SiC-particulate metal matrix composites

Journal Pre-proof Tension and fatigue behavior of Al-2124A/SiC-particulate metal matrix composites Ji Xia, John J. Lewandowski, Matthew A. Willard PII:

S0921-5093(19)31304-8

DOI:

https://doi.org/10.1016/j.msea.2019.138518

Reference:

MSA 138518

To appear in:

Materials Science & Engineering A

Received Date: 24 June 2019 Revised Date:

1 October 2019

Accepted Date: 5 October 2019

Please cite this article as: J. Xia, J.J. Lewandowski, M.A. Willard, Tension and fatigue behavior of Al-2124A/SiC-particulate metal matrix composites, Materials Science & Engineering A (2019), doi: https://doi.org/10.1016/j.msea.2019.138518. This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2019 Published by Elsevier B.V.

Tension and Fatigue Behavior of Al-2124A/SiC-particulate Metal Matrix Composites Ji Xia, John J. Lewandowski, and Matthew A. Willard Department of Materials Science and Engineering, Case Western Reserve University, Cleveland, OH

1. Abstract The tension and fatigue properties of SupremEX® 225XE composites (Al2124A/25%/SiCp/3 μm) were determined for extruded samples tested in the longitudinal orientation in the T4 condition. Tension testing was conducted at 0.001/sec on as-machined cylindrical samples tested with high alignment fixtures according to ASTM E-8. The fatigue tests were conducted on polished hourglass fatigue samples with stress ratio R = 0.1 and test frequency of 20 Hz using hydraulic grips on a Model 810 MTS servo-hydraulic testing machine according to ASTM E466-15. A total of 26 samples were tested with stress levels between 448 MPa and 557 MPa. The resulting stress vs. cycles to failure (i.e. S-N curve) was compared to the unreinforced matrix alloy, other unreinforced high strength aerospace aluminum alloys, as well as other conventionally processed MMCs. The SupremEX® 225XE MMC exhibited improvements to both the low cycle fatigue (LCF) and high cycle fatigue (HCF) performance compared to these materials. The fatigue strength at 107 cycles was 448 MPa, well in excess of those for monolithic aluminum alloys as well as conventionally processed MMCs. Detailed fractography revealed the presence of a unique cone-shaped fracture feature for samples failing in HCF. Calculations based on fracture morphologies suggested catastrophic fatigue fracture occurred as the material reached its critical fracture toughness. Initial modeling of this fatigue performance used a variant of the Universal Slopes Criterion proposed in early work at Case Western Reserve University by S. M. Manson. Keywords: Al metal matrix composites; SiC reinforcement; tension; fatigue SiCp MMC) is one of the most popular MMCs

2. Introduction

that have been studied. This class of materials matrix

has high specific strength and stiffness

composites (MMCs) are a class of materials

compared to Al alloys and other conventional

that

materials such as steels [1], making them

Particulate-reinforced consist

of

metal

reinforcement

particles

imbedded in a metallic matrix. The SiC

great

candidates

for

automotive

and

particulate reinforced aluminum alloy (Al-

aerospace applications. In practice, MMCs have been applied to both military and 1

commercial aircraft to replace metallic alloys

dislocation substructure. The reinforcement

and plastic composites [2]. In the field of

particles, depending on their size, can also act

transportation, these materials are also used

as obstacles during deformation of the

in engine parts [3] and railway vehicles [4].

composite and introduce a back stress during deformation.

The major strengthening contributor is the reinforcement

the

The overall improvement in composite

strengthening mechanism can be divided into

strength and stiffness is a superposition of all

direct and indirect strengthening methods [5].

the

Direct strengthening is related to load

contribution from each mechanism is also

transfer from the matrix to the reinforcement.

dependent on various factors, such as

This can provide significant enhancement to

reinforcement

the yield strength of the composite as a

volume fraction [5]. Studies have shown that

function of reinforcement volume fraction [6],

the tensile strength of the Al MMCs can be

which is described by the modified shear lag

improved by around 50 MPa with either 10%

model:

increase in reinforcement volume fraction or =

where

and

particles,

where

+2 +

(1)

aforementioned

size,

mechanisms.

and

The

reinforcement

several microns of decrease in reinforcement size [13]–[17], although this depends on the

are the yield strength for

the composite and the matrix, and

matrix

and

alloy,

heat

treatment,

and

reinforcement size and volume fraction.

are the volume fractions for particulate and

While these strength improvements are

matrix, respectively. The aspect ratio ( ) for a

positive, some of the greatest mechanical

spherical particle is taken as 1.

improvements relate to enhanced high cycle fatigue performance, as reviewed below.

