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

Подождите немного. Документ загружается.

from the mono-ball end. The bar was sup-

posedly swaged at 650



C (1200



F) and

then stress relieved.

 The bar was then ﬁnish machined and heat

treated to a tensile strength range of 1930 to

2070 MPa (280 to 300 ksi). The part was

Fig. 7

SEM fractographs documenting typical topographic features of the T-head. (a) Typical portion of fracture surface, showing

typical origins at arrows (500 mm). (b) Typical fatigue striations on fracture surface of T-head (1 mm)

Fig. 8

Micrographs documenting the microstructure of the failed strain bar. (a) Section 1-1 from Fig. 4 showing a banded structure

(670 mm). (b) Location A, retainer ring groove (300 mm). (c) Section 2-2 from Fig. 4 showing banded structure (670 mm).

(d) Location B at origin showing banded structure (300 mm)

356 / Failure Analysis of Heat Treated Steel Components

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:45PM Plate # 0 pg 356

then magnetic particle inspected. Electro-

deposited nickel was speciﬁed for plating on

the inside diameter surface.

This plating is relatively soft, ductile, and

has a relatively high melting point. However,

electroless nickel, which has high phosphorus

content, was substituted without authorization.

This plating is hard and brittle and starts to melt

at the eutectic temperature of 880



C (1616



F).

Figures 16 and 17 show the MLG linear actu-

ating piston rod cylinder components as re-

ceived for examination. Two circumferential

fractures were located approximately 165 and

215 mm (6

and 8

in.) from the mono-ball

end of the piston rod cylinder. These two frac-

tures were joined by a longitudinal crack

approximately 50 mm (2 in.) long.

The longitudinal fracture was opened, and the

fracture surfaces were examined visually and at

up to 50 · magniﬁcation using a stereomicro-

scope. Figure 18 shows macrographs of the

fracture surfaces. The circumferential fractures

originated and terminated on the longitudinal

crack. The longitudinal crack exhibited a single,

intergranular-appearing origin on the inside

diameter surface of the piston rod cylinder. The

origin was located at approximately midlength

of the longitudinal crack, 200 mm (8 in.) from

the mono-ball end. An arrest mark was present

on the fracture surface at approximately 85% of

the wall thickness.

Fig. 9

Macrograph documenting the appearance of the

as-received main landing gear lever attach pin

Fig. 10

Magnetic particle inspection indications on the

ﬂange of the main landing gear lever attach pin

Fig. 11

Overtempered (darkened areas) in the ﬂang e

revealed by temper etching (2.5 mm)

Case Studies of Steel Component Failures in Aerospace Applications / 357

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:45PM Plate # 0 pg 357

An SEM was used to examine the fracture

surface and to document the mode of failure.

Figure 19 shows typical SEM fractographs made

of the fracture surface. The fracture topography

at the origin was intergranular. The dimensions

of the origin were 0.5 mm wide by 0.19 mm

deep (0.02 in. by 0.0075 in.). Away from the

origin was a static rupture area, which was

characterized by dimples. Approximately 85%

across the thickness, an arrest mark was present.

As shown in Fig. 19, the fracture topography in

the arrest was intergranular. Past this zone to the

edge of the specimen, the fracture topography

consisted of dimples (static rupture).

Hardness measurements were made on the

piston rod cylinder. The hardness was 54 HRC,

and this met the hardness requirement of 53 to 55

HRC.

Fig. 12 Discoloration of the fracture surface (330 mm)

Fig. 13

Topographic features of the fracture surface as observed by the SEM. (a) Overall view of the fracture surface (111 mm).

(b) Intergranular fracture observed on the fracture surface (10 mm)

Fig. 14

Metallographic examination of cracks evident in the ﬂange of the main landing gear lever attach pin, showing loss of

chromium at cracks (20 mm)

358 / Failure Analysis of Heat Treated Steel Components

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:45PM Plate # 0 pg 358

A hydrogen analysis was conducted on the

piston rod cylinder. The hydrogen content was

determined to be 2.3 ppm. This amount of hy-

drogen is considered sufﬁcient to cause hydro-

gen embrittlement (Ref 5).

