheating can temper the martensite that forms Successful welding of high-carbon steel requires the development of a specific welding procedure for each application The composition, thickness, and configuration of the component parts must be considered in process and consumable selections

High-strength low-alloy steels are designed to meet specific mechanical properties rather than a chemical composition the alloy additions to HSLA steels strengthen the ferrite, promote hardenability, and help to control grain size Weldability decreases as yield strength increases For all practical purposes, welding these steels is the same as welding plain carbon steels that have similar carbon equivalents Preheating may sometimes be required, but postheating

is almost never required

Quenched and tempered steels are furnished in the heat-treated condition with yield strengths ranging from approximately 350 to 1000 MPa (50 to 150 ksi), depending on the composition The base metal is kept at less than 0.22%

C for good weldability Preheating must be used with caution when welding QT steels because it reduces the cooling rate

of the weld HAZ If the cooling rate is too slow, the reaustenitized zone adjacent to the weld metal can transform either to ferrite with regions of high-carbon martensite, or to coarse bainite, of lower strength and toughness A moderate preheat, however, can ensure against cracking, especially when the joint to be welded is thick and highly restrained A postweld stress-relief heat treatment is generally not required to prevent brittle fracture in weld joints in most QT steels

Heat-treatable low-alloy steels. Examples of HTLA steels include AISI 4140, AISI 4340, AISI 5140, AISI 8640, and 300M The high hardness of these steels requires that welding be conducted on materials in an annealed or overtempered condition, followed by heat treatment to counter martensite formation and cold cracking However, high preheating is often used with a low-hydrogen process on these steels in a quenched and tempered condition, as in motor shaft applications Preheating, or interpass heating, for both the weld metal and the HAZ are recommended Hydrogen control is also essential to prevent weld cracking Extremely clean vacuum-melted steels are preferred for welding

Low sulfur and phosphorus, as described previously, are required to reduce hot cracking Segregation, which occurs because of the extended temperature range at which solidification takes place, reduces high-temperature strength and ductility Fillers of lower carbon and alloy content are highly recommended Preheat and interpass temperatures of 315 °C (600 °F) or higher are very harsh environments for welders because of the physical discomfort and because an oxide layer forms at the weld joint However, the cooling rate must be controlled to allow the formation of a bainitic microstructure instead of the hard martensite The bainitic microstructure can be heat treated afterward to restore the original mechanical properties of the structure Specifications and procedures should be followed rigorously for difficult-to-weld materials

Precoated Steels. Thin plates and steel sheets are often precoated to protect them from oxidation and corrosion The coatings commonly used are aluminum (aluminized), zinc (galvanized), and zinc-rich primers As expected, the coating originally at the weld region is destroyed during fusion welding, and the effectiveness of the coating adjacent to the weld

is significantly decreased by the welding heat In the case of aluminized steels, the formation of aluminum oxide may adversely affect the wetting and weld pool shape The welding electrode and filler metals should be selected carefully A basic coating shielded metal arc (SMA) welding electrode is recommended

For galvanized steels, weld cracking is generally attributed to intergranular penetration by zinc Zinc dissolves considerably in iron to form an intermetallic compound at temperatures close to the melting temperature of zinc Thus, molten zinc penetrates along the grain boundaries, leaving behind a brittle champagne fracture during cooling with the onset of a tensile stress state Cracking occurs primarily at the throat region of a fillet weld, where shrinkage strain is more significant The use of hot-dipped coatings results in more severe cracking, while thin electrogalvanized coatings are the least susceptible to cracking

Low-silicon electrodes and rutile-base SMA welding rods are both good for galvanized steel welding Specific welding and setup procedures should be followed, such as removing the zinc coating by an oxy-fuel process or by grinding, ensuring a large root opening, and using a slower welding speed to allow zinc vaporization and to prevent zinc entrapment in the weld metal Adequate ventilation and fuel extraction should be mandatory in welding galvanized steels because of the health hazard of zinc fumes

Weldability of Steels

S Liu, Center for Welding and Joining Research, Colorado School of Mines; J.E Indacochea, Department of Civil Engineering, Mechanics, and Metallurgy, University of Illinois at Chicago

Weldability Tests

Weldability tests conducted to provide information on the service and performance of welds However, the data obtained

in these tests can also be applied to the design of useful structures Frequently these data are obtained from the same type

of test specimens used in determining the base metal properties Predicting the performance of structures from a laboratory-type test is very complex because of the nature of the joint, which is far from homogeneous, metallurgically or chemically Along with the base metal, the weld joint consists of the weld metal and the HAZ Thus, a variety of properties are to be expected throughout the welded joint Careful interpretation and application of the test results are required

