fatgue paper for mn alloys good concept

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Stress corrosion cracking of a wrought Mg–Mn alloy

under plane strain and plane stress conditions

Bharat S. Padekar a,b,c, V.S. Raja b,⇑, R.K. Singh Raman c,d

a IITB-Monash Research Academy, Indian Institute of Technology Bombay, Mumbai 400 076, Indiab Department of Metallurgical Engineering and Materials Science, Indian Institute of Technology Bombay, Mumbai 400 076, Indiac Department of Mechanical and Aerospace Engineering, Monash University, Clayton, Victoria 3800, Australiad Department of Chemical Engineering, Monash University, Clayton, Victoria 3800, Australia

a r t i c l e i n f o

Article history:

Received 12 July 2012

Received in revised form 29 December 2012

Accepted 1 February 2013

Keywords:

Magnesium alloy

Stress intensity factor

Stress corrosion cracking

Life prediction

a b s t r a c t

Stress corrosion crack growth of a rare earth containing Mg–Mn hot rolled alloy was stud-

ied using compact tension and circumferential notch tensile specimens employing con-

stant load in 0.1 N NaCl solution saturated with Mg(OH)2. Circumferential notch tensile

specimens satisfied plane strain conditions; Log crack velocity versus stress intensity plots

yielded slope of 0.5 and 1 for stage I region and threshold stress intensity of 10 and

13MPa m1/2 for circumferential notch tensile and compact tension specimens respectively.

Fractography showed transgranular features involving hydrogen in stress corrosion

cracking.

2013 Elsevier Ltd. All rights reserved.

1. Introduction

Magnesium alloys are among the best lightweight metallic structural materials available for engineering applications.

Their relative high strength-to-weight ratio makes them potential candidates for applications in transport and aerospace

industries. Mg alloys are used in both cast and wrought conditions. Later type of alloys is used where components are made

through fabrication.

Development of the wrought Mg alloys can be divided into four broad categories, i.e., Mg–Zn, Mg–Al, Mg–Th, and Mg–Mn

alloys [1,2]. Among them Mg–Mn alloys i.e. M1A (Mg–1.2Mn), MA1 (Mg–1.9Mn) and MA8 (Mg–2Mn–0.25Ce), the relatively

low cost alloys are available in the form of extrusions, sheets, and plates. As the precipitation-hardening effect is small, M1A

is nonheat-treatable alloy. Presence of Mn in the alloy has several beneficial effects. Manganese is usually added to most of

the Mg-based alloys to remove, right at the melting stage, the impurity elements, such as Fe; for the purpose of improving

the corrosion resistance [3]. Furthermore, Mn enhances rollability [4] of magnesium alloys. During solidification of Mn-containing Mg-alloys, Al–Mn compounds precipitate, in preference to Mg–Al precipitates at high temperature. Therefore,

increasing the Mn content reduces Al concentration in the solidifying liquid and, consequently, the amount of Mg–Al pre-

cipitates (such as Mg17Al12) is cathodic to alpha matrix thereby facilitating localised corrosion. Mn addition is also reported

to refine the grains [4]. It should be noted that Mg–Mn alloys generally contain rare-earth such as Ce, which is known to

improve tensile, and creep properties by refining the grains [5]; addition enhances passivation behaviour [6] which, in turn,

can affect the SCC behaviour of alloys.

As magnesium alloys are generally susceptible to SCC [7], their successful application and failure avoidance depend on

evaluating and improving their resistance to crack growth. SCC behaviour of Mg–Mn wrought alloys has generally received

0013-7944/$ - see front matter 2013 Elsevier Ltd. All rights reserved.http://dx.doi.org/10.1016/j.engfracmech.2013.02.018

⇑ Corresponding author. Tel.: +91 22 2576 7892; fax: +91 22 2572 3480.

E-mail address: [email protected] (V.S. Raja).

