Canale L.C.F., Mesquita R.A., Totten G.E. Failure Analysis of Heat Treated Steel Components

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dependent on austenitization time, temperature,

cooling history of the particular alloy, and the

composition of the alloy. The transformation

products from austenite decomposition change

from a mixture of ferrite and pearlite to bainite

or martensite with increasing cooling rate.

Isothermal transformation (time-temperature

transformation, or TTT) and continuous cooling

transformation (CCT) diagrams can be used to

predict steel microstructures and hardness after

heat treatment. These diagrams are a set of

curves drawn in a semilogarithmic coordinate

system with logarithmic time/temperature

coordinates that deﬁne, in the case of the TTT

diagram, for each level of temperature, the

beginning and end of the transformation of

austenite under isothermal conditions, and, in

the case of the CCT diagram, deﬁne each var-

iation in temperature as a function of time during

cooling, the temperature at which the austenite

begins and ends its transformation. Industrial

heat treatments consist of heating and soaking

at a suitable temperature, followed by cooling

at an appropriate rate in order to obtain a struc-

tural state closer to that equilibrium (annealing),

reduce the internal stresses without substantially

modifying the structure (stress relieving),

increase hardness by more or less complete

transformation of austenite to martensite and

possibly bainite (quench harden), or obtain a

uniform and ﬁne-grained structure with pearlite

(normalizing). Other heat treatments are applied

to ferrous products after quench hardening

(tempering) or solution treatment (aging) to

bring the properties to the required level. The

process conditions that shall be taken into

account usually include the furnace atmosphere

(for example, temperature and carbon potential),

heating rates, and quench conditions.

Thermomechanical Modeling

Primary stress is a normal or shear stress

developed by the imposed loading that is

necessary to satisfy the laws of equilibrium of

external and internal forces and moments. Sec-

ondary stress is a normal or shear stress devel-

oped by the constraint of adjacent parts or by

self-constraint of a structure. The basic char-

acteristic of the secondary stress is that it is self-

limiting. An example of secondary stress is a

general thermal stress.

Thermal stresses are related to temperature

by the thermal expansion coefﬁcient. The

deformation is proportional to the thermal

gradient through the thermal conductivity, l.

Thermal expansion has a strong inﬂuence on the

development of residual stress; whenever the

thermal expansion or contraction of a body is

prevented, thermal stresses appear. If large

enough, they cause yielding, fracture, or elastic

collapse (buckling).

The total induced strain is the result of the

sum of the strain produced by thermal (aE

DT=e

) and transformation (e

) strain due to

local transformation-induced volume expan-

sion. The total induced strain must be accom-

modated through either elastic (e

) or plastic

) strain, which sums to the total strain:

=  (e

)(Eq 11)

Thermal Shock. The ability of a material to

resist thermal shock, due to a sudden immersion

in a cold ambient, without cracking depends on

its thermal expansion coefﬁcient, a; the tensile

strength, s

, for metals; the Young’s modulus, E;

the thermal conductivity, l; and the heat-transfer

coefﬁcient, h. A temperature change of DT

applied to a constrained body or a sudden

change, DT, of the surface temperature of the

unconstrained component induces a thermal

strain.

Stress Concentration. The elastic strain and

stresses calculated previously are nominal

values that do not take into account local dis-

continuities such as holes, notches, or section

changes. Even on a structure where stress

intensity has been limited by yield criteria, there

may exist highly localized regions where

peak stresses are several times higher than

yield. Maximum local stresses on a structure can

be determined considering nominal stresses

multiplied by a stress-concentration factor and

can be estimated through a detailed stress ana-

lysis or by using approximate formulas that

account for the most common cases. Design

should be veriﬁed to conﬁrm whether there are

stress-concentration points that may activate

failure mechanisms due to brittle fracture, cor-

rosion, or fatigue. A sharp corner concentrates

and magniﬁes applied stresses, stresses that arise

in tool and die manufacturing (such as during

quenching), or stresses that occur during service.

In addition to promoting cracking during liquid

quenching, sharp corners promote buildup of

residual stresses that may not be fully relieved

by tempering and can therefore reduce service

life. The largest possible ﬁllet should be used

at all sharp corners. Air-quenching grades of

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steel are more tolerant of sharp corners than

liquid-quenching grades. Fatigue fractures pro-

pagate from the sharp edges.

Residual stresses can be deﬁned as those

stresses that remain in a material or body after

being manufactured and processed in the

absence of external forces or thermal gradients.

After the cooling processes have been com-

pleted, the residual-stress distribution between

the surface and core is obtained. If the surface

stresses exceed the hot yield strength of the

material, it plastically deforms, resulting in

thermally induced dimensional changes and

thermal and transformational stresses. In heat

treated parts, residual stresses may be classiﬁed

as those caused by a thermal gradient alone and a

thermal gradient in combination with a micro-

structural change (phase transformation).