There are currently many tests that evaluate not only the strength requirements of steel structures, but also the fracture characteristics and the effect of environmental conditions on early failure of the weldments Selected major tests are described below

Weld Tension Test. To obtain an accurate assessment of the strength and ductility of welds, several tension test specimens can be used; all weld metal specimens, transverse weld specimens, and longitudinal weld specimens are shown

in Fig 9 In the all weld metal test, base metal dilution should be minimized if the test is to be representative of the weld metal However, the resulting properties may not be easy to translate into those properties achievable from welds made in

an actual weld joint

Fig 9 Typical tension test specimens for evaluating welded joints Both plate-type specimens have identical

dimensions All dimensions given in millimeters

Interpreting test results for the transverse butt weld test is complicated by the different strengths and ductilities generally found in the various regions of the joint The primary information gained from the test is the ultimate tensile strength Yield strength and elongation requirements are generally not specified

Tests of HAZ properties that are unaffected by the presence of either base metal or weld metal are not easy to conduct because it is practically impossible to obtain specimens made up entirely of the HAZ In addition, as indicated earlier, the HAZ is composed of various regions, each with its own distinct properties Simulated HAZ specimens that are generated and tested using a Gleeble thermomechanical testing system can be used to provide a more accurate assessment of the tensile properties of this region

Bend Test. Different types of bend tests are used to evaluate the ductility and soundness of welded joints Bend test results are expressed in various terms, such as percent elongation in outer fibers, minimum bend radius prior to failure, go/no-go for specific test conditions, and angle of bend prior to failure Various specimen designs, both notched and unnotched, and testing techniques have been used Today, unnotched specimens can be used in quality control tests, while notched specimens may be used to predict in-service behavior; however, most notched bend tests are used for research purposes and are not in common industrial use Transverse bend tests are useful because they quite often reveal the presence of defects that are not detected in tension tests However, the transverse specimen suffers from the same weakness as the transverse weld tension test specimen in that nonuniform properties along the length of the specimen can cause nonuniform bending, although this is often compensated for by the use of a wraparound bend fixture

Hardness testing can be used to complement information gained through tension or bend tests by providing information about the metallurgical changes caused by welding Routine methods for the hardness testing of metals are well established In carbon and low-alloy steels, the hardness near the fusion line in the HAZ may be much higher than in the base metal because of the formation of martensite In the HAZ areas where the temperature is low, the hardness may be lower than in the base metal because of tempering effects

The drop-weight test design is based on service failures resulting from brittle fracture initiation at a small flaw located in a region of high stress The drop-weight test can be considered a limited-deflection bend test that uses a crack starter to introduce a running crack in the specimen The specimen is a bar on which a brittle crack starter weld is deposited This overlay cracks when the bar is deflected by the drop weight A series of test is performed at different temperatures to determine the testing temperature below which the crack will propagate to the edges of the specimen This critical temperature is also called the nilductility temperature (NDT), defined as the highest temperature at which the propagating crack reaches the edge of the specimen Therefore, the drop-weight test is also known as the NDT test

The Charpy V-notch (CVN) test is the most popular technique for evaluating the impact properties of welds The energy absorbed by a sample at fracture determines the toughness of the specimen In this test, specimens at different temperatures are broken using a pendulum hammer A typical plot of CVN results for a carbon and low-alloy steel is illustrated in Fig 10 The plot shows that there is a transition from low- to high-energy fracture over a narrow temperature range; this is associated with a change from trans-crystalline to ductile fracture Therefore, material quality can be defined

in terms of this transition temperature

Fig 10 Schematic impact transition curve for steel FATT50 , fracture appearance transition temperature

In the CVN testing of welds, the notch is typically located at the weld centerline For CVN testing of the heat-affected zone, the notch is more typically introduced at the CGHAZ However, because precise location of a notch is never simple

in the HAZ, simulated weld samples are used instead

The crack tip opening displacement test measures toughness, primarily for elastic-plastic conditions In CTOD tests, the clip gage opening at the onset of fracture is measured and used to calculate the crack opening displacement at

the crack tip The critical value of CTOD at fracture, ∆c, is a critical strain parameter that is analogous to the critical stress-intensity parameter KIc The CTOD test provides a useful method of determining the critical flaw size Nevertheless, the test is very sensitive to changes in sample thickness, hardness, and strength, and it is difficult to obtain valid results in practical specimen thicknesses