Engineering Fracture Mechanics 102 (2013) 180–193

Contents lists available at SciVerse ScienceDirect

Engineering Fracture Mechanics

j o u r n a l h o m e p a g e : w w w . e l s e v i e r . c o m / l o c at e / e n g f r a c m e c h

http://dx.doi.org/10.1016/j.engfracmech.2013.02.018mailto:[email protected]://dx.doi.org/10.1016/j.engfracmech.2013.02.018http://www.sciencedirect.com/science/journal/00137944http://www.elsevier.com/locate/engfracmechhttp://www.elsevier.com/locate/engfracmechhttp://www.sciencedirect.com/science/journal/00137944http://dx.doi.org/10.1016/j.engfracmech.2013.02.018mailto:[email protected]://dx.doi.org/10.1016/j.engfracmech.2013.02.018http://crossmark.dyndns.org/dialog/?doi=10.1016/j.engfracmech.2013.02.018&domain=pdf


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very limited research attention. However, there are reports on the role of Mn-addition on SCC of other Mg alloys. Logan and

Gessing [8] have shown that an alloy of MA3 (Mg–6Al–1Zn–0.3Mn) type sandwiched between sheets of the MA1 type alloy

does not suffer SCC in a marine atmosphere under a tensile stress of 210 MPa. Similarly, another Mg alloy (8% Al, 0.5% Zn, and

0.2% Mn) sandwiched between MA1 sheets [9] also did not suffer SCC in 200 days at a tensile stress of 150 MPa, while the

unprotected sheets cracked within a few days of exposure. Among extremely limited SCC study on Mg–Mn alloys, Nozaki

et al. [10] found an Mg–2% Mn–0.5% Ce alloy to undergo SCC in distiled water and 0.5% potassium hydrogen fluoride solution.

In their critical review of SCC of magnesium alloys, Winzer et al. [7] have reported transgranular SCC behaviour of Mg–Mnalloys. However, no published literature is available on the stress corrosion crack growth of Mg–Mn alloys. Hence this work

aims at understanding SCC behaviour of Mg–Mn (0.1Al–1.75Mn–0.3Zn–0.3 other RE–0.25Ce) alloy. As SCC crack growth data

is very important not only to understand the SCC tendency of the alloy but also design structures against SCC failure, this has

also been obtained. Since, circumferential notch tensile (CNT) specimen allows use of thinner section specimens that still

comply with plane strain conditions [11–18] this has been used along with CT specimen. Work was carried out on both

pre-cracked CT and CNT specimen under constant load test condition in chloride containing environment.

2. Experimental procedures

2.1. Material

The test alloy was commercially available magnesium based hot rolled alloy that was annealed at 300–350 C for 30 min

and then air cooled. Its chemical composition in weight% was analysed to be 0.1Al–1.75Mn–0.3Zn–0.007Ni–0.002Be–


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For microstructural examination using optical microscopy, specimens were prepared by standard metallographic proce-

dures and were etched in a solution of ethanol 100 ml and nitric acid 8 ml. Grain size was determined using the mean linear

intercept method [19]. Energy dispersive X-ray spectroscopy (EDX) analysis was carried out for the characterisation of the

broad chemical composition of the secondary-phase particles, which included X-ray mapping of precipitates and surround-

ing areas.

Tensile tests were carried out in triplicate using extensio metre on Zwick/Roell universal testing machine (UTM) using

cylindrical tensile specimens of standard proportions (diameter 6 mm 25 mm gauge length) were machined from rolled

plate; these being aligned both in longitudinal and transverse axis direction. Yield strength ( rYS

), Young’s modulus (E ), ulti-

mate tensile strength (rUTS), and elongation (El) are respectively 140 MPa, 44 GPa, 200 MPa, and 30% (28% in transverse).

2.2. Specimen design

A few slow strain rate (106 s1) tests were carried out in 0.1 N NaCl solution saturated with Mg(OH)2 using smooth ten-

sile specimens machined along the longitudinal and transverse of the rolling direction. From this study transverse specimens

were found to be more susceptible to SCC than the longitudinal specimens. Accordingly the specimens for CT and CNT tests

were pre-cracked along the rolling direction to obtain more conservative values.

As the available thickness of plate was only 12.2 mm, the dimension proportions of CT specimen were chosen to be 1/2 T

i.e. thickness, (B) of 1/2 in. (12.7 mm) and for a W /B ratio of 2, W becomes 1 in. (25.4 mm), (where W is width of CT specimen)

for which pulling grip set for fatigue pre cracking and constant load test were available in the laboratory. The drawing of CT

specimen is shown in Fig. 1a. The specimen was machined in the T–L alignment [20] so that the crack propagated in the roll-

ing direction. A few specimens were provided with dovetail notch for fitting the clip gauge. Side milling and drilling of holeon CT specimens were carried out using a CNC milling machine. Straight through notch was produce by EDM wire cut with

1 mm end slot of 0.16 mmradius. The notch was polished manually using 1000 grit size silicon carbide paper to remove EDM

marks. The two side surfaces of CT specimen were polished successively till 5000 grit size silicon carbide paper with a final

finish using diamond paste of 0.25 lm grit. Lines to limit the fatigue pre crack length were marked on both sides of specimen.