Distortion is any change in the shape and

original dimensions of a ferrous product occur-

ring during heat treatment. For a given geometry

and heat ﬂow, the distortion is minimized by

selecting materials with large values of l/a.

Shape and volume changes during heating and

cooling can be attributed to three fundamental

causes:

 Residual stresses that can cause shape

change when they exceed material yield

strength

 Stresses caused by differential expansion

due to thermal gradients

 Volume changes due to transformational

phase change

Quench Cracking. If thermally induced

stress exceeds the local tensile strength of the

material, cracking results.

Thermochemical Model

Thermochemical treatments may be applied

to a ferrous product in the austenite state to

obtain a surface enrichment in carbon (carbur-

izing), which is in solid solution in the austenite.

Carburizing can be done in gas atmosphere,

in solid medium, or in a bath of molten salt.

Other thermochemical treatments can be applied

to produce surface enrichment in nitrogen

(nitriding); in nitrogen and carbon (nitro-

carburizing); in sulfur, carbon, and nitrogen

(sulﬁdizing); in silicon (siliconizing); in chro-

mium (chromizing); in boron (boriding); and in

aluminum (aluminizing). By error in the heat

treatment, the surface can be decarburized.

Table 3 shows guidelines to avoid heat treat-

ment failures during the design review (phase 2

design review); the recommendations come from

the examples presented in the previous sections.

Failure Aspects of Welded Components

Brittle Fracture. Low-temperature/low-

toughness fracture is sudden failure of a struc-

tural component that is usually initiated at a

crack or defect. This is an unusual occurrence,

because design stresses are normally sufﬁciently

low to prevent such an occurrence. However,

Table 3 Phase 2 design review to avoid heat treatment failures

Characteristic of the material

prior to heat treatment

operation

Always use the grade or composition of steel most suitable for the work that the part has to perform.

The carbon content of the steel is one of the determining factors for quenchant selection; plain carbon steels

with less than 0.35% C rarely crack on hardening.

Avoid structural and compositional material heterogeneity.

Avoid “dirty” steels (clustered, stringer-type inclusions, high sulfur).

Geometry: shape and

dimensions

Avoid asymmetric design.

Avoid points of stress concentration (sharp corners, blind holes, reentrant angles, single internal or external

keys, deep keyways, splines, holes, grooves, very coarse machining marks, deep scratches, and tool

marks). Provide adequate ﬁllet or radius at the base of gear teeth, splines, and serrations. Do not have

holes in direct line with the sharp angles of cutouts. Do not drill screw holes closer from edges.

Add extra holes, if possible, on heavy, unbalanced sections to allow for faster and more uniform cooling

when quenched.

Order stock large enough to allow for machining to remove decarburized surfaces and surface

imperfections, such as laps and seams, and to correct the distortion.

Heat treatment operation Verify design suitability for the type of furnace equipment available.

Avoid thermal shock.

Avoid decarburized surface.

Consider contact with furnace hearths or placement near walls.

Avoid heat treatment errors.

Control positioning in the furnace and during quenching.

Incorporate straightening operations as required.

Avoid creep of the part at elevated temperature under its own weight or as a result of ﬁxturing.

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some older equipment with thick walls, equip-

ment that may be subjected to low temperature

due to an upset, or equipment that may be

modiﬁed could be susceptible to varying degrees

of embrittlement.

Brittle fracture is associated with very little or

no plastic deformation and with a cleavage

fracture or intergranular surface, unlike ductile

fracture, which is associated with a ﬁbrous sur-

face. A material may fail in an unstable and

catastrophic brittle manner under stresses even

lower than the allowable design stresses used to

avoid ductile failures. This may occur with a

combination of material properties and applied

stress levels.

The material property that controls brittle

fracture strength is toughness. Other factors that

have an effect on brittle fracture are material

thickness; local stress level, including nominal

stresses, residual stresses, and stress-concentra-

tion factors; temperature; and loading rate.

Carbon and low-alloy steels undergo a transition

from ductile failure mode to brittle failure

mode at low temperatures. Resistance to crack

propagation is measured through fractomechan-

ical tests. Crack propagation will occur when the

stress intensity at the crack tip, K, reaches a

critical value, K

(MPam

1/2

). Brittle materials

are those that remain elastic until breaking

(break occurs before yield).

According to API 581 (Ref 21) low-

temperature/low-toughness fracture of steel is

affected by:

 Applied loads: Fracture is less likely at low

applied loads.

 Materials speciﬁcation: Some materials are

manufactured to exhibit good fracture

properties or toughness properties. Materials

are often qualiﬁed for use by performing an

impact test that measures the energy needed

to break a notched specimen. Fine-grained

structures, such as tempered martensite, with

low impurity content are associated with

a high degree of toughness. Other micro-

structural elements, such as precipitates,

second-phase particles, dislocations, and

solutes in a solid solution, contribute to

increased yield strength but reduce tough-

ness.

 Temperature: Many materials (especially

ferritic steels) become brittle at a tempera-

ture called the transition temperature. Brittle

fracture is typically not a concern above

300



C (570



F).