The application of fracture mechanics to the prevention of catastrophic failure in weldments is, however, complicated by the nature of the weldment In addition to their metallurgical heterogeneity, weldments often contain high residual stresses Consequently, it is inadequate to fracture test the base metal and assume that the critical crack length thus determined is valid when the base metal is made into a weldment The fracture toughness criterion must be determined for the base metal, the HAZ, and the weld metal By first determining the zone with the lowest toughness value, it is then possible to evaluate a more realistic critical flaw size However, the plane-strain fracture toughness tests are preproduction or pilot plant type tests that provide a rational means for designs and engineers to estimate the effects of new designs, materials, or fabrication practices on the fracture-safe performance of structures Other popular tests include compact tension (CT) and wedge opening load (WOL) tests, which are commonly used in the evaluation of structural weldments Further discussion of CTOD and other fracture toughness testing of welds is available in the "Selected References" at the end of this article

Stress-Corrosion Cracking Test. The presence of corrosive environments in a steel weldment may accelerate the initiation of a crack Usually, the higher the strength of the steel, the more susceptible it becomes to stress-corrosion cracking The steels considered in this article are not usually exposed to severely corrosive environments, but rather to the atmosphere, moisture, hydrocarbons, fertilizers, and soils Nevertheless, welding can lower corrosive resistance by the introduction of:

• Compositional differences that promote galvanic attack between weld metal, HAZ, and base metal when the joint in immersed in a conducting liquid

• Residual stresses that can cause stress-corrosion cracking

• Surface flaws that can act as sites for stress-corrosion cracking

Stress-corrosion cracking is generally delayed cracking, with longer time to failure at lower stresses Most corrosion tests are fairly long in terms of time because of the slow crack initiation that occurs in unnotched test bars However, it has been found that the long initiation period can be eliminated by testing precracked specimens Additional information on the stress-corrosion cracking test is available in the "Selected References" at the end of this article

stress-Weldability of Steels

S Liu, Center for Welding and Joining Research, Colorado School of Mines; J.E Indacochea, Department of Civil Engineering, Mechanics, and Metallurgy, University of Illinois at Chicago

Fabrication Weldability Tests

There are various types of tests for determining the susceptibility of the weld joint to different types of cracking during fabrication They are:

• Restraint tests

• Externally loaded tests

• Underbead cracking tests

• Lamellar tearing tests

Table 1 summarizes the applications, controllable test variables, and typical test data of several fabrication weldability tests to illustrate the differences among them Of the many tests identified in Table 1, the Lehigh restraint test, the Varestraint test, and the controlled thermal severity test are described below

Table 1 Comparison of weldability tests for fabrication

Test Fields of use Controllable variables Type of data Specific

Joint geometry, process, filler metal, restraint level, heat input, preheat, postweld heat treatment

Critical restraint, or % hindered control

None Low cost

Rigid restraint

test

Weld metal hot and cold cracks, root cracks, HAZ hydrogen cracks

Joint geometry, process restraint level, filler metal, heat input, preheat

Critical restraint

Restraint jig Costly

machining and setup

Tekken test Weld metal root cracks,

HAZ hydrogen cracks

Joint geometry, process filler metal, heat input, preheat

Critical preheat None Low cost

Circular groove

test

Weld metal hot and cold cracks, HAZ hydrogen cracks

Process, filler metal, preheat Go/no-go None Costly

Loading jig Intermediate

Loading jig Costly

machining and setup

Varestraint test Weld metal and HAZ hot

None Costly

preparation

Cruciform test HAZ hydrogen cracks,

weld metal root cracks

Process, heat input, preheat, filler metal

Go/no-go None Costly

preparation

Loading jig Costly specimen

preparation

Cranfield test Lamellar tearing Filler metal Number of

passes to crack

None Low cost

Nick bend test Weld metal soundness Filler metal Go/no-go None Low cost

Source: Ref 1

The Lehigh restraint test (Fig 11) is particularly useful for quantitatively rating the crack susceptibility of a weld metal as affected by electrode variables This test provides a means of imposing a controllable severity of restraint on the root bead that is deposited in a butt weld groove with dimensions suitable to the application Slots are cut in the sides and ends of a plate prior to welding By changing the length of the slots, the degree of plate restraint on the weld is varied without significantly changing the cooling rate of the weld Therefore, a critical restraint for cracking can be determined for given welding conditions This sample is also useful for hydrogen cracking