On the other hand, various indirect strengthening mechanisms can contribute to

Hall et al. [18] evaluated the effects of

various changes in the matrix microstructure

reinforcement size on fatigue performance of

as a result of the addition of reinforcement

MMCs containing a 2124A aluminum matrix

[7]. Due to the thermal expansion mismatch

reinforced with 20 vol.% SiC particulates. The

between the Al matrix and SiC particles [8],

average reinforcement size was varied from

residual stresses can remain in the composite

2 μm to 9 μm to 20 μm, while fatigue testing

after

large

was conducted at R = -1, and 30 Hz. The

population of dislocations that can be

composite with 2 μm SiC particles exhibited

generated around the reinforcement particles

the best fatigue life in both low cycle fatigue

[6], [9], [10]. This can produce accelerated

(LCF) and high cycle fatigue (HCF), followed

aging during heat treatment [11], [12], as well

by the one with 9 μm SiC particles. The

as higher initial work hardening due to the

composite

processing,

resulting

in

a

2

with

20 μm

reinforcement

exhibited the worst fatigue performance and

225XE

this was attributed to premature fracture of

(Al2124A/25%/SiCp/3 μm).

the large SiC particles, thereby producing

were produced via a novel powder processing

early fatigue initiation and enhanced damage

technique developed by Materion Brush Inc.

accumulation during cyclic loading. Similar

with subsequent consolidation by HIP and

observations were made by Chawla et al. [14]

extrusion and tested in the T4 condition.

Al-SiCp

MMC These

MMCs

on a 2080 Al alloy reinforced with 20 vol.%

3. Materials

SiCp and tested at R = -1 and 30 Hz, where the fatigue lifetimes were improved in both the

The materials were processed by Materion

LCF and HCF regimes with a decrease in

Brush Incorporated, with the brand name of

particle size. Chen and Tokaji [19] also found

SupremEX® 225XE and processed using a

that 2024Al-SiCp MMCs containing 5 μm SiC

novel powder premixing route. High quality

exhibited fewer fractured SiCp compared to

grade 2124A aluminum alloy powders were

MMCs containing either 20 μm or 60 μm SiCp

premixed and mechanically alloyed with SiC

reinforcements. This suggests that smaller

particulate reinforcement using a proprietary

reinforcement size provides better fatigue

process to create agglomerate composite

performance.

powders [21]. The 2124A aluminum alloy, The effects of changes in volume fraction

with composition shown in Table 1 [22], was

SiCp from 0 to 30 vol.% with constant SiCp

reinforced with 25 volume percent SiC

size was also examined in the work presented

particulate, with an average size (d50) of 3 μm,

by Chawla et al. [14], which showed a

and were initially consolidated into billets by

continuous improvement in the composite

hot isostatic pressing (HIP) using cans that

fatigue strength with increasing volume

were 20.32 cm (8 in.) in diameter and

fraction SiCp, and well in excess of the

68.58 cm (27 in.) in length. Argon gas at

unreinforced 2080 Al matrix alloy. For

103 MPa pressure was used along with HIP

instance, fatigue strength at

107

cycles for the

temperatures ranging from 420 to 520°C.

composite with 30 vol.% reinforcement was

Extrusion was then performed at a ratio of

250 MPa, whereas the value for 2080 Al alloy

30:1 at 430°C with extruded bars heat treated

was only 140 MPa. Improvements in the

to the T4 condition (i.e. solution treated at

fatigue life were also observed in Kaynak and

505°C and then cold water quenched). All

Boylu’s work [20] on a SiCp reinforced 4xxx

tension and fatigue specimens were taken

series Al MMC with reinforcement content

from the middle of the extruded bars.

increasing from 0 to 15 SiCp volume percent. The The present work investigates the tension