 Residual stresses and postweld heat treat-

ment

 Thickness

Temper embrittlement is one of the main

causes of toughness degradation in ferritic steels

during high-temperature service. This degrada-

tion may lead to component failure during ser-

vice. The problem arises when some types of

steels are exposed to temperatures between 345

and 565



C (650 and 1050



F). Typically, 2

Cr-

Mo steels with a bainitic structure are

the most susceptible to this phenomenon.

Temper embrittlement can also occur in

C-

Mo, 1Cr-

Mo, 1

Cr-

Mo, 3Cr-1Mo, and

5Cr-

Mo steels. Conversely, 9Cr-1Mo steels

are less susceptible. Welded joints (weld metal

and heat-affected zone) are the most susceptible

zones.

In all cases, the solution to the problem lies in

alloy purity. Exposures within the critical tem-

perature range may occur during temper or

postweld heat treatments or during service, and

these conditions should be avoided. However,

many components operate within the critical

temperature range.

The segregation of residual elements anti-

mony, arsenic, phosphorus, and tin toward

austenitic grain boundaries is the main cause

of temper embrittlement. Also, manganese and

silicon play an important role in this segre-

gation, and their content should be limited.

Both residual and alloy elements can segre-

gate, but the former can be concentrated up to

300 times their average value in the material.

Segregation only occurs in ferrite within a 315

to 540



C (600 to 1005



F) temperature range

but never occurs during austenitization. In

addition to segregation in grain boundaries, a

ﬁne precipitation can occur within the grains,

resulting in a strength increase (Mo

C pre-

cipitation).

The phenomenon associated with changes in

grain boundaries causes intergranular brittle

fracture. In general, ductility and rupture

strength are not affected; nevertheless, both

can be reduced under severe conditions.

Toughness is affected by up to a 100



(212



F) shift toward the right of the ductile-

brittle transition curve, as evidenced from

impact testing.

When a material is exposed to a 370 to 565



(700 to 1050



F) temperature range, property

degradation may become irreversible; in this

case, temper embrittlement and creep operate

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simultaneously. During equipment operation

in a hydrogen service environment, hydrogen

may diffuse into the metal. During cooling

from the operating temperature (shutdown), the

material becomes oversaturated with hydrogen.

The combination of thermal stresses and

hydrogen oversaturation may lead to hydrogen-

induced cracking. If the material toughness has

been reduced considerably due to temper

embrittlement, the risk of a catastrophic crack-

ing is high.

Material degradation due to temper embrit-

tlement may result not only in brittle fractures

with catastrophic consequences but also in a

reduction of the equipment useful life and in a

decrease in the equipment reliability and efﬁ-

ciency, since it may be necessary to operate at

lower temperatures to avoid temper embrittle-

ment or to depressurize to avoid stresses when

the equipment is cold.

Temper embrittlement is reversible. Heat

treatment for a short period of time at tempera-

tures above 565



C (1050



F), followed by

quick cooling, can restore the initial properties.

However, if material thickness is high and

cooling rates required by precipitation kinetics

are not achieved, embrittlement may reoccur.

Moreover, the component may crack during

heat treatment due to the effect of thermal

stresses.

Material thickness over 25 mm (0.98 in.) is

more susceptible to brittle fracture due to temper

embrittlement. Postweld heat treatments mini-

mize the susceptibility to brittle fracture through

this mechanism.

Aging Tendency. The tendencies for aging

after cold working increase hardness and tensile

strength, with a simultaneous reduction in duc-

tility. Therefore, the risk of embrittlement due to

welding in cold-worked areas increases. The

killed steels show enhanced resistance to aging

in normalized conditions. When the degree of

deformation is high, the material should be

thermally treated (normalized or stress relieved)

before the deformed zone is welded. The risk of

embrittlement due to welding in cold-worked

areas to be welded, without special require-

ments, is allowable under the following con-

ditions. The relationship between the inner

bending radius (r) and the plate thickness (t)

shall be:

 r/ti1.0 and tj4.0 mm (0.16 in.)

 r/ti1.5 and tj8.0 mm (0.32 in.)

 r/ti2.0 and tj12.0 mm (0.47 in.)

 r/ti3.0 and tj 24.0 mm (0.94 in.)

 r/ti10.0 and all thicknesses

These conditions correspond to an elongation,

e=1/(2r/t+1), in the cold-worked areas of 33,

25, 20, 14, and 5%, respectively.