Fig 11 Basic outline of a Lehigh test specimen

The Varestraint test (Fig 12) determines the susceptibility of the welded joint to hot cracking The test utilizes external loading to impose controlled plastic deformation in a plate while a weld bead is being deposited on the long axis

of the plate The specimen is mounted as a cantilever beam, and a pneumatically driven yoke is positioned to force the specimen downward when the welding arc reaches a predetermined position By the choice of the radius to which the plate is bent, the severity of deformation causing cracking can be determined Strain from 0 to 4% can be chosen according to the susceptibility of the joint to hot cracking When the bending moment is applied transverse to the weld

axis, the test is termed transvarestraint A spot Varestraint test can also be conducted by keeping the arc stationary; bending is applied at the moment the arc is extinguished

Fig 12 Schematic of the Varestraint test A, weld location; B, die; C, arc; D, load; r, radius of deformation

The controlled thermal severity test (Fig 13) is designed to measure the cracking sensitivity of steels under cooling rates controlled by the thickness of the plates and the number of paths available for dissipating the welding heat

It is conducted with a plate bolted and anchor welded to a second plate in a position to provide two fillet (lap) welds The fillet located at the plate edges has two paths of heat flow The lap weld located near the middle of the bottom plate has three paths of heat flow, thus inducing faster cooling The fillet welds are made first and allowed to cool, followed by the lap welds After a holding time of 72 h at room temperature, the degree of cracking is determined by measuring the crack length on metallographic specimens

Fig 13 Schematic of the controlled thermal severity test

A number of other tests have been developed that contain welds in a circular configuration The circular patch test has probably the most severe testing conditions; the two varieties are the Navy circular patch restraint test and the segmented circular patch restraint level Cracking is detected by visual, radiographic, and liquid penetrant inspection The cracking susceptibility of a material is measured as the total crack length and expressed as a percentage of the weld length These tests can be used to determine both hot and cold cracking in the weld metal and the HAZ Depending on the results, a go/no-go criteria is established for weld qualification Detailed descriptions of these tests can be found in the "Selected References" at the end of this article

Reference cited in this section

1 R Stout, Weldability of Steels, 4th ed., Welding Research Council, 1987

• O Blodgett, "Why Preheat? An Approach to Estimating Correct Preheat Temperature," Brochure

G-231, Lincoln Arc Welding Foundation, June 1970

• B.F Brown, "Stress Corrosion Cracking and Corrosion Fatigue of High Strength Steels," Report 210, Defense Metals Information Center, 1964, p 91-102

• "Classification of Microstructures in Low Carbon Low Alloy Steel Weld Metal and Terminology," DOC IX-1282-83, International Institute of Welding, 1983

• J Cornu, Advanced Welding Systems: Fundamentals of Fusion Welding Technology, IFS/Springer

• K Easterling, Introduction to the Physical Metallurgy of Welding, Butterworths, 1983

• D.P Fairchild, Brittle Zones in Structural Welds, in Welding Metallurgy of Structural Steels, J Koo,

Ed., The Metallurgical Society, 1987, p 303-318

• H Granjon, "Notes on the Carbon Equivalent," DOC IX-555-67, International Institute of Welding,

• J Koo and A Ozekan, Local Brittle Zone Microstructure, in Welding Metallurgy of Structural Steels, J

Koo, Ed., The Metallurgical Society, 1987, p 119-135

• S Kou, Welding Metallurgy, Wiley Interscience, 1987

• S Liu, D.L Olson, and D.K Matlock, A Thermodynamic and Kinetic Approach in the Development of

Expressions for Alloy Behavior Prediction, J Heat Treat., Vol 4 (No 4), 1986, p 309-316

• F Matsuda, T Hashimoto, and T Senda, Fundamental Investigations on Solidification Structure, Trans

Natl Res Inst Met (Jpn.), Vol 11, 1969, p 43-58

• H.G Pisarski and J Kudoh, Exploratory Studies on the Fracture Toughness of Multi-Pass Welds With

Locally Embrittled Regions, in Welding Metallurgy of Structural Steels, J Koo, Ed., The Metallurgical