A metallographic specimen, section 1-1, was

removed at the origin location shown in Fig. 19.

The specimen was prepared using standard

metallographic techniques. The microstructure

at the origin is shown in Fig. 20. The micro-

structure of the steel, which consisted of tem-

pered martensite, was normal.

The plating at the origin consisted of three

distinct layers. The layers were subjected to an

energy-dispersive x-ray analysis. The outermost

two layers consisted essentially of nickel and

phosphorus, whereas the inner layer consisted of

nickel and iron. The outermost two layers con-

tained cracks, and the inner layer did not. Also

evident in the micrograph is plating that had

covered a small part of the fracture surface. This

gave evidence that a crack was present prior to

plating or prior to the heating of the plating,

which occurred during swaging.

A single intergranular origin was associated

with the fracture surface, and this origin was

relatively clean. These factors suggest a delayed

mode of failure, as occurs in a hydrogen

embrittlement failure. The bulk hydrogen was

analyzed to be 2.3 ppm; this concentration

would most likely be higher at the crack tip.

The intergranular origin was separated from

the intergranular arrest zone by a ductile rupture

region, indicating that the crack had propagated

in a ductile manner before arresting. Hydrogen

most likely diffused to and concentrated at the

crack tip, which caused the intergranular arrest

region to occur. The crack then propagated to

failure by ductile rupture.

Based on the results of this investigation, it is

concluded that the MLG linear actuating piston

rod cylinder failure was most likely due to

hydrogen embrittlement.

Failure Analysis of AISI 420 Stainless

Steel Roll Pin

Several failures of AISI stainless steel roll

pins were reported. This pin is a standard part

and is used to hold pin components together by

expanding after compression. The pin is manu-

factured from AISI 420 stainless steel and is heat

treated to 46 to 55 HRC.

Figure 21 shows the as-received failed roll

pin. The fracture extended the length of the pin.

An SEM was used to examine the fracture

surface of the service failure and a laboratory-

created overload failure from the same lot of

material. Figure 22 shows the results of the SEM

examination. In this ﬁgure, the topology of

the fractures was intergranular, with some dim-

ples present. This failure mechanism indicated

that the failure was due to overload (static

rupture).

Fig. 15 Micrographs showing the morphology of the cracks. (a) Overall view (50 mm). (b) Closeup of crack (20 mm)

Fig. 16

As-received main landing gear linear actuating

piston rod cylinder

Case Studies of Steel Component Failures in Aerospace Applications / 359

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:46PM Plate # 0 pg 359

Fig. 17 Appearance of longitudinal crack in the main landing gear linear actuating piston rod cylinder

Fig. 18 Fracture surfaces. (a) Circumferential crack 1 (4.3 mm). (b) Circumferential crack 2 (4.3 mm). (c) Longitudinal crack (5 mm)

360 / Failure Analysis of Heat Treated Steel Components

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:46PM Plate # 0 pg 360

Hardness measurements were made on the

roll pin to verify the heat treated condition.

The hardness of the roll pin was measured to

be 49 HRC, which met the hardness requirement

of 46 to 55 HRC.

A metallographic specimen was prepared

through the fracture surface. The specimen was

prepared using standard metallographic tech-

niques. The microstructure (Fig. 23) consisted

of tempered martensite, with carbides out-

lining the prior-austenitic grain boundaries.

This is not a typical microstructure for quen-

ched and tempered AISI 420 stainless steel

(Ref 2).