Hardening Tendency. The maximum hard-

ness in the heat-affected zone (HAZ) depends

on the chemical composition (carbon equiva-

lent) and the cooling rate during welding (t

8/5

Carbon most affects hardenability, and its

effect and that of other elements have been

included in carbon equivalent formulas. The

International Institute of Welding carbon equi-

valent formula, recommended for steel with

more than 0.18% C, is:

IIW

=C+

Cr+Mo+V

Cu+Ni

(in wt%)(Eq12)

When carbon is 50.18%, it is generally re-

commended that the percentage of cementite

) formula be used:

=C+

+5B (in wt%)(Eq13)

Hardenability of steel is not necessarily an

indicator of HAZ hardness. It is important to

control the maximum hardness in the HAZ in

order to avoid two main problems: cold

cracking (hydrogen-assisted cold cracking)

during welding fabrication, in which the limit

of maximum hardness is usually 350 HV; and

in-service cracking in hydrogen environments,

where the maximum hardness shall be less

than 220 HV. Maximum HAZ hardness can

be accurately calculated as a function of

chemical composition and the cooling time

from 800 to 500



C (1470 to 930



F) (t

8/5

)

(Ref 22).

The HAZ maximum hardness for low-alloy

steels can be calculated using the following

expressions by Du

ren (Ref 22). For 100% mar-

tensite (alloying elements in wt%):

=802 C+305 (Eq 14)

For 100% bainite:

350  C+



+101 (Eq15)

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For 05%martensite5100:

=2019  C  (170:5  log t

8=5

)

+0:3 



+66 (170:8  log t

8=5

)(Eq16)

where the hardness of martensite is HV

, bai-

nite is HV

, the variable amount of martensite is

, and t

8/5

is the cooling time between 800

and 500



C (1470 and 930



F).

Cold cracking can occur after the welding

process if four factors are present: local hydro-

gen concentration, susceptible weld metal or

HAZ, local metal hardness, and a high level of

residual stress remaining after welding, which

may cause cracking at temperatures less than

100



C (212



F).

The causes of cold cracking are related to

many factors, including initial weld metal

hydrogen content, residual hydrogen content at

100



C, steel carbon equivalent, yield stress of

steel or weld metal, heat input, preheat tem-

perature, material thickness, joint restraint

intensity, notch concentration factor, welding

process thermal efﬁciency, and others. Cold

cracking is a diffusion-controlled phenomenon

that requires days or weeks at room temperature

to develop cracks.

Hydrogen-induced cracking in the HAZ can

be parallel to the fusion boundary adjacent to a

ﬁllet weld or in the form of toe cracks. Weld

metals are by no means immune when the steels

possess high yield strength. Thus, cold cracking

induces surface-connected cracking or subsur-

face cracking, which may provide initiation

points for further cracking by brittle fracture or

fatigue.

Segregation Tendency. Element (phos-

phorus, sulfur, carbon, etc.) segregation impairs

weldability. Chemical heterogeneity can con-

tribute to localized increases of hardenability.

Thus, a normal chemical composition of the heat

of the steel may exhibit hardness higher than the

maximum allowable hardness in certain parts of

the HAZ, despite a normal chemical composi-

tion. From the point of view of segregation

behavior, semikilled and killed steels are better

than rimmed steels. If segregation zones are

involved—as in butt welding—care should be

taken to limit penetration and hence minimize

weld metal dilution. In addition, suitable ﬁller

metal and low-hydrogen basic electrodes should

be used. The annealing treatment relieves

internal stresses, restoring ductility and tough-

ness, reﬁning grains, reducing gaseous content

in the steel, and improving homogenization of

alloying elements.

Stress-corrosion cracking (SCC) of auste-

nitic stainless steels can be caused by chlor-

ides and polythionic acids, by hydrosulfuric acid

on carbon and low-alloy steels, and by caustic

corrosion on carbon steels. Stress-corrosion

cracking may arise when a susceptible material

is simultaneously combined with certain levels

of tensile stresses and a critical environment

within a speciﬁc temperature range. Tensile

residual stresses, resulting from manufacturing

processes such as welds, contribute to cause this

type of damage. Also, cold plastic deformation

causes hazardous residual stresses. Tensile

stresses must be reduced by controlling manu-

facturing processes and design. Stress relieving

postweld heat treatments can be used to mini-

mize susceptibility to SCC in austenitic stainless

steel, attempting to avoid sensitization by using

stabilized or low-carbon grades.

Heat Treatment Procedures Applied

to Welded Components

Carbon and Low-Alloy Steels

Postweld heating (soaking) is an option for

carbon-manganese and low-alloy steels. Mat-

erial of thick-walled low-alloy steel pressure

vessels is susceptible to hydrogen-induced

cracking, and there are many difﬁculties in

detecting small cracks in the HAZ of heavy-

section steel by conventional nondestructive

examination methods. Therefore, it is very

important to reduce to a minimum the risk of

hydrogen-induced cold cracking of weldment of

thick plate under the heavily restricted condition

during the entire welding process. One possible

solution is to apply both preheating and post-

heating (Ref 23). There is a temperature range

where weld hydrogen-induced cracking may

occur and a temperature above the upper limit

where no delay cracking occurs, although the

hydrogen content or the restraint intensity is

high. Therefore, if postheating is carried out

above this critical temperature, the discharge of

hydrogen is allowed; then, the material can

reach room temperature with a minimum risk of

hydrogen-induced cracking.