The alignment of CNT specimen was such that its axis was along the transverse direction of the plate. The drawing of CNT

specimen is shown in Fig. 1b. CNT specimens were machined on lathe machine and finally groove of 60 angle with root ra-

dius of 0.15 mm was produced.

2.3. Pre-cracking

Fatigue pre-cracking of CT specimen was done at the laboratory temperature with 4 Hz frequency and load ratio (R) as 0.1

using load control sine wave. The initial K max was 6 MPa m

1/2

with 2

10

4

loading cycles, and for the final 1 mm length theload was lower such that the final K max was 3.7 MPa m

1/2 and loading cycles were of the order of 5 103. Side grooves using

an end mill cutter of 60 nose angle and end radius of 0.15 mm were milled to a depth of 5% of the specimen thickness per

side on both the sides of the specimen. Deburring and cleaning were done before loading the specimen on a constant load

test rig.

Fig. 1. Drawing of: (a) CT, and (b) CNT specimen (All dimensions are in mm unless specified).

182 B.S. Padekar et al. / Engineering Fracture Mechanics 102 (2013) 180–193


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CNT specimens were pre-cracked at laboratory temperature using a rotating cantilever bending machine. The rotating

frequency was about 46 Hz (2800 RPM) and R as 1. The distance between groove centre and point of application of load

was kept constant; the dead load was so applied that the resulting stress intensity was approximately 25% of the K IC. To

achieve the required fatigue pre-crack depth, time was pre-set from the initial trials. The measured fatigue pre-crack depth

varied from 0.4 to 1 mm. Fatigue pre-cracked specimens were wrapped with PTFE tape around the notch and were stored in

a desiccator to avoid any undue corrosion at the crack-tip, due to the laboratory environment.

2.4. Experimental setup

A horizontal loading frame was used with a calibrated load cell of 1000 kg capacity and least count of 1 kg for measure-

ment of the applied load for testing CT specimens. The load was applied through a compressed helical spring having spring

index (250 N mm1) such that the load did not decrease by greater than 1% of the applied load, due to the change in load line

deflection before resetting time (normally 12 h). Load was applied slowly so that the applied K I rate was below 100 MPa m1/2

per minute [21]. Load was recorded at least twice a day throughout the test period and the load was adjusted, as and when

required, by tightening the loading nut. Pull rod was provided with a flat locator with loading frame so as to avoid its rotation

while tightening the loading nut. The pulling grips were provided with loading pins coated with alumina and wrapped with

PTFE so as to facilitate the adjustment of the specimen to proper loading line and allow rotation of the specimen during test-

ing [20]. Loading clevis was provided with PTFE washers on each side of the specimen to avoid galvanic coupling with the

clevis material. The corrosive environment for CT and CNT testing was 0.1N NaCl solution saturated with Mg(OH)2 which has

a pH of 10.5. The pre-cracked CT specimen was subjected to this corrosive environment by dripping of the solution at a rate

of 1 ml per minute. The drip solution was replaced with fresh solution every week throughout the test period. The drippingarrangement further ensured avoidance of the galvanic coupling with other gripping members.

For CNT specimens constant load test rigs that work on reactive force of helical compression springs were used which are

vertically arranged. The springs were calibrated on UTM for load versus deflection. The load was applied by tightening the

loading nut in such way that the standard loading rate was not exceeded [21]. Pull rod of the rig is provided with flat locator

to avoid rotation of grip. A clock was connected to the end cup of spring (also for CT specimens) through banana clip so that

after failure of specimen, banana clip disconnects and stops the clock, enabling determination of an accurate time to fail.

Three similar CNT rigs were used for this work. For CNT specimen, a corrosion cell of 40 ml capacity, made of transparent

acrylic was placed such that the notch is at the mid of the solution level in cell. The environment used was the same as

for the CT tests. Once again, the corrosive solution was replaced with fresh solution every week throughout the test period.

Loss in solution level because of evaporation was compensated by topping up with distilled water.