SupremEX®

designated

and fatigue performance of SupremEX 3

225XE

material

2124A/SiC/25p (3 μm),

is and

shortened

as

remainder

of

225XE

the

automated polishing system to examine its

naming

microstructure in different sections using

convention follows the format of [Geometry] -

standard metallographic procedures. The

[Material] - [Heat treatment] - [Extrusion

polishing sequence was 220 grit SiC polish,

direction] - [Test ID]. For example, R-225XE-

550 grit SiC polish, 9 μm diamond polish,

T4-L1 refers to one cylindrical (round) 225XE

3 μm diamond polish, and OP-U colloidal

specimen with T4 heat treatment, tested in

silica suspension polish with each step

the longitudinal direction, labeled as L1. The

duration ranging from 60 to 120 seconds. The

naming convention used for the fatigue

specimen was then cleaned in a Struers

specimens is similar to that for tension

Lavamin ultrasonic cleaning chamber using

specimens, with “H” standing for hourglass

tap water between each step for a 30-second

fatigue geometry.

rinse followed by a 30-second ultrasonic

the

throughout

paper.

This

cleaning.

Triplicate cylindrical tension specimens were prepared in the longitudinal orientation and

with

T4

heat

treatment.

After these preparations were completed,

Sample

the microstructures in different orientations

dimensions followed ASTM E8 [23] with a

were examined with an FEI Nova Nanolab

gage length and diameter of 24.9 mm and

200 SEM using secondary electron mode, at

5.1 mm, respectively. Twenty-six hourglass

magnifications of 1000x, 2500x, and 5000x.

fatigue specimens were machined from the

The operating voltage was 15 kV, and the

longitudinal orientation and polished by

current was 1 nA.

Materion Brush Incorporated following ASTM

4.2 Tension tests

E466-15 [24]. The hourglass gage length was 35.2 mm with a gage radius of 50.8 mm and

The uniaxial tensile tests were performed

minimum diameter of 6.4 mm. Final polishing

using a high alignment grip on an Instron

of the gage region was conducted with

1125 Universal Testing

micron-sized polishing compound along the

ambient conditions with a 100 kN load cell. A

longitudinal axis of the sample to minimize

standard MTS contact extensometer was used

any transverse scratches.

to record sample displacement data (Δl), with

Machine under

a gage length of 12.7 mm (0.5 in). The

4. Experiments

extensometer was calibrated prior to each

4.1 Microstructure Evaluation

and was removed from test samples after

test session, using an analog calibration tool,

One grip end of a broken tension sample

about 1% strain to prevent damage to the

(R-225XE-T4-L3) was sectioned, mounted in

extensometer. An UVID Arion 1DTM non-

epoxy, and polished on a Struers Tegramin

contact video extensometer was also used to 4

record

strain

data

simultaneously extensometer.

up

with This

to the

%= 1−

failure, contact

non-contact

where

video

and

extensometer was utilized to capture strain measurements

after

the

#⁄

(5)

is the load, Δ is the displacement, are the initial gage length and cross

sectional area, and

contact

$ × 100%

is the area of the

#

fracture surface. Elastic modulus (') and 0.2%

extensometer was removed. It operates by

offset yield strength (0.2%

recording high-resolution measurement of

) were also

calculated from the engineering stress-strain

spacing changes between three fiducial

curves, according to ASTM E111 [26].

markers painted on the sample surface, along the loading direction. These fiducial markers

4.3 Fatigue tests

were painted using an orange oil-based

Twenty-six

uniaxial

stress−controlled

applicator pen, with marker size on the order

fatigue tests were conducted at R = -0.1 and

of 0.5 mm in diameter to be recognized by the

20 Hz using a MTS 810 Material Test System,

camera. The gage length of this non-contact

and followed ASTM E466-15 [24]. The

extensometer refers to the spacing between

maximum load

the outermost markers, which was 12.7 mm.

determined by:

The data collected from this non-contact extensometer is aligned with that from the contact extensometer with reference to time.

()

Matlab codes were created to analyze and align

the

data

collected

from

where

both

and

extensometers, as documented in related

for each test was

=

⁄

= * ,

() ⁄4

(7)

was the desired maximum stress, ()

was the cross-sectional area at the

minimum diameter, ,

work [25].