The purpose of soaking is to allow hydrogen

diffusion to avoid critical values in the

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weldments. Welding must be performed, main-

taining preheating and interpass temperatures

that depend on the material, process, thickness,

type of joints, and heat input. The postheating

must be carried out immediately after welding,

not allowing the temperature to be lower than

120



C (250



F). The temperature for post-

heating generally is 300



C (570



F) for 4 to

6 h, depending on the steel type. Then, the

weldment is cooled to room temperature. Gen-

erally, the same devices recommended for pre-

heating are used. This treatment does not

produce either stress relief or microstructural

changes.

Postweld heat treatment (stress relieving)

may be used for carbon-manganese and low-

alloy steels. Postweld heat treatment (PWHT) is

a uniform heating of a weldment at a tempera-

ture below the critical range to relieve the major

part of the residual stresses, followed by uniform

cooling in still air. The PWHT is carried out to

fabricate vessels to increase fracture toughness

and minimize the levels of residual stress, which

confers resistance to brittle fractures, tempers

material structure, and removes the possibility

of SCC. The accrued beneﬁts depend on the

material under consideration.

It is usually required by codes and customers

that manufacturers should heat treat all welds on

a thick-wall pressure vessel, including any repair

welds. The PWHT is conducted at a temperature

and for the period speciﬁed in the applicable

fabrication codes. There are two possibilities:

one is to stress relieve the completed vessel in

the furnace as a whole; the other is to stress

relieve the subassemblies separately in the fur-

nace and then heat treat the ﬁnal circumferential

seam locally. Techniques for PWHT are some-

what similar to those thermal methods in pre-

heating and soaking. When local treatment is

carried out, care must be taken to ensure full

through-thickness heating, and similarly, the

temperature gradient is such that the length

of material on either side of the weld at a

temperature exceeding half the treatment tem-

perature is at least 2.5(rt)

1/2

, where r is the bore

radius, and t is the material thickness. Problems

can arise where butt welds attaching pipes to

nozzles positioned close to nozzle/vessel welds

require separate heat treatment. In such cases,

it may be necessary to apply individual stress

analysis to verify the proposed conditions

(Ref 23).

For both local and furnace treatment techni-

ques, it is very important to exercise control over

heating rates (particularly at temperatures below

300



C, or 570



F, with complex components),

time at soak temperature, soak temperature, and

cooling rate to avoid such undesirable events as

ﬂame impingement, distortion, overheating, air

quenching, and reheat cracking. Very large-

diameter pressure vessels that have transport

difﬁculties may be erected on site and PWHTed

from inside by gas burners, using the vessel as its

own furnace (Ref 23).

It is well known that the mechanical properties

of material are degraded due to stress relieving in

some materials (Ref 23). Generally, in low-alloy

pressure vessel steels, the yield strength, tensile

strength, and toughness diminish, and elongation

and reduction of area increase as the temper

parameter (TP) is increased:

TP=T (log t+20) · 10

3

(Eq 17)

where T is temperature (K), and t is time (h). The

PWHT is not considered to be a signiﬁcant

variable for carbon and carbon-manganese steels

up to 50 mm (2 in.) (Ref 23).

Postweld Heat Treatment of Stainless Steels

Postweld heat treatment is used for stainless

steels (Ref 24–26). Welding stainless steels in

thick sections, when the thickness exceeds

approximately 20 mm (0.8 in.), is a complex

operation. In addition to codes or engineering

speciﬁcations, which may impose deﬁnitive

procedures, it is important to have sufﬁcient

metallurgical background to understand what

may happen during welding and the subsequent

heat treatment operation.

The stresses induced by welding often need

to be eliminated if dimensional stability of the

construction is to be guaranteed and if resistance

to SCC is mandatory. The properties of welded

joints of stainless steels in thick sections that

may be modiﬁed during a PWHT are principally

the corrosion resistance and the mechani-

cal properties. The possibility of distortion

occurring during heat treatment must also be

considered.

For the purpose of analyzing PWHT in

stainless steels, the materials considered were

divided into four groups.

Chromium Steels. For both ferritic and

martensitic types, some codes recommend pre-

heating in the temperature range from 150 to

400



C (300 to 750



F) to avoid problems with

hydrogen-induced cracking in welding these

steels.

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Martensitic stainless steels are normally

PWHTed between 600 and 800



C (1110 and

1470



F), whereas ferritic steels are sometimes

heated between 730 and 800



C (1345 and

1470



F), with rapid cooling in order to avoid

embrittlement. Treatments above 900



(1650



F) in these materials are usually intended

as homogenizing treatments in order to achieve

better properties after the tempering treatment

that follows. The mechanical properties are

usually improved after this double heat treat-

ment. Low-temperature stress relieving should

not be applied to the straight chromium stainless

steels, since it may markedly affect the ductility

and toughness. The phenomenon is known as

475



C (885



F) embrittlement, and it is due to

the coherent precipitation of chromium-rich

ferrite, known as alpha prime, within the mis-

cibility gap of the iron-chromium system. This

precipitation leads to a slow increase in hardness

accompanied by a corresponding loss of tough-

ness. The alpha-prime phase also decreases the

corrosion resistance. This type of structural

change can be reversed by an annealing treat-

ment at approximately 600



C (1110



F). In

17% Cr steels containing nickel and molybde-

num, the toughness is increased by tempering at

630 to 650



C (1165 to 1200



F), below the

temperature where austenite or ferrite is formed.