2.5. Measurement of crack size

Crack lengths for CT specimens were measured on the fractured surface using a travelling microscope. One side of CT

specimen was placed parallel to the long slide scale of travelling microscope. Normally the crack front of CT specimen

was observed to be parabolic in shape. In order to average the crack size, more accurately the ‘‘nine lines’’ method was used

as shown in Fig. 2a. All the nine measurements were made at equispaced distances, keeping a margin of 0.005W (0.1 mm)

from the side grooves as shown in Fig. 2a. The initial crack length in CT specimen, ai corresponds to the distance between the

centre line (C/L) of the loading pin hole and the end of the fatigue pre-crack zone. The final crack size, a f corresponds to the

final crack front measured on the fractured surface considering SCC region [22]. Magnifying optical system was used with

10 in microscope to observe the crack front. For a CT specimen, the average crack length was obtained from the nine mea-

surements using the formula [22]:

a ¼1

8

a1 þ a92

þ

X8

i¼2

ai

( ) ð1Þ

For a CNT specimen the crack size was also measured on the fractured surface using a travelling microscope. The ligament

of CNT was generally found to be eccentric after fatigue pre-cracking (as shown in schematic Fig. 2b). In order to account for

the eccentric (e) ligament, the fracture surface of specimen was divided into 12 wedge-shaped elements [12] at intervals of

30 for which an indexing jig was used. Schematic of the indexing jig is shown in Fig. 2c. A section of the CNT specimen with

the fracture surface facing upward was pasted on the central circular slot of jig. Approximate centre of outer diameter was

obtained by sliding both the axes of the travelling microscope. After every straight measurement (such as 0 and 180 ), the

arm was rotated and locator pin was put in the next index hole of the jig. The radial depth of a f and am was measured at the

12 locations.

2.6. Determination of K I

CT specimens were loaded to certain loads assuming initial crack size such that the initial K I was below K IC. Parameters

like B, Bn

, and W were measured before loading. After failure of the specimen or after the surviving specimen was forced to

fracture, the initial crack size was measured on the fracture surface. K I was calculated as per ISO and ASTM standards

B.S. Padekar et al. / Engineering Fracture Mechanics 102 (2013) 180–193 183


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[20,21,23]. Similar to CT specimens after measuring the crack size on fractured surface of CNT specimens K I was determined

using Eq. (2) [13]:

K I ¼ ðrt þ rbÞ ffiffiffiffiffiffiffiffiffiffiffiapF o

p ð2Þ

Further details and calculation method can be found elsewhere [13,14].

2.7. Allowance for eccentricity (e)

One of the critical issues in determination of K IC by CNT method is the possible effects of the eccentric shape of the final

ligament produced during fatigue pre-cracking by rotating-bending (as shown in schematic Fig. 2b). When the ligament is

off-centre, the final fracture of the specimen will result not only from tension force but will also have a component of the

bending forces. Therefore, to account for the effect of the eccentricity, there is a bending stress term (rb) in Eq. (2). The spe-

cific method [12] briefly described in Section 2.5 was used for the determination of the centroid of the ligament, in order to

determine the bending stress applied to the specimen as a result of the eccentricity. The eccentricity of the ligament is cal-

culated using the first moment of area. The details of the eccentricity’s calculations are discussed in other publications by

Ibrahim and Stark [13].

2.8. Validity requirement

Linear elastic fracture mechanics (LEFMs) is applicable with the assumption that the materials deform in elastic/brittle

manner. In reality, engineering materials undergo some plastic deformation at the crack tip before they fail. The LEFM ap-

proach is only valid if the size of the plastic zone is negligible compared to the specimen geometry.

Validity check for every CT specimens was done according to the criteria as described elsewhere [20,21,23].

Specifications for determining the validity for CNT of measuring K I have been used [11–18]. Validity limits require: (a) the

fatigue crack depth to be at least twice the Irwin plastic zone [12,14], and (b) the average stress across the ligament after

fatigue cracking should not exceed 2.5 times the yield strength. These limits were also applied to those of CNT specimens

Fig. 2. Crack measurement method (a) CT [22], (b) CNT specimen [13], and (c) schematic of indexing jig used for CNT specimen.


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where the final ligaments were eccentric. With an eccentric ligament, the maximum nominal stress considered was then a

combination of tensile and bending stresses, which should not exceed 2.5 times the yield strength. Further, for such eccentric

ligament cases, the fatigue crack depth considered for the purposes of validity was taken to be the greatest depth of the

crack.

