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

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Fig. 22

(a) Surface of an AISI A4 primer cup plate showing spalling at one of the 3.2 mm diameter holes made by electrical discharge

machining (EDM) Original magniﬁcation: 2.5· . (b) Microstructures associated with the spalled hole in (a) caused by

improper EDM technique. Source: Ref 16

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to 0.040 in.). Determination of case depth by

visual examination on a microspecimen was

0.89 mm (0.035 in.).

The principal causes of failure were inade-

quate heat treatment of the case and a design that

incorporated a raised central section of the inner

diameter, which acted as a stress raiser. Rough

machining of the inner diameter aggravated the

situation. The case with the cementite grain-

boundary network had not been heated to a high

enough temperature or long enough to take the

cementite into solution in the austenite. It was

suspected that after slow cooling from the car-

burizing temperature, pins were heated slightly

above the Ae

prior to quenching and then given

a low temper. The case was reﬁned, and the core

was unreﬁned.

Thus, poor heat treatment, resulting in a brittle

grain-boundary network of cementite, and a

design that formed locations of stress con-

centration in the inner diameter were the most

probable causes of failure.

The pins should be carburized to a double heat

treatment to reﬁne both case and core and to

eliminate the brittle grain-boundary network of

cementite. The pin design should be changed to

eliminate the central raised section of the inner

diameter to avoid the ﬁllets acting as stress rais-

ers. The machining of the inner diameter should

be improved to avoid a rough surface. The depth

of the carburized case should be reduced to

approximately 0.38 mm (0.015 in.) to increase

pin toughness.

Example 12 features cracking of an alloy

steel bolt. A heat treated, cadmium-plated AISI

8740 steel bolt broke through the head-to-

shank ﬁllet while being handled during assem-

bly (Ref 19). Dimensions of the alloy steel bolt

(MSD 21250-10070) were 15.9 mm (0.625 in.)

diameter, 111 mm (4.375 in.) grip length, and

134.5 mm (5.294 in.) overall length. It was heat

treated to a tensile strength of 1240 to 1380 MPa

(180 to 200 ksi) and a hardness of 39 to 43 HRC

and then cadmium plated per QQ-P-416 type II

class 2 (23 h bake).

The bolt fractured through the head-to-shank

ﬁllet, a type of failure usually traceable to a

poorly controlled manufacturing process, such

as heat treating (quench cracking) or chemical

plating (hydrogen embrittlement). In this in-

stance, delayed cracking caused by hydrogen

embrittlement was initially suspected, because

the bolt reportedly had passed a magnetic par-

ticle inspection.

The fracture surface (Fig. 32), has two distinct

zones. Zone 1 was covered with a thick layer of

baked-on scale. The scale was removed and the

area examined using an SEM. The fracture

topography shown in Fig. 33 is a combination of

tearing and intergranular hairline cracks, fea-

tures often associated with both hydrogen

embrittlement and quench cracking. However,

the heavy adherent nature of the scale suggests

that it formed on the crack surface at high tem-

perature, that is, during heat treating but before

conventional quenching and subsequent cad-

mium plating. Zone 2 is characterized by

equiaxed dimples, a common feature of ductile

tension overload (ﬁnal stage of fracture). The

bolt head was cut through its centerline, and

the crack cross section was metallographically

examined in a zone 1 area to further explore

whether failure was caused by hydrogen em-

brittlement (step cracking, no decarburization)

or quench cracking (temper scale). The crack

had both heavy decarburization and temper scale

(Fig. 34), ruling out hydrogen embrittlement

cracking. However, because quench cracking

(a) (b) (c)

Fig. 23

Plastic mold die made from AISI S7 tool steel that was found to be cracked before use. A crack followed the lower recessed

contour of the large gear teeth and had an average depth of 1.6 mm. Smaller cracks were also observed on the ﬂat surfaces.

Source: Ref 16

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occurs at relatively low temperatures, de-

carburization of the surface could only have

occurred if the crack was present prior to heat

treating. When asked to clarify the situation, the

bolt manufacturer admitted that the part had

been “quenched in water from high temperatures

to verify dimension integrity and was returned to

the production lot, instead of being scrapped.” It

was assumed that this uncontrolled quench

between hot heading and heat treating caused the

bolt to crack. Decarburization and scaling

occurred during subsequent heat treating of the

cracked part.