• A.B Rothwell, CAN/MET Report 79-6, Can Weld Fabr., Vol 20, 1980

• C.P Royer, A User's Perspective on HAZ Toughness, in Welding Metallurgy of Structural Steels, J

Koo, Ed., The Metallurgical Society, 1987, p 255-262

• H Suzuki "Carbon Equivalent and Maximum Hardness," DOC IX-1279-83, International Institute of Welding, 1983

• Welding, Brazing, and Soldering, Vol 6, 9th ed., Metals Handbook, American Society for Metals, 1983

• Welding Handbook, Vol I and II, 7th ed., American Welding Society, 1983

Elevated-Temperature Properties of Ferritic Steels

Introduction

CARBON STEELS and low-alloy steels with ferrite-pearlite or ferrite-bainite microstructures are used extensively at elevated temperatures in fossil-fired power-generating plants, aircraft power plants, chemical-processing plants, and petroleum-processing plants Carbon steels are often used up to about 370 °C (700 °F) under continuous loading, but also have allowable stresses defined up to 540 °C (1000 °F) in Section VIII of the ASME Boiler and Pressure Vessel Code Carbon-molybdenum steels with 0.5% Mo are used up to 540 °C (1000 °F), while low-alloy with 0.5-1.0% Mo in combination with 0.5-9.0% Cr and sometimes other carbide formers (such as vanadium, tungsten, niobium, and titanium) are often used up to about 650 °C (1200 °F) For temperatures above 650 °C (1200 °F), austenitic alloys are generally used However, these general maximum-use temperature limits do not necessarily apply in specific applications with different design criteria Tables 1 and 2, for example, list maximum-use temperatures in two specific application areas with different design criteria

Table 1 Temperature limits of superheater tube materials covered in ASME Boiler Codes

Maximum-use temperature

Oxidation/graphitization criteria, metal surface (a)

Strength criteria, metal midsection Material

SA-106 carbon steel 400-500 750-930 425 795

Ferritic alloy steels

0.5Cr-0.5Mo 550 1020 510 950

1.2Cr-0.5Mo 565 1050 560 1040

2.25Cr-1Mo 580 1075 595 1105

Austenitic stainless steel

(a) In the fired section, tube surface temperatures are typically 20-30 °C (35-55 °F) higher than the tube midwall temperature In a typical U.S utility boiler, the maximum metal surface temperature is approximately 625 °C (1155 °F)

Table 2 Suggested maximum temperatures in petrochemical operations for continuous service based on creep or

ruptured data

Maximum temperature based on creep rate

Maximum temperature based on rupture Material

Type 304 stainless steel 595 1100 815 1500

This article covers some elevated-temperature properties of carbon steels and low-alloy steels with ferrite-pearlite and ferrite-bainite microstructures for use in boiler tubes, pressure vessels, and steam turbines In these applications, the selection of steels to be used at elevated temperatures generally involves compromise between the higher efficiencies obtained at higher operating temperatures and the cost of equipment, including materials, fabrication, replacement, and downtime costs The highly alloyed steels, which depend on an austenitic matrix for their high-temperature properties, generally have higher resistance to mechanical and chemical degradation at elevated temperatures than the low-alloy ferritic steels However, a higher alloy content generally means higher cost Therefore, carbon and low-alloy ferritic steels are extensively used in several forms (piping, pressure vessel plates, bolts, structural parts) in a variety of applications that involve exposure to elevated temperatures In addition, interest in ferritic steels has increased recently because their relatively lower thermal expansion coefficient and higher thermal conductivity make them more attractive than austenitic steels in applications where thermal cycling is present

To illustrate the tonnage requirements for carbon and low-alloy steels in industrial construction, 1360 Mg (1500 tons) of pressure tubing were required for the construction of a single 500 MW coal-fired generating plant The quantities of the various carbon and low-alloy steels used in the pressure tubing were as follows:

Steel type Tons % of total tonnage

Carbon 540 36

C- 1

2Mo

150 10

Elevated-Temperature Properties of Ferritic Steels

Carbon and Low-Alloy Steels for Elevated-Temperature Service

The numerous types of steels used in elevated-temperature applications include the following:

in the article "High-Strength Structural and High-Strength Low-Alloy Steels" in this Volume, may be effective substitutes for carbon steels in elevated-temperature applications Another category of ferritic steels for elevated-temperature service are manganese-molybdenum-nickel ferritic steels (ASTM A 302 and A 533), which are commonly used for pressure vessels in light-water reactors High-alloy steels, stainless steels, hot-work tool steels, and the iron-base superalloys are discussed in the Section "Specialty Steels and Heat-Resistant Alloys" in this Volume