2.9. Determination of crack velocity

SCC crack length for CT specimens was obtained from the difference of final and initial crack lengths, (a f –ai) and the aver-age crack velocity was obtained as:

da

dt ¼

SCC length ðmÞ

Time to fail or testing ðsÞ ð3Þ

The crack velocity for CNT specimens was determined by using Eq. (4) [16–18]:

da

dt ¼

dK I dt

da

dK I ð4Þ

The term dK I /dt of Eq. (4) is determined from the relationship between K I and time-to-failure, t f , obtained by fitting the

data of initial K I and time to fail in constant load test. The second term, da/dK I , can be determined from Eq. (2) [16–18]. Inthe

current work the crack growth rate was also calculated from the difference of average final and initial crack depths (crack

extensions) and time to failure or time of test in case of survived specimens [18]. These two methods were used for CNT data

and plotted on the same plot for comparison.

2.10. Fractography

Fractographic features were characterised using scanning electron microscope (SEM) of Hitachi make, model S-3400N.

However a stereomicroscope was first used to examine the overall characteristics of fractured specimen. In the case of CT

specimens, the region, containing fatigue, SCC, and mechanical failure were separated. Surface oxides were removed by soak-

ing the sectioned specimens in boiling solution of 20 wt.% chromic acid for 20 s then rinsed in distiled water, dried, and ultra-

sonically cleaned with acetone before fractographic examination.

3. Results

3.1. Microstructure and corrosion resistance

Typical optical micrograph of the alloy is shown in Fig. 3a. Equiaxed grains observed in the microstructure are typical of

wrought magnesium alloys. Based on the linear intercept method [19], average grain size of the alloy was measured to be

1 2 ± 3 lm. The back scattered electron imaging reveals the presence of second phases ( Fig. 3b), which are found to align

along the rolling direction. In addition, fine globular particles are also noticed. EDX analysis revealed that the globular par-

ticles containing Mg, Mn and Ce (Fig. 4), where as the X-ray mapping of precipitates aligned in the rolling direction (Fig. 5)

confirms these precipitate to be rich in Ce and Mn as well as lean in Mg with respect to the adjacent area. These precipitates

are reported to be ternary intermetallic compound (Mg, Mn)12Ce having the same tetragonal structure as Mg12Ce [24]. The

Mn content in Mg12Ce is reported to very between 0.3 and 0.6 at.%, depending on alloy composition and quenching temper-

ature [24]. This precipitate composition is in agreement with those reported by Liu et al. [6] for AM Ce0.5 magnesium alloy.

According to them, Ce addition facilitated the formation of a new phase having a needle-like morphology. Pan and Yang [5]

also suggested that adding 0.87 wt.% Ce to the Mg–3Sn–1Mn alloy leads to the formation of a new phase namely Mg12Ce.

Impurities such as Fe, Cu and Ni have profound influence on the corrosion rate of Mg-alloys. That can be easily affected in

the rate of hydrogen evolution [25]. For quantitative assessment of hydrogen evolution, the alloy sample was mounted in an

epoxy resin so that a surface area of 2 cm2 was exposed to NaCl solutions. The working surface was polished to a 0.25 lm

diamond finish and then immersed in the test solution. The hydrogen evolved during the immersion test was collected, fol-

lowing the procedure of Song et al. [26]. The hydrogen evolution results are presented in Fig. 6, the corrosion rate was cal-

culated to be 0.27 and 0.56 mg cm2 day1 in 0.1 N and 3.5 wt.% NaCl solutions respectively. It is suggested that when the

corrosion rate of Mg alloys in 3 wt.% NaCl solution is below 10 mg cm2 day1 then microstructure strongly influences the

corrosion rate and when the corrosion rate is between 10 and 100 mg cm2 day1 then Fe impurity has been found to influ-

ence the corrosion rate [25]. Accordingly, the considerably low corrosion rate of the present alloy agrees with the high purity

level (


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of 3802 N and a P Q of 3427 N. After measuring the specimen dimensions and crack size (ai), K I was calculated from load, P Q which turns out to be 23 MPa m1/2. It was found that the CT specimen would satisfy plane strain conditions only when its

thickness exceeds 67.5 mm. Accordingly the estimated value of 23 MPa m1/2 (within 3% error) corresponds to K QC.

As CNT specimen satisfy plain strain conditions [11], pre-cracked CNT specimens were subjected to axial tensile tests for

obtaining valid plane strain fracture toughness, K IC. A pre-load of 200 N was applied; the extension rate during the test was

0.025 mm s1, and the maximum load attained was 4620 N. Other test parameters were; (a) initial crack size, ai, 2.364 mm,

(b) ligament diameter, d, 4.74 mm, and (c) eccentricity of ligament, e, 0.167 mm. K IC thus calculated using Eq. (2) was found

to be 15 MPa m1/2 which was considerably lower than K QC of 23 MPa m1/2. Each of the four specimens satisfied the valid

plane strain conditions [15].