Generally, because of their high harden-

ability, martensitic stainless steels can be quen-

ched in either oil or air. Oil quenching ensures

maximum ductility and corrosion resistance. Air

Fig. 19

SEM fractographs documenting the appearance of the fracture surface. (a) Origin location (670 mm). (b) View of box A

showing fracture origin. (c) Location A (5 mm). (d) Location B (2 mm). (e) Location C (5 mm). (f) Location D (2 mm)

Case Studies of Steel Component Failures in Aerospace Applications / 361

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:46PM Plate # 0 pg 361

quenching may cause decreases in ductility and

corrosion resistance. If slow cooled or an

inadequate quench through the critical range

of 870 to 540



C (1600 to 1000



F), this steel

will precipitate carbides at the grain boundaries

(Ref 2). Because of the presence of the inter-

granular carbides, it is likely that this part

experienced an inadequate quench.

Based on the results of this investigation, it

was thought that the failure mechanism was

overload, with limited ductility caused by

improper quenching during heat treatment.

Failure Analysis of a Main Landing

Gear Lever

The MLG lever was removed from service

after a hard carrier landing. Multiple cracks

developed during removal of the ion vapor-

deposited (IVD) aluminum with sodium hydro-

xide to analyze for residual stresses by x-ray

diffraction.

The lever was machined from 300M steel

forging into a hollow conﬁguration, machined,

then heat treated to a tensile strength level of

1930 to 2070 MPa (280 to 300 ksi). The lever

was then IVD coated on the outside.

Figure 24 shows the failed MLG lever as

received for examination. The location of a

primary crack and a series of secondary cracks,

which were between the up-latch and oleo lugs,

is also shown in Fig. 24. The primary crack was

approximately 23 mm (0.9 in.) long and is

shown in Fig. 24.

The primary crack was opened, and the frac-

ture surfaces were examined visually and at up

to 50 · magniﬁcation using a stereomicroscope.

Figure 25 shows a macrograph of the fracture

surface. Also shown in Fig. 25 are the appear-

ances of origins (some of which were dis-

colored), which were located on the outside

surface of the lever. The origin areas had a

faceted appearance, which indicate a delayed

mode of failure, that is, stress-corrosion crack-

ing or hydrogen embrittlement (Ref 6).

An SEM was used to examine the fracture

surface. Figure 26 shows SEM fractographs

documenting the topographic features of the

fracture origins. The fracture topography was

intergranular, which is indicative of a delayed

mode of failure.

The wall thickness at the origin location was

measured to be 5.8 mm (0.229 in.). The drawing

Fig. 20 Microstructure at fracture origin (12.5 mm)

Fig. 21

As-received 420 roll pin. (a) Visual view of

as-received roll pin (2.5 mm). (b) SEM view of

as-received roll pin (200 mm)

362 / Failure Analysis of Heat Treated Steel Components

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:46PM Plate # 0 pg 362

requirements for the maximum and minimum

wall thicknesses were 5.33 and 4.45 mm (0.210

and 0.175 in.), respectively. Therefore, the

drawing requirement was exceeded by 0.48 mm

(0.019 in.).

Hardness measurements were made on the

lever to verify heat treatment. The hardness of

the lever was 54.2 HRC, and this met the hard-

ness requirement of 53 to 55 HRC.

Hydrogen analyses conducted on the lever

yielded values of 3.0, 4.6, 3.4, and 3.5 ppm. The

average of the four values was 3.63 ppm, which

is considered high enough to produce a hydro-

gen embrittlement failure.

A metallographic specimen was removed

through the fracture origin area. The specimen

was prepared using standard metallographic

techniques. Figure 27 shows the microstructure

of the specimen. The microstructure of the lever

was tempered martensite, and this is normal for

a quenched and tempered high-strength, low-

alloy steel.

Based on the results, it is thought that the

cracks formed in the MLG lever as a result of

Fig. 22

SEM examination of the failed roll pin and laboratory-produced overload fractures. (a) Location A of the service failure

(20 mm). (b) Location A of the service failure showing intergranular fracture with some dimples (5 mm). (c) Laboratory-

produced overload failure showing intergranular fracture (20 mm). (d) Laboratory-produced failure showing intergranular fracture with

dimples (5 mm). Compare to (b)

Fig. 23

Microstructure of failed roll pin. Micro structure

consists of tempered martensite with carbides de-

corating the prior-austenite grain boundaries (10 mm)

Case Studies of Steel Component Failures in Aerospace Applications / 363

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:46PM Plate # 0 pg 363

the application of sodium hydroxide to remove

the IVD coating in addition to the residual

stresses present as a result of the hard carrier

landing. It is likely that accelerated stress-

corrosion cracking occurred because of the high

residual stresses and electrolyte.