Alloy Designations and Specifications

Carbon and low-alloy steels used for elevated-temperature service are usually identified by American Iron and Steel Institute (AISI) designations; aerospace material specification (AMS), American Society of Mechanical Engineers (ASME), or American Society for Testing and Materials (ASTM) specification number; nominal composition; or trade name These steels have also been assigned numbers in the Unified Numbering System In addition, there are Military and Federal specifications covering many of these steels

Steel products manufactured for use under the ASME Boiler and Pressure Vessel Code must comply with provisions of the appropriate ASME specification Each specification includes information on ranges and limits of composition, dimensions and tolerances, minimum mechanical properties, and other functional requirements The designations applied

to these products include the letters "SA," the number of the specification, and possibly other letters or numbers to distinguish among the various types, grades, and classes within a single specification Most ASTM specifications are identical to the ASME specification of the same number except that the ASTM designations begin with the letter "A." Some examples of ASME specifications for elevated-temperature steels, as well as their compositions and typical room-temperature mechanical properties, are given in Tables 3(1) and 3(2)

Table 3(1) Compositions of steels for elevated-temperature service

Composition, % ASME

specification

UNS designation

Nominal composition

Product form

carbon steel pipe

0.17(a) 0.90(a) 0.035(a) 0.045(a) 0.25

Cu (a)

SA-299 K02803 C-Mn-Si C-Mn-Si

steel PV plate

0.28(a)

0.90-1.40

0.30

0.20(a)

0.95-1.30

0.30

PV plate

0.25(a)

1.15-1.50

0.30

0.15-0.035(a) 0.040(a)

0.40-0.70

0.60

0.45-0.10

Cu (a)

SA-517F K11576

High-strength alloy steel

PV plate

0.20

0.10- 1.00

0.60- 0.35

0.15-0.035(a) 0.040(a)

0.40-0.65

1.00

0.70- 0.60

0.40- 0.006

0.002-B, 0.15- 0.050

Cu, 0.03- 0.08

V

SA-335 P12 K11562

1Cr- 1

2Mo

Seamless ferritic alloy steel pipe for high- temperature service

0.44- 0.44- 0.44-

SA-217WC6 J12072

1 1

41

Cr-2Mo

Alloy steel castings

MPa ksi MPa ksi

Minimum elongation in 50

Table 4 Compositions and mechanical properties of AISI steels for elevated-temperature service

Nominal composition, % AISI

designation

AMS

designations

Commercial designation

UNS designations

Typical applications

C Mn Si Cr Mo V

601 6304 K14675 Bolting and structural parts 0.46 0.60 0.26 1.00 0.50 0.30

602 6302, 6385, 6458 17-22 AS K23015 Bolting and structural parts 0.30 0.55 0.65 1.25 0.50 0.25

603 6303, 6436 17-22 AV K22770 Turbine rotors and aircraft

parts

0.27 0.75 0.65 1.25 0.50 0.85

610 6437, 6485 H11 mod T20811

K74015

Ultrahigh-strength components

0.40 0.30 0.90 5.00 1.30 0.50

Room-temperature tensile properties Temperature at which 70

MPa (10 ksi) will cause rupture in

Temperature to produce min creep

rate at 70 MPa (10 ksi)

Yield

strength

Tensile strength

603 1000 145 1100 160 17 52 650 1200 613 1135 565 1050

610 1480 215 1805 262 10 36 630 1170 595 1100 560 1040 540 1000

The AISI designation for steels intended for elevated-temperature service is a three-digit number beginning with a 6, such

as 601 The AISI designations are also included in Table 4

Fig 1 Effect of elevated-temperature exposure on the room-temperature tensile properties of normalized

0.17% C steel after exposure (without stress) to indicated temperature for 83,000 h

Fig 2 Effect of exposure to elevated temperature on stress-to-rupture of carbon steel Stress-to-rupture in

1000 and 10,000 h at the indicated temperature for specimens of normalized 0.17% C steel exposed to the test temperature (without stress) for 83,000 h and for similar specimens not exposed to elevated temperature prior

to testing

Creep-Resistant Low-Alloy Steels

Creep-resistant low-alloy steels usually contain 0.5 to 1.0% Mo for enhanced creep strength, along which chromium contents between 0.5 and 9% for improved corrosion resistance, rupture ductility, and resistance against graphitization Small additions of carbide formers such as vanadium, niobium, and titanium may also be added for precipitation strengthening and/or grain refinement The effects of alloy elements on transformation hardening and weldability are, of course, additional factors