Fig. 3. (a) Optical micrograph of Mg–Mn alloy obtained on the rolled surface (aligned horizontal) and (b) SEM image obtained on short transverse surface

showing precipitate that are rich with Ce aligned in the rolling direction.

Fig. 4. Point EDX analysis on precipitate of Mg–Mn alloy.



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3.3. Constant load test

On the basis of the P Q obtained in K QC tests. CT specimens were loaded. Similarly, maximum load employed during K ICtests on CNT specimens provided idea for loads for CNT tests in corrosive solutions. This practice is based on the fact that

K ISCC/K QSCC is generally lower than K IC/K QC. K I , crack velocity, and validity check for plain strain condition were calculated

for each specimen. The results are presented in Figs. 7 and 8. As expected the time-to-fail in both the tests increased con-

siderably as the initial applied K I approached to the threshold for stress corrosion cracking.

The equations of the curves that best fit data of initial K I against time to fail for both CT and CNT (shown in Fig. 7) are

respectively given below:

Fig. 5. SEM/BSE image of precipitate aligned in the rolling direction and corresponding X-ray mapping for different elements of Mg–Mn alloy.

Fig. 6. Corrosion behaviour, as revealed by hydrogen evolution measurements for Mg–Mn alloy in 0.1 N and 3.5 wt.% NaCl solution.



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K I ¼ 10et f =28 þ 13 ð5Þ

K I ¼ 5et f =400 þ 10 ð6Þ

In both the equations, K I and t f corresponds to stress intensity (in MPa m1/2), and time to fail (in h) respectively. K QSCC and

K ISCC obtained from the above equations are 13 MPa m1/2 and 10 MPa m1/2. It is further noted that at the time to failure equal

to zero, K I equals to 23 MPa m1/2 and 15 MPa m1/2 which are respectively the plane stress fracture toughness, K QC (for 1/2 T

CT testing) and the plane strain fracture toughness, K IC (for CNT testing).

Crack growth data for the alloy tested using CT and CNT specimens are shown in Fig. 8; where straight line having a slope

(m) and an intercept with Y axis is fitted to the data points in stage I to obtain the equations corresponding to the CT and CNTspecimens (as shown in Eqs. (7) and (8) respectively). The best line fit for the data points generated by calculation using Eq.

(4) for CNT specimen is expressed in Eq. (9).

log da

dt

¼ K I 23 ð7Þ

log da

dt

¼ 0:5K I 15 ð8Þ

log da

dt

¼ 0:3K I 13 ð9Þ

where da/dt is in ms1 and K I in MPa m1/2.

Fig. 7. Plots of applied K I vs. time to fail using 1/2 T CT and CNT specimen.

Fig. 8. Plot of log crack velocity vs. applied K I for the test data of CT and CNT specimens (along with calculated data points for CNT using Eq. (4)).



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3.4. Fractography

Representative stereo fractographs of a failed CT and CNT specimens are shown in Fig. 9. Various regions, viz., machined

notch region, fatigue pre-cracked region, the SCC region, and mechanically failed region are delineated on the fractograph.

White arrows (Fig. 9i) mark the deformation of the specimen along the thickness suggesting that the available thickness did

not satisfy the plane strain conditions, again supporting our earlier suggestion that CT specimen does not meet the plane

strain criterion. Similar deformation along the thickness direction was reported by Winzer et al. [27] for CT specimen of

AZ31alloy tested in distiled water using constant extension rate test (2lm h1) their specimens also did not meet the thick-

ness requirement for plane strain condition. Machined notch and fatigue pre-cracked regions of CNT specimen (Fig. 9ii), were

covered with oxides which was cleaned using boiling solution of 20 wt.% chromic acid, before SEM fractography. However,

before such cleaning, the crack size was measured using travelling microscope.