The crack in the bolt occurred subsequent to

the hot heading operation prior to the production

run. The bolt was quenched in water, dimen-

sionally inspected, and returned to the produc-

tion lot instead of being scrapped. The heavy

decarburization layer on the crack surface sup-

ports this scenario. A schematic of the quench

crack formation is shown in Fig. 35.

Example 13 involves hydrogen embrittle-

ment failure of several cadmium-plated carbon

steel socket head cap screws (Ref 20). The cap

screws were part of a slide valve assembly on a

regenerator line in a petrochemical plant. The

screws were exposed to Gulf Coast atmosphere,

with no exposure to a chemical process or sig-

niﬁcant temperatures. The cap screws failed

during initial loading, while maintenance was

being performed on the valve. One failed and

one unfailed cap screw were sent to a laboratory

for analysis.

The as-received cap screws (Fig. 36) were

visually examined. One of the two screws had

fractured at the head-to-shank radius and was

missing its head. Both screws had been sec-

tioned in the threaded part of the shank (17th and

19th threads) approximately 25 mm (1.0 in.)

Fig. 24

(a) AISI S5 tool steel hammer head that cracked

during heat treatment. The fracture was caused by

quench cracking by the decarburized surface (b) and deep stamp

mark (arrows). Actual size. Source: Ref 16

Fig. 25

Quench crack promoted by the presence of a deep,

sharp stamp mark in a die made of AISI S7 tool steel.

Source: Ref 16

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from the head of the screw. A crack was

observed in the ﬁrst thread root below the

unthreaded part of the shank in the fractured

screw. The screws appeared to have been plated.

The crack in the fractured screw was opened to

reveal its fracture surfaces. Both the initial and

laboratory-opened fracture surfaces of the screw

were examined with a stereomicroscope at

magniﬁcations of 7 to 45 · and with an SEM at

much higher magniﬁcations. The fracture sur-

faces displayed similar fracture modes. Figure 37

shows the overall fracture surface. Figure 38

(a) (b) (c)

(e)(d)

Fig. 26

A 4140 grade steel seamless tubing that failed because of quench cracks. (a) Cross section of tube showing extensive cracking

revealed by dye-penetrant inspection. (b) SEM micrograph showing intergranular fracture at a crack origin. Original mag-

niﬁcation: 90· . (c) SEM micrograph illustrating the brittle mode of failure associated with the fracture. Original magniﬁcation: 50· . (d)

Micrograph showing the typical concentrations of nonmetallic stringers in the tube material. (e) Micrograph showing a quench crack.

Note the intergranular branching and heavy oxide. Original magniﬁcation: 400· . Source: Ref 17

Fig. 27

Failed wrist pin (sample 1), showing fractured faces.

Source: Ref 18

Fig. 28

Schematic of wrist pin. Note stress raisers at “A” and

the rough machining on surface “B.” Source: Ref 18

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shows a ductile fracture mode, which was

observed over the majority of the fracture sur-

faces. Figure 39 shows an intergranular fracture

mode, which was observed around the circum-

ference of the screw, next to the plated surface.

The part of the shank containing the fracture

surfaces was metallurgically prepared in cross

section to look for secondary cracking and

pitting. No secondary cracking or pitting was

observed.

To determine whether the cap screws were

plated, a fracture surface and the outside surface

of the unthreaded part of the shank were cleaned

of oil and other deposits and analyzed using

energy-dispersive x-ray spectroscopy (EDS) in

conjunction with SEM examination. The EDS

Fig. 30

Surface structure along a longitudinal axis of speci-

men 2. The dark matrix is tempered martensite; the

light-colored grain-boundary network is cemen tite. Nital etch.

Original magniﬁcation: 200 · . Source: Ref 18

Fig. 31

Macrograph of sample 2 taken along the longitudinal

axis, showing cracks emanating from both the inner

and outer diameters. Unetched. Original magniﬁcation: 15·.

Source: Ref 18

Fig. 32

Close-up view of the bolt-shank fracture surface.