The three general types of creep-resistant low-alloy steels are chromium-molybdenum steels, vanadium steels, and modified chromium-molybdenum steels Chromium-molybdenum steels are used primarily for tube, pipe, and pressure vessels, where the allowable stresses may permit creep deformation up to about 5% over the life of the component Typical creep strengths of various chromium-molybdenum steels are shown in Fig 3 Figure 3 also shows the creep strength of a chromium-molybdenum steel with vanadium additions Chromium-molybdenum-vanadium steels provide higher creep strengths and are used for high-temperature bolts, compressor wheels, or steam turbine rotors, where allowable stresses may require deformations less than 1% over the life of the component

chromium-molybdenum-Fig 3 General comparison of creep strengths of various creep-resistant low-alloy steels

Chromium-molybdenum steels are widely used in oil refineries, chemical industries, and electrical power generating stations for piping, heat exchangers, superheater tubes, and pressure vessels The main advantage of these steels is the improved creep strength from molybdenum and chromium additions and the enhanced corrosion resistance from chromium The creep strength of chromium-molybdenum steels is derived mainly from two sources: solid-solution strengthening of the matrix ferrite by carbon, molybdenum, and chromium: and precipitation hardening by carbides Creep strength generally, but not always, increases with higher amounts of molybdenum and chromium The effects of chromium and molybdenum on creep strength are complex (see "Effects of Composition" in this article) In Fig 3, for example, 2.25Cr-1Mo steel has a higher creep strength than 5Cr-0.5Mo steel

Chromium-molybdenum steels are available in several product forms (see Table 24 in the article "Classification and Designation of Carbon and Low-Alloy Steels" in this Volume) In actual applications, boiler tubes are used mostly in the annealed condition, whereas piping is used mostly in the normalized and tempered condition Bend sections used in piping, however, are closer to an annealed condition than to a normalized condition As a result of the cooling rates employed in these treatments, the microstructures of chromium-molybdenum steels may vary from ferrite-pearlite aggregates to ferrite-bainite aggregates Bainite microstructures have better creep resistance under high-stress, short-time conditions but degrade more rapidly at high temperatures than pearlitic structures As a result, ferrite-pearlite material has better intermediate-term, low-stress creep resistance Because both microstructures will eventually spheroidize, it is expected that over long service lives the two microstructures will converge to similar creep strengths

The 0.5Mo steel with 0.15% C is used for piping and superheater tubes operating at metal temperatures to 455 °C (850

°F) Above this temperature, spheroidization and graphitization may increase the possibility of failure in service Use of

carbon-molybdenum steel has been largely discontinued for the higher temperatures because of graphitization Chromium steels are highly resistant to graphitization and are therefore preferred for service above 455 °C (850 °F)

The 1.0Cr-0.5Mo steel is used for piping, cracking-still tubes, and boiler tubes for service temperatures to 510 or 540

°C (950 or 1000 °F) The similar 1.25Cr-0.5Mo steel is used up to 590 °C (1100 °F) and has comparable stress-rupture and creep properties as that of the 1.0Cr-0.5Mo alloy (Fig 4)

Fig 4 Creep strength (0.01% 1000 h) and rupture strength (100,000 h) of 1Cr-0.5Mo and 1.25Cr-0.5Mo steel

Source: Ref 1

The 2.25Cr-1.0Mo steel has better oxidation resistance and creep strength than the steels mentioned above The 2.25Cr-1Mo steel is a highly favored alloy for service up to 650 °C (1200 °F) without the presence of hydrogen or 480 °C (900 °F) in a hydrogen environment This steel, which has substantial documentation of its elevated-temperature properties (Ref 2, 3, 4, 5, and 6), is discussed in more detail in the section "Elevated-Temperature Behavior of 2.25Cr-1Mo Steel" in this article

The 5, 7, and 9% Cr steels are generally lower in stress rupture and creep strength that the lower-chromium steels because the strength at elevated temperatures typically drops off with an increase in chromium However, this may not always be the case, depending on the service temperature (Fig 5) and the exposure (Fig 6 and 7) Heat treatment is also

an important factor The main advantage of these steels is the improved oxidation resistance from the increased chromium content