(6)

()

() ,

of the hourglass

specimen, which was measured prior to each

All tension tests followed ASTM E8

test using a micro-projector. Apart from

standard [23]. They were conducted at an initial strain rate of

10-3/s, ,

the number of cycles to failure (.# ) was also

with a sampling

rate of 20 Hz. Engineering stress ultimate

recorded for each experiment to produce the

,

stress-cycles-to-failure curve, or S-N curve.

tensile

Normalized S−N curves against average UTS

engineering

strain

strength (

), and reduction of area (

%)

and 0.2%

were calculated based on following equations: (2)

=Δ⁄

(3) ⁄

were plotted to compare to

other conventionally processed MMCs.

= ⁄

=

,

(4)

5

4.4 Fractography Fractography was performed using both optical and scanning electron microscopes (SEM). A Keyence VHX-5000 series digital

microscope was used to characterize the

on the images in Figure 1. In the longitudinal

fracture surfaces at magnifications between

direction,

20x and 200x. It was also used for

homogenously inside the matrix. However, in

approximate 2D measurements, such as

the transverse direction, bands of SiC

diameters of fracture surfaces. SEM imaging

deficient regions exist as a result of extrusion.

was performed using either an FEI Quanta 3D

This difference in microstructure (i.e. SiC-rich

Environmental SEM (for specimens taller

and lean regions) contributes to orientation-

than 15 mm), or an FEI Nova Nanolab SEM

dependent tensile properties as shown in a

(specimen shorter than 15 mm). Secondary

related work [27].

electron

5.2 Tension results results

imaging

voltage and

parameters

current were

including

adjusted to

In

terms

of

fracture

particulates

distribute

The calculated tensile properties are listed

optimize images for specific features or details.

SiC

in Table 2, and an example of a typical stress-

surface

strain curve is shown in Figure 2. The values

characterizations, the voltage was chosen as

were calculated from

15 kV or 20 kV, with current chosen as 3 nA.

of E, and 0.2%

Positions of fracture initiation were identified

corresponding

and imaged. A unique cone-shaped fracture

curves. Calculations of elongation, UTS, and

feature exhibited in certain HCF fatigue tests

ROA% followed Equation 4, 5, and 6. The

was further examined in the region of fatigue

average value /0 and standard deviation (SD)

crack growth (i.e. the side of the cone) and

for each property are calculated based on the

catastrophic failure (i.e. outside of the cone).

three measured specimens.

engineering

stress-strain

As a supplement to the FEI Nova Nanolab

5.3 Fatigue results

SEM, energy dispersive X-ray spectroscopy (EDS) was used at 20 kV to perform

The S-N curve generated based on σmax and

elemental analysis of fracture initiation

Nf, is shown in Figure 3. In the plot, the

features of the specimen.

average 0.2%

listed in Table 2 is indicated,

and the fatigue strength at 107 cycles is

5. Results

shown as 448 MPa. The S-N curve was

5.1 Microstructure

partitioned into a low cycle fatigue (LCF) and a high cycle fatigue regions (HCF) with

Figure 1 shows the microstructure of R-

conventionally defined separation at 105

225XE-T4-L3 in the longitudinal (L) and

cycles [28]. As can be seen, most of the fatigue

transverse (T) directions. No attempt was made

to

determine

the

particle

tests were conducted at

size

stresses that

exceeded the yield strength of the material. In

distribution, however, the reported average

addition, fracture initiation sites for each

particle size (3 μm) seems reasonable based 6

specimen were identified by SEM or Keyence

exhibiting

optical microscope as occurring from the

revealed a unique cone-shaped fracture

sample edge or inside the sample. Samples

feature on some of the HCF fatigue samples,

exhibiting fracture from internal defects

such as the one shown in Figure 7. In Figure 7,

typically exhibited a unique cone-shaped

the apex of the cone points towards the

fracture morphology, as discussed in section

viewer and contained one of the Fe-Cr rich

6.4. These specimens are marked with

inclusions.

triangles in Figure 3.

typically observed at the cone apex of such

5.4 Fractography

samples, as summarized in Table 3. The base

internal

Fe-Cr

fracture

rich

nucleation

inclusions

were

of the cone is observed at the far right side of A typical tensile fracture surface shown in

Figure 7 as a change in content.

Figure 4 exhibited dimpled fracture with void Keyence optical microscopy was used to

growth around fractured and/or decohered

determine the macroscopic cone morphology

SiC particulates. Shear lips marked by red

as shown in Figure 8. The apex of each cone

arrows in Figure 4 were also present on the

roughly formed a 90-degree angle while

cylindrical tension sample fracture surfaces.