Note the heavy scale on the zone 1 surface. Source:

Ref 19

Fig. 29

Central longitudina l zone of sample 2, showing

banded structure of white ferrite and dark unresolved

pearlite with MnS inclusions (light gray). 2% nital etch. Original

magniﬁcation: 200·. Source: Ref 18

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Fig. 33

SEM fractography of a ﬁeld on the zone 1 surface (see

Fig. 32). Note the combination of tearing and

intergranular fracture. Source: Ref 19

Fig. 34

Optical micrograph of a portion of the crack along a

cross section of the fractured bolt head. Note the

decarburization at the surface of the crack. Source: Ref 19

Fig. 35 Schematic of quench crack formation. Source: Ref 19

Fig. 36

As-received socket head cap screws. Arrow indicates

a secondary crack in the screw thread root. Source:

Ref 20

Fig. 37

Fracture surface of the crack in the failed screw after

the crack was opened in the laboratory. “L” indicates

the laboratory-induced overload region. Original magniﬁcation:

20·. Source: Ref 20

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results are shown in Fig. 40. It was determined

that the screws were plated with cadmium.

Superﬁcial Rockwell hardness measurements

were taken on the metallographic section. The

average hardness for the failed cap screw was

80.5 HRN, which converts to approximately

40 HRC.

The presence of a ductile fracture mode at the

core and an intergranular fracture mode at the

outer surface of a plated bolt is typical of

hydrogen embrittlement but could also be stress-

corrosion cracking (SCC). However, SCC can

be eliminated, because the metallographic

results showed no evidence of secondary cracks

or other corrosion mechanisms, such as pitting.

Many hydrogen embrittlement mechanisms

have been proposed, but none is universally

accepted. However, the phenomenon of hydro-

gen embrittlement is widely known. The pre-

sence of hydrogen in steel reduces the ductility

of the steel and causes premature failure under a

static load. The time for failure depends on

the stress applied to the component and the

amount of hydrogen that has diffused into the

steel. A component may fail initially when put

under load or may fail several weeks after being

loaded. Because of this characteristic, hydrogen

embrittlement is sometimes called hydrogen-

induced delayed failure.

Electroplating is a common cause of hydro-

gen embrittlement in bolts and screws, because

hydrogen is evolved (or liberated) during the

process. The screws had been plated, and be-

cause no other source of hydrogen was iden-

tiﬁed, it is likely that the plating process was the

source of the hydrogen.

After most plating processes, bolts and screws

are usually baked for several hours at 190



(375



F) to diffuse any hydrogen out of the steel.

However, when a bolt or screw is cadmium

plated, it requires a much longer time (approxi-

mately 24 h) for baking, because hydrogen

Fig. 38

SEM micrograph showing a ductile fracture mode

observed over the majority of both fracture surfaces.

Original magniﬁcation: 1000 · . Source: Ref 20

Fig. 39

SEM micrograph showing an intergranular fracture

mode, observed around the entire circumference at

both fractures in the screw. Structure at top is the base metal;

structure at bottom is cadmium plating. Original magniﬁcation:

1000·. Source: Ref 20

LT= 100 SECS

20K

Iron

Manganese

15K

10K

5000

(a)

(b)

5000

ENERGY keV

0.000 1.000 2.000 3.000 4.000 5.000 6.000 7.000 8.000 9.000 10.000

ENERGY keV

0.000 1.000 2.000 3.000 4.000 5.000 6.000 7.000 8.000 9.000 10.00

COUNTSCOUNTS

BASE METAL OF SCREW

PLATING ON THE SCREWS SURFACE

Cadmium

Fig. 40

Energy-dispersive x-ray spectroscopy spectra of (a)

the base metal of the screw and (b) the plating on the

outside surface. Source: Ref 20

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diffuses less readily through cadmium than other

electrodeposited metals. In the case of the cap

screws, the screws either were not baked at all or

were not baked for a sufﬁcient period of time

and/or at a high enough temperature.

The cap screws failed because of hydrogen

embrittlement. The most probable root cause

was absence of baking or insufﬁcient baking

of the cap screws after the cadmium plating

process.

To eliminate the possibility of future hydro-

gen embrittlement failures, the screws should

be baked at approximately 190



C (375



for 24 h.

Heat Treatment Design

As was analyzed in previous sections, heat

treatments are a series of operations in the course

of which a solid ferrous product is totally or

partially exposed to thermal cycles to achieve

the desired change in structures and properties

(Ref 2). The chemical composition of the

material may possibly be modiﬁed during these

operations (thermochemical treatment).