The overall SEM fractographs of CT and CNT specimens that failed in CLT were obtained after cleaning in boiling solution

of 20 wt.% chromic acid. The SCC region is distinct for its appearance of fingers extending from fatigue region in the direction

of rolling until mechanical failed region (Fig. 10a). The fatigue region of CT specimen is shown in Fig. 10b which shows stri-

ations with evidence of precipitates that were undermined because of long contact with the corrosive environment during

CLT. Overall fracture surface of a CNT specimen is shown in Fig. 11a; whereas the transition from fatigue to SCC is shown in

Fig. 11b. TGSCC region for CT specimen is shown in Fig. 10c, and, at higher magnification, the precipitates that are aligned

in rolling direction are shown in Detail A of Fig. 10. EDX analysis of these precipitates shows similar contents as that shown

in Fig. 4. SCC region showing parallel facets (Fig. 11c) are the evidence of hydrogen assisted cracking and discontinuous crack

growth [28–30]. Mechanical failed region posses fine dimples as an evidence of ductility (Fig. 11d).

4. Discussion

In Fig. 7, for time to failure equal to zero, K I for CT specimen corresponds to 23 MPa m1/2 which is plane stress fracture

toughness K QC, and for CNT K I corresponds to 15 MPa m1/2 (i.e., plane strain fracture toughness K IC). The specimens for which

the tests were terminated after 1008 h are identified by arrows, these specimens were forced to fracture by tensile loading to

enable crack size measurement and fractographic analysis.

The applied K I has a considerable influence on the crack propagation velocity; and there is considerable difference in CT

and CNT data, as shown in Fig. 8. The data generated using CT specimen showed both stage I and stage II crack growth (which

is similar to those in the schematic represented in Fig. 12 by Dietzel [22]). Notably, Speidel et al. [31] have shown ZK60 alloy

to exhibit two stage crack growth behaviour in distiled water, 1.4 M Na2SO4 and 5 M NaBr when tested using double can-

tilever beam specimen. On the other hand, CNT did not show stage II crack growth. It is useful to examine previous results

published in literature where in CNT specimens were employed to study SCC behaviour of various alloy systems. Notably,

steel [17,18], and magnesium alloy [32] have not been shown to exhibit stage II crack growth. It is not clear if the authorshad planned to examine this aspect in their work or it is parse the short come of CNT specimen it does not show stage II crack

growth. However, in the present case we were not able to obtain stage II crack growth. This is because of the fact that the

CNT specimen failed when it was loaded near K IC (14 MPa m1/2) and below this value, no stage II crack growth tendency was

noticed. On the contrary, stage II crack growth is set in CT specimen data well below the alloy K QC value. Hence, it can be said

that CNT test specimen may lack the ability to show stage II crack growth region. It is not clear, what could be the plausible

reason for the absence of stage II crack growth in CNT specimen.

It is also noticed from Fig. 8 that CNT and CT specimens show two different slopes 0.5 and 1 respectively in stage I crack

growth region. Not only the slopes of stage I crack growth between CNT and CT specimens are different, but also the

crack growth velocity. The latter can be seen by the fact that for a K I value of 13 MPa m1/2 CT and CNT specimens show crack

growth rates of 1 1010 and 3.1 109 ms1 respectively.

Fig. 9. Optical stereo fractograph of the fractured surface: (i) CT, and (ii) CNT specimens marked with different regions; (a) machined notch, (b) fatigue precrack, (c) SCC region, and (d) mechanical failed region (The arrows show deformation in the thickness direction of CT specimen).



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Fig. 10. Fractographs of failed CT specimen: (a) SEM fractograph showing finger like appearance extending from fatigue crack front in rolling direction

(black ink spot shows stress corrosion crack front), (b) fatigue region showing striations, (c) TGSCC region showing evidence of precipitates that had mine

out which are mostly aligned in the rolling direction (Detail A shown above arrow in the BSE/SEM image).

Fig. 11. Fractographs of failed CNT specimen: (a) SEM fractograph showing finger like appearance extending from crack front in rolling direction, (b) SEM

fractograph shows transition from fatigue to SCC, (c) SCC region showing parallel facets and evidence of precipitates that had mine out shown by feintarrows, and (d) evidence of ductility in mechanical failed region.



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is less severe than 0.1 N NaCl solution saturated with Mg(OH)2 employed in the present test. The table shows that ZK60A-T5

exhibits K ISCC as 8 and 6 MPa m1/2 in 1.4 M Na2SO4 and 5 M NaBr respectively even at higher crack velocity (1 10

7 ms1)

[31]. So this alloy (ZK60A) seems to be more susceptible to SCC should the environment contains aggressive species such as

sulphate and bromide. Hence, it is suggested that the alloy having composition Mg–0.1Al–1.75Mn–(0.3 other RE)–0.3Zn–

0.1Si–0.25Ce can exhibit higher K ISCC in reasonably identical environment.