In straight 17% Cr steels, the precipitation of

sigma phase can occur between 550 and 800



(1020 and 1470



F), and it is accompanied by a

loss of ductility. The sigma phase is formed only

after a very long time and may be eliminated by

heat treating above 800



C (1470



F).

Soft martensitic stainless steels have re-

sulted in an increasingly worldwide use in

petrochemical and chemical plants or industries,

gas turbine engines, turbine blades, compressors

and discs, and in a variety of aircraft structural

and engine applications (Ref 27). They have

high proof strength and high toughness even in

very low temperatures or thick cross sections

(Ref 28, 29). If the 12% Cr stainless steels are

used as high-strength structural steels, they must

be weldable, formable, and have good impact

toughness (Ref 30). Hence, in soft martensitic

stainless steels, the carbon content is kept below

mass 0.1% to improve weldability by promoting

a structure with fewer tendencies for cold

cracking, better corrosion resistance, and better

toughness. Because of the lower carbon, the

addition of 4 to 6% Ni (the most powerful

austenite former after carbon and nitrogen)

is required to avoid delta ferrite, which is

deleterious to impact toughness. For enhanced

corrosion, temper embrittlement, and tempering

resistance, 0.5 to 2% Mo is added, depending on

the intended use. In order to develop the max-

imum strength and toughness, the steel must be

mostly martensitic after cooling, with limited

delta ferrite. The martensite must be tempered to

obtain good toughness, ductility, and stress-

corrosion resistance. In the as-welded condition,

the microstructure consists of low-carbon mar-

tensite, some presence of delta ferrite, and

retained austenite in agreement with the nickel

content of the alloy. Postweld heat treatments

are necessary to satisfy the service mechanical

property requirements (Ref 31). If high impact

values are required, PWHTs such as solution

annealing plus tempering or double tempering

are necessary (Ref 32, 33). The aim of solution

annealing is the homogenization of the micro-

structure by dissolution of the delta ferrite,

which is a nonequilibrium solidiﬁcation product.

The delta ferrite is harmful since it increases the

ductile-brittle transition temperature. On the

other hand, intercritical tempering at 600



(1110



F) or double tempering (hypercritical

plus intercritical) at 670+600



C (1240+

1110



F) produces tempered martensite with

ﬁnely dispersed austenite that is stable and not

transformable during cooling (Ref 33). It is

known that this austenite, which can be observed

only by scanning electron microscopy, increases

toughness sharply, although it slightly reduces

the strength. It has been argued that when

retained austenite is present near a propagating

crack, the concentrated strain at the crack tip

induces transformation into martensite. This

mechanically induced transformation would

absorb energy and thus increase the toughness.

The associated volumetric expansion of the

martensitic transformation would tend to close

the crack and relieve stresses at its tip. The

latter mechanism absorbs strain energy during

fracture and therefore limits crack extension

(Ref 34).

In many codes for austenitic chromium-

nickel stainless steels with low ferrite content

and fully austenitic alloys, no PWHT is pre-

scribed. In the case of austenitic chromium-

nickel steels, low-temperature treatment (400 to

525



C, or 750 to 975



F) will help to achieve

dimensional stability of the construction by

reducing peak stresses, although this treatment is

not frequently used. Treatment in the tempera-

ture range from 550 to 1170



C (1020 to

2140



F) is a true stress-relieving treatment. The

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highest part of the range (960 to 1170



C, or

1760 to 2140



F) involves solution treatment

and achieves maximum relief of stress; it may be

performed to dissolve most of the carbides and

sigma phase or delta ferrite. Treatments in the

range from 1040 to 1170



C (1905 to 2140



F),

followed by water quenching after annealing,

are applied to prevent intercrystalline corrosion

and SCC. When PWHT is recommended, it only

applies to heavy thickness, and a temperature

between 900 and 1000



C (1650 and 1830



F) is

chosen, followed by water quenching or air

cooling, depending on the thickness of the

component. Although it may seem difﬁcult to

imagine such treatments being applied to heavy

components, there are some industrial examples.

The stress relieving of austenitic steels is not

usually applied except for very thick sections.

An exception is in the case of cladding, where

the stress-relieving temperature is chosen with

respect to the base material, and it is frequently

in the range of 540 to 700



C (1005 to 1290



F).

Austenitic stainless steels have a higher coefﬁ-

cient of thermal expansion and lower thermal

conductivity than ordinary ferritic steels, so

greater distortion of welded components must be

expected.