In addition to the ability of the heat treatment

to achieve desired mechanical properties, heat

treatment also produces dimensional changes

and residual-stress patterns that, in some cases,

can lead to component cracking and distortions.

In the following section, a procedure is ana-

lyzed to improve the performance of the design

process. A typical component manufacturing

process includes metalworking, machining, or

other forming operations, followed by heat

treatment.

Different types of heat treatment of steels

are usually employed by industry: hardening,

austenitization, annealing, normalizing, stress

relieving, quenching and tempering, and aus-

tempering.

Heat treatment processes include component

heatup, holding at temperature for through-heat

solutionizing, or thermochemical treatments

such as carburizing or nitriding, quenching from

elevated temperature, postquench tempering, or

aging treatment. All the steps can inﬂuence

dimensional changes, residual-stress patterns,

and cracking in heat treated components.

The Process of Component Design

In order to avoid failures associated with

heat treatments, it is necessary to develop a

systematic component design process. The

typical phases of component design include

planning and requirements deﬁnition, concept

design, detail design, and test and validation.

The component is designed to provide a speciﬁc

mechanical, thermal, and chemical function

throughout its life cycle and is often limited by

space, cost, and safety considerations. The

selection of materials and manufacturing pro-

cesses for a cost-effective component design is a

complex process and often involves iterative

decision making. The iterative nature of design

requires a continuous analysis and redesign

process.

Process design employs stress-analysis tools

with stress-concentration factors, design rules

based on experimental data, material property

databases, and mechanical properties resulting

from a broad range of heat treatment processes.

Computer modeling is a valuable design tool for

heat treated components. A computer process

simulation model allows a particular design to

be tested under a speciﬁc set of process condi-

tions. The computer software can graphically

display not only the resulting residual stresses

and distortions in the component but also the

associated transient evolution of temperature,

metallurgical phases, volume changes, and

stresses.

Regardless of the procedure used for devel-

oping the design, at the end of the design pro-

cess, the required quality of the product should

be deﬁned and then described in a technical

document, deﬁnitive layout, or ﬁnal design

intended to accomplish the product manu-

facturing. Only the minimum quality needed

for the product to perform the function intended

should be speciﬁed. Overspecifying and in-

cluding restrictive features in the quality des-

cription can cause delays and increase costs to

the buyer.

The design process requires developing

operational deﬁnitions. Operational deﬁnitions

are, for example, a speciﬁc test of a piece of

material or criterion for judgment. Without

operational deﬁnitions, a speciﬁcation is mean-

ingless. A speciﬁcation for heat treatment con-

taining the clause “Avoid long, thin sections”

requires operational meaning of “long, thin

section.” This deﬁnition depends greatly on the

quenching media; however, any section length

greater than 15 times the diameter is almost

always characterized as such, and the slightest

nonuniformity in quench will cause it to distort.

As the quenching medium becomes more severe

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(i.e., water, caustic quench), this criterion is

reduced to as low as 5 times the diameter. For

larger length-to-diameter ratios, consideration

should be given to ﬁxture quenching or induc-

tion hardening.

The component design process can be

divided into two phases:

 Phase 1, which corresponds to the basic

deﬁnition of the product (including concept

and detail design)

 Phase 2, which corresponds to the design

review aiming to prevent failures and mini-

mize risks

Phase 1 deals with the basic deﬁnition of the

product. As a ﬁrst step, the generic type of the

material as well as its geometric conﬁguration

should be selected. The criterion applied to

avoid plastic deformation states that the cal-

culated effective stress must be lower than the

yield and design life creep-rupture stresses of

the material. If the product to be designed is an

element that must withstand not only tension

loads but also bending, torsional, and axial-

compressive loads, then the combined effect of

the applied load type, shape, and size and the

material properties should be analyzed. Design

for those elements subjected to axial compres-

sion should be intended to avoid not only

plastic collapse but also elastic instabilities.

Elastic instabilities may cause Euler buckling

and local buckling. The occurrence of this

failure mode depends on the geometry of the

elements and on the Young’s modulus of the

material. In general, the relationship between

the Young’s modulus (E) and the density (r),

E/r, should be maximized in order to increase

stiffness. By doing so, yield occurs before

buckling, whereas by increasing speciﬁc

strength (s

/r), strength also increases; hence,

buckling occurs before yield. During the basic

design step, brittle fracture should be avoided.