5. Conclusions

Stress corrosion crack growth, and stage I cracking behaviour of rare-earth containing hot rolled Mg–Mn alloy in 0.1 N

NaCl solution saturated with Mg(OH)2 were studied using CT specimens of 12.2 mm thickness and CNT specimens of

9.5 mm diameter under constant load conditions. To the authors knowledge this is the first study which reports the crack

growth behaviour of any Mg–Mn alloy. The main conclusions of this study are as follows.

(1) Circumferential notch tensile (CNT) technique was successfully used to determine the threshold stress intensity for

stress corrosion cracking under plane strain condition (in spite of the insufficient thickness of the plate of the test

material). The K IC and K ISCC are found to be15 MPa m1/2 and 10 MPa m1/2 respectively and the alloy exhibited a slope

of 0.5 during stage I stress corrosion crack growth.

(2) The K QC and K QSCC determined using CT specimens (that could only meet plane stress criterion) were found to be

23 MPa m1/2 and 13MPa m1/2 respectively, the alloy exhibited a slope of 1 during stage I stress corrosion crack growth.

(3) The SCC cracks propagated in transgranular manner with the evidence of flat parallel facets which indicate hydrogen

assisted cracking.

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Table 2

Comparison of K ISCC data.

Material Specimen type and environment K ISCC (MPa m1/2) V c (ms

1) Ref.

AZ91D sand cast CNT in modified body fluid 5.5 1 1014 [32]

AZ91E-T6 sand cast SENB in 3.5% NaCl 7.5 – [35]

ZK60A-T5 extruded DCB in 5 M NaBr 6 1 107 [31]

ZK60A-T5 extruded DCB in 1.4 M Na2SO4 8 1 107 [31]

ZK60A-T5 extruded DCB in distiled water 12 1 1010 [31]

AZ31 extruded CT in distiled water 13.7a

2.5 109

[27]Mg–9Al extruded, ST 4 h at 350C, water quenched SENB in 2% NaCl + 0.5–2% K2CrO4 7

a–14a 0.1–5 106 [34]

AZ61 rolled Single edge notched in 3.5% NaCl + 2% K2CrO4 14a–17a – [36]

AZ61 rolled, 1.5 h at 475 C, quenched, aged 48 h at 200 C Single edge notched in 3.5% NaCl + 2% K2CrO4 12a–3a – [36]

Mg–Mn hot rolled (present work) CNT in 0.1 N NaCl sat. Mg(OH)2 10 1 1010 –

a These values do not satisfy plane strain conditions.



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2008;10(5):453–8.

[28] Chakrapani DG, Pugh EN. Hydrogen embrittlement in a Mg–Al alloy. Metall Trans A 1976;7:173–8.

[29] Meletis EI, Hochman RF. Crystallography of stress corrosion cracking in pure magnesium. Natl Assoc Corros Eng 1984;40(1):39–45.

[30] Winzer W, Atrens A, Dietzel W, Song G, Kainer KU. Comparison of the linearly increasing stress test and the constant extension rate test in the

evaluation of transgranular stress corrosion cracking of magnesium. Mater Sci Engng A 2008;472:97–106.

[31] Speidel MO, Blackburn MJ, Beck TR, Feeney JA. Corrosion fatigue: chemistry, mechanics and microstructure. Natl Assoc Corros Eng 1971;2:324–45.

[32] Choudhary L, Raman RKS. Magnesium alloys as body implants: fracture mechanism under dynamic and static loading in physiological environment.

Acta Biomater 2012;8:916–23.

[33] Jones RH, Simonen EP. Stage I corrosion cracking behaviour. Mater Sci Engng A 1993;160:127–36.

[34] Ebtehaj K, Hardie D, Parkins RN. The influence of chloride-chromate solution composition on the stress corrosion cracking of a Mg–Al alloy. Corros Sci

1993;28:811–21.

[35] Stephens RI, Schrader CD, Goodenberger DL, Lease KB, Ogarevic VV, Perov SN. Corrosion fatigue and stress corrosion cracking of AZ91E-T6 castmagnesium alloy in 3.50/0NaC1 solution. SAE 930752. 1993; 843–57.

[36] Moccari A, Shastry CR. An investigation of stress corrosion cracking in MgAZ61 alloy in 3.5% NaCl +2%K2CrO4 aqueous solution at room temperature. J

Mater Tech (Zeitschrift für Werkstofftechnik) 1979; 10: 119–23.


fatgue paper for mn alloys good concept

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