Unstabilized Austenitic Stainless Steels

(UNS S30400, S31600, S30403, and

S31603). These grades normally possess

excellent weldability, provided they are welded

with ﬁller metals that yield an austenitic-ferritic

weld metal to avoid hot cracking during welding

(5 to 15 ferrite number). It is necessary to follow

certain procedures in order to achieve sufﬁcient

corrosion resistance, cracking resistance, and

toughness. It is well known that austenitic

stainless steels are always subjected by the steel

manufacturer to a solution-annealing treatment,

normally in the range of 1050 to 1100



C (1920

to 2010



F). In the course of this heat treat-

ment, carbide M

, sigma phase, and delta

ferrite are completely dissolved, and the anneal-

ing process produces a homogeneous, fully

austenitic structure. With a subsequent quench-

ing treatment, this state is maintained up to

room temperature. If possible, the PWHT of

welded components should be avoided, with the

exception of a solution-annealing treatment.

However, if heat treatment cannot be avoided,

special attention must be paid to the inﬂuence of

carbide and phase precipitations on the corro-

sion-resistance and toughness properties of the

weld. The precipitation of sigma phase is the

most important of all precipitation phenomena,

apart from M

precipitation, particularly

regarding mechanical properties. Due to the

highest chromium content of delta ferrite, the

weld metal containing delta ferrite is often more

precipitation prone than the base metal of similar

composition (because the solution-annealing

treatment normally applied to the base metal

dissolves any delta ferrite).

Stabilized Austenitic Stainless Steels. This

group of steels contains grades that are alloyed

with titanium or niobium in order to improve

their intergranular corrosion resistance. Tita-

nium- and niobium-stabilized steels can be

welded using niobium-stabilized ﬁller metals

with delta ferrite contents in the range of 7 to 15

ferrite number. Whenever possible, heat treat-

ment after welding, for example, stress-relieving

treatment, should be avoided. If it cannot be

avoided, it is important to use special ﬁller

metals with less delta ferrite content to avoid

sigma precipitation and detrimental effects on

intergranular corrosion resistance. Stabilized

steels are somewhat more susceptible to sigma-

phase precipitation or to knife-line attack.

Fully Austenitic Stainless Steels. This

group of stainless steels has a stable austenitic

structure that must normally be welded with

fully austenitic ﬁller metals that do not produce

any ferrite in the weld deposit. In the event of a

PWHT, the intergranular corrosion attack range

is relatively strongly inﬂuenced by the chro-

mium and nitrogen content. The presence of a

small amount of sigma-phase or chi-phase pre-

cipitation is sufﬁcient to give a marked drop

in the pitting resistance. A 475



C (890



embrittlement does not occur in fully austenitic

weld metal, because this only occurs in the

presence of a ferritic structure.

Duplex austenitic-ferritic chromium-

nickel stainless steels having delta content in

the range of 30 to 60% are considered. An alloy

containing 25% Cr and 5% Ni, which, in

metallurgical terms, is very close to the widely

used duplex stainless steel 22Cr-3Mo-5Ni,

solidiﬁes completely to delta ferrite from melt-

ing. During further cooling, d-c transformation

starts at approximately 1200



C (2190



F) with

the precipitation of predominantly nodular aus-

tenite at the ferrite grain boundary. During fur-

ther cooling to room temperature, there is only a

partial transformation to austenite. The structure

now contains some 60% primary delta ferrite

and approximately 40% secondary precipitated

austenite. If such an alloy is now subjected to a

solution-annealing treatment, this ratio can be

Component Design / 39

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shifted to slightly higher austenite contents.

Preheating of duplex steels is not normally

required. With thick material, a preheating

between 100 and 150



C (212 and 300



F) may

be advantageous. In order to obtain high ducti-

lity in the welded joint, a solution-annealing

treatment, followed by water cooling of the

completed welded component, is not normally

necessary. If it is required, however, the tem-

perature must be set according to manufacturer’s

speciﬁcations. With a solution-annealing treat-

ment in the range of 1020 to 1100



C (1870 to

2010



F), very close to a metallurgical equili-

brium, it is possible to reverse any harmful

structures in the HAZ that may have occurred

during welding. It should be noted that this may

lead to severe distortion.

Heat Treatment of Austenitic-Ferritic

Dissimilar Joints. The PWHT of austenitic-

ferritic dissimilar joints or weldments should be

avoided whenever possible. However, heat

treatments such as annealing or stress relieving

may at times be unavoidable or even mandatory.