Failure Analysis of an Inboard Flap

Hinge Bolt

An inboard ﬂap hinge bolt was found to be

failed after 286.5 ﬂight hours. The bolt had been

machined from 4330V-modiﬁed steel bar and

Fig. 24

Appearance of the main landing gear lever showing the location of the primary and secondary cracks. (a) Overall view

(33 mm). (b) Location of primary and secondary cracks at site of ion vapor deposit (IVD) removal (10 mm)

364 / Failure Analysis of Heat Treated Steel Components

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:46PM Plate # 0 pg 364

heat treated to the 1515 to 1655 MPa (220 to

240 ksi) tensile strength range. A nut with the

bolt was machined from 4140 steel bar heat

treated to the 1240 to 1380 MPa (180 to 200 ksi

tensile strength range.

Figure 28 shows the inboard ﬂap hinge bolt

and nut as received for examination. The bolt

was approximately 102 mm (4 in.) long. The

bolt consisted essentially of the shank, which

was approximately 23 mm (0.9 in.) in diameter

and 64 mm (2.5 in.) long, and a threaded por-

tion, which was 23 mm (0.9 in.) in length with

thread type 0.5000-20 UNJF-3A THD in accor-

dance with MIL-S-8879. The unthreaded por-

tion of the bolt was chromium plated, while the

threaded portion was aluminum IVD coated.

The fracture had occurred at the transition of the

shank to the threaded portion of the bolt. The

fracture surface was examined visually and at up

to 30 · magniﬁcation using a stereomicroscope.

Figure 29 shows a macrograph of the fracture

surface. Fracture ridges emanated from an origin

that was located on the outer diameter surface

of the bolt at the location shown in Fig. 29.

The origin exhibited a reﬂective, intergranular

appearance. Apparent corrosion products were

also observed on the fracture surface.

Residual stresses were measured on the sur-

face of the bolt at locations 1 and 2, shown in

Fig. 28. The residual stresses were measured

using an x-ray residual-stress analyzer. Prior to

measuring the residual stresses, the chromium

plating (location 1) and the aluminum IVD

coating (location 2) were removed. The residual

stress measured at location 1 was 790+

100 MPa (–115+15 ksi), which indicated that

this area was properly shot peened as required

according to the engineering drawing. The resi-

dual stress measured at location 2 was 140+

100 MPa (20+15 ksi). This indicated that

this area was not shot peened. The engineering

drawing does not require this location to be shot

peened.

An SEM was used to examine the fracture

surface and to document the topographic fea-

tures. Figure 30 shows SEM fractographs taken

of the fracture surface. The fracture topography

at the origin was intergranular. The intergranular

topography extended for approximately

the way across the fracture surface, at which

point dimples indicative of ductile rupture were

present.

A laboratory-produced failure yielded a struc-

ture that consisted only of dimples.

The corrosion products on the service failure

fracture surface were subjected to an energy-

dispersive x-ray analysis. Calcium, potassium,

magnesium, and chlorine were identiﬁed as

being present.

Rockwell hardness measurements were made

on the inboard ﬂap hinge bolt to verify heat

treatment condition. The hardness of the bolt

averaged 46 HRC. This met the hardness

Fig. 25

Appearance of primary crack removed from part.

(a) Region of primary crack where ion vapor-

deposited coating had been removed (13 mm). (b) Primary crack

showing branching (2 mm). (c) Fracture surface of crack after

opening (3 mm)

Case Studies of Steel Component Failures in Aerospace Applications / 365

Name ///sr-nova/Dclabs_wip/Failure_Analysis/5113_351-393.pdf/Chap_11/ 18/8/2008 3:46PM Plate # 0 pg 365