Brittle fracture is associated with very little or

no plastic deformation. A material may fail in a

catastrophic—brittle—manner under stresses

even lower than the allowable design stresses

used to avoid ductile failures. 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, resi-

dual stresses, and stress-concentration factors;

temperature; and loading rate. Carbon and low-

alloy steels undergo a transition from ductile

failure mode to brittle failure mode at low

temperatures. If the material is heat treated

during fabrication, it will have adequate

toughness in its ﬁnal condition. Design must be

reviewed in order to minimize the presence of

notches and defects that concentrate stresses.

When mechanical loads are an alternative,

there may be a risk of fatigue failures.

In fatigue failures, a crack grows in each

loading cycle until the remaining ligament fails

due to ductile or brittle fracture. This phe-

nomenon can occur at stress levels lower than

the allowable stresses for static loads. It should

be emphasized that fatigue failures strongly

depend on design and manufacturing quality,

which is accomplished by increasing fatigue

strength and minimizing stress concentrators.

Technical requirements should be complied

with at the lowest cost. A phase 1 basic design

detailed analysis is beyond the scope of this

work.

Phase 2 design review has the purpose of

assuring that the basic design fulﬁlls the

requirements and reviews the design to avoid

failures. This step veriﬁes that the basic types of

failure modes have been properly controlled by

design and determines the types of damages

associated with each failure mode in order to

implement methods for detection.

Material behavior can be analyzed by

developing models that relate materials attri-

butes, required functions, and manufacturing

processes. Due to the large number of aspects

involved, the problem can be simpliﬁed by

considering blocks of knowledge that corre-

spond to speciﬁc mechanisms and functions.

Each block of knowledge represents a simpliﬁed

model that relates some properties to the

required functions, through the knowledge pro-

vided by materials science and engineering. The

use of state-of-the-art criteria, which, in some

cases, are based on practical experience, can

optimize the accomplishment of the analyzed

functions. The results of the analysis are syn-

thesized in the deﬁnition of design.

Consider two groups of behavior models.

The ﬁrst group relates materials attributes—

generally known as properties—that are well

deﬁned and individually determined. These

properties are component shape and size,

modulus of elasticity, Poisson’s ratio, ultimate

tensile strength, yield strength in tension, shear

strength, compressive strength, ductility, elon-

gation, fracture toughness, hardness, thermal

conductivity, thermal diffusivity, thermal

expansion, speciﬁc heat, density, fracture

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toughness, creep strength, creep-rupture

strength, and so on. For these properties, test

and criteria have accurate operational deﬁni-

tions.

The second group of behavior models in-

cludes attributes that involve the complex

interrelation of a number of variables associated

with materials, manufacturing (heat treatment)

processes, and service conditions. For this

second group of blocks of knowledge, phy-

sical metallurgy is intensely used, together

with the laws of mechanics and empiric knowl-

edge.

Modeling of Heat Treatment

Modeling of heat treatment processes, like

other materials processes such as casting and

welding, is quite complex due to the tight

coupling of various metallurgical transfor-

mations and the associated changes in thermal

and mechanical states. A heat-transfer model,

coupling with a phase transformation model, a

thermomechanical model, and a thermo-

chemical model (Fig. 41), is considered.

Heat Transfer Model

Heat treatments are a series of operations in

the course of which a solid ferrous product is

exposed to thermal cycles. There are different

types of heat transfer: conduction, convection,

and radiation. During heat treatments, the tem-

perature varies in time as well as in space; these

processes are called unsteady, nonstationary, or

transient. The factor of proportionality thermal

diffusivity, a, deﬁnes the rate of change of

temperature, and the heat-transfer coefﬁcient, h,

controls the heat ﬂow through the component

surface. The thermophysical properties, l, r,

and C

, vary with the temperature. The geo-

metric design will try to avoid mass asymmet-

ries. Large differences in section size and

distribution of material causes differential

heating and cooling during heat treating. The

thermal cycle induces phase transformation and

thermomechanical stresses.

Solid-Phase Transformation Model

The thermal cycle induces solid-phase

transformation. Microstructures that are formed

upon cooling and the proportions of each are

Fig. 41 Behavior model to analyze heat treatment design

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