They should always be adapted to suit the

requirements of the low-alloy steel section of

the joint. Often, the steel in question is a low-

alloy creep-resistant steel type used in boiler

and pressure vessels that demand a PWHT to

suit each particular steel grade. In austenitic-

ferritic dissimilar joints, such heat treatments

may lead to the occurrence of the following

phenomena:

 Carbon enrichment in the weld metal due to

the diffusion of carbon from the low-alloy

steel into the austenitic weld metal

 As a consequence thereof, carbon depletion

in the HAZ of the low-alloy steel

 Coarse grain formation in the HAZ of the

low-alloy steel due to recrystallization pro-

cesses

 Embrittlement of the austenitic weld metal

due to precipitation of brittle phases, for

example, sigma phase

All these processes are time and temperature

dependent. Considering that the normal anneal-

ing time is in the range of 2 to 10 h, eventual

damage will normally occur, depending on the

material combination, at temperatures above

600



C (1110



F). Nickel-base weld metals,

due to their high nickel content outside the range

of sigma-phase precipitation, show no signs

of embrittlement during heat treatments. The

use of nickel-base ﬁller metals is recommended

if austenitic-ferritic dissimilar joints are subject

to PWHT at temperatures above 600



(1110



F). Nickel alloy ﬁller metals, abundant

and versatile, are frequently used for piping and

pressure vessel applications in reﬁneries, che-

mical plants, and power plants (Ref 35).

The Risk-Based Approach

and Heat Treatments

Risk analysis is a powerful tool to rationalize

the decision-making process and is applicable to

heat treated steel components. Increasing inci-

dents led to regulatory action. Some regulations

require that recognized and generally accepted

good engineering practice (RAGAGEP) must be

followed. This also created the desirable objec-

tive for industry to document what RAGAGEP

was.

The concepts of risk-based design, inspection,

and maintenance have been developed and are

being implemented (Fig. 42). Risk is the com-

bination of the probability (or frequency of

occurrence) and consequence (or severity) of a

hazard. Its scope is limited to a speciﬁc envir-

onment during a certain period of time. The

intent of risk-based initiatives is to use ﬁnite

resources and allocate these resources in a

manner that achieves the greatest overall

reduction in risk. Flaws that exceed the limits

permitted by codes may be found. After a defect

that is not acceptable is found via a ﬂaw eva-

luation/ﬁtness-for-service analysis, a repair or

replacement is required for continued operation.

Various types of damage may occur

throughout the different heat treatment stages.

These types of damage may lead to failures

in the product during its useful life. The term

failure refers to the inability of part of the

component or the entire component to perform

the functions it was designed for, leading to an

unproﬁtable and technically useless product.

Failure modes analysis is a procedure in

which each potential failure mode is analyzed to

determine its effects and the criticality of these

effects on the system and to rank each potential

failure according to its severity. The criticality

analysis involves the use of risk analysis tech-

niques based on the assessment of the likelihood

of failure and its potential consequences.

Among the most widely used tools in the

analysis of design is failure mode and effects

analysis (FMEA), which is accepted by the

Occupational Safety and Health Administration

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(OSHA). These methodologies are based on

reliability and safety.

By using the FMEA methodology in the

analysis of a new design, it is possible to identify

single-point failures (a single-point failure refers

to an individual failure that may cause the entire

system to collapse) and redesign the product

to avoid them, thus eliminating them completely

or achieving a more robust redesign that is less

sensitive to failures. The FMEA has evolved

from an ad hoc technique, dependent on a

designer’s experience, to a formal and accepted

analysis technique. Failure modes can be

eliminated by removing their causes or at least

having their probabilities of failure reduced to

acceptable levels.

There are several methodologies to assist in

logical thinking to resolve undesirable events.

These tools include, among others, reliability-

centered maintenance, risk-based inspection,

FMEA, modiﬁed FMEA, and root-cause

analysis.

Working into an integrated system with risk-

based efforts, fractomechanical-based structural

integrity approaches, and failure analysis will

allow the development of new tools to assist

the designers and manufacturers in minimizing

failures related to heat treating operations.

REFERENCES

1. L. Campos Franeschini Canale, G. Totten,

and D. Pye, Heat-Treating Process Design,

Handbook of Metallurgical Process Design,

G. Totten, K. Funatani, and L. Xie, Ed.,

Marcel Dekker Inc., New York, 2004

2. M. Solari, Risk Based Design, Chapter 2,

Handbook of Mechanical Alloy Design, G.

Totten, K. Funatani, and L. Xie, Ed., Marcel

Dekker Inc., New York, 2003

3. G. Totten, M. Narazaki, R.R. Blackwood,

and L.M. Jarvis, Failures Related to Heat

Treating Operations, Failure Analysis and

Prevention, Vol 11, ASM Handbook, ASM

International, 2002, p 192–223

4. M. Jacob and G.A. Hawkins, Elements of

Heat Transf er, John Wiley & Sons, Inc.,

1957

Risk-Based

Design

Risk-Based

Inspection

Fitness-

For-Service

Failure

Analysis

Mechanical Integrity

Initiatives

Risk Based Initiatives

Risk Assessment,

HAZOP, FMEA, WI, FTA,

RCA, RCM, RBI,

Fitness-For-Service,

Life Extension

Material Science

FEA, Welding, etc.

Economic

Factor

Human

Factor

Fig. 42 Strategies to minimize risk throughout the cycle life based on risk and mechanical integrity initiatives

Component Design / 41

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