Table 4 documents the cone apex angle and As shown by the S-N curve in Figure 3,

maximum cone diameter (i.e. at its base).

fatigue fracture initiated from either edge

Figure 9 provides an SEM image of the side

nucleated or internal defects. Figure 5 shows

surface of the cone in addition to an image

an example of edge initiated fracture due to

showing the typical catastrophic overload

surface inhomogeneity. In this case, a defect

failure that occurred beyond the base of each

beneath the polished surface layer initiated

cone. Interpretations of these images and

the crack. Step-like features that radiated

features are provided in the discussion.

from fracture initiation sites were observed in such specimens. On the other hand, Figure

6. Discussion

6 shows an example of fracture initiation from an internal defect that was identified by

6.1 Mechanical properties compared to

EDS as a large (i.e. 30 µm) SiC particle. The

monolithic 2124 Al alloy

majority of internal initiation sites were

Table 5 summarizes the R-225XE-T4-L

associated with Fe-Cr rich inclusions (as

tensile

determined by EDS) on the order of 40 µm,

monolithic Al 2124 [22]. The improved elastic

shown in Figure 7.

modulus as a result of the addition of

analyses

of

comparison

to

(ESiC ~ 400 GPa [29]) has been captured by

fracture initiation (i.e. edge vs. internal), surface

in

homogenously distributed SiC reinforcement

In addition to documenting the location of fracture

properties

using the Hashin-Shtrikman relationship in

samples 7

previous work [25]. For this study, the upper

where

and

fraction, and r = 1.5 µm is the average particle

lower

bound

at

25

vol.%

SiC

= 0.25 is the reinforcement volume

reinforcement were calculated from Equation

radius.

1 as 130 GPa and 105 GPa, respectively.

spacing is on the order of 6 µm. While the SiC

Extensive testing of elastic modulus (E) by

particles of this size are too large to produce

Materion [21] has shown E values greater

Orowan strengthening [30], they will act as

than 110 GPa suggesting some slippage of the

barriers to dislocation movement and create

extensometer in the present test.

a backstress during tension testing as well as

The calculated

SiC

interparticle

constrained flow of the surrounding matrix

Improvements to both the 0.2% offset

[31], [32]. These factors will contribute to an

yield strength and UTS were also exhibited in

additional increase in strength and work

the 225XE composite. Improvements to the

hardening behavior in the absence of damage

yield stress were captured by using a

(e.g. voids or cracked SiC particles) that may

modified shear lag model described by

evolve

Equation 2, as summarized elsewhere [25].

during tensile deformation. The

combination of these factors also contribute

The calculated yield strength from the

to the increase in the UTS, and that is

modified shear lag model was 496 MPa, with

consistent with present observations [25].

an assumption of spherical reinforcement. However, this model does not completely

However, in addition to the presence of the

capture the strengthening observed presently,

brittle

and it typically under-predicts the yield

constrained flow of the matrix, their presence

strength based on other experimental results

will also increase the rate of damage

[25]. As outlined earlier, other sources of

accumulation during tension testing [33] at

strengthening include potential contributions

ambient pressure and produce lower strains

from

density

to failure, consistent with the reduced

thermal

elongation in comparison to monolithic 2124.

expansion coefficients (i.e. a factor of 10)

These effects have been clearly documented

between the SiC particulate and Al matrix, as

in other work and further show the important

well as accelerated aging of the matrix

effects of imposed stress state on the damage

facilitated by the high dislocation density.

accumulation rate as well as resulting tensile

Additional contributions to strength result

ductility [31], [32], [34]. However, the

from the barriers to dislocation movement

relatively uniform SiCp dispersion and small

provided by the SiC particles. Assuming

SiCp size in the present MMC appear to delay

spherical SiC reinforcement, the interparticle

the

spacing λ can be estimated as:

accumulation, thereby producing sufficient

the

resulting

1=

2

enhanced from

345 $6

7 45

dislocation

differences

in

SiC

onset

particles

of

damage

increasing

initiation

the

and

work hardening and higher ductility than

(8) 8

typical DRAs [33] in addition to exhibiting

shown in Figure 3 to an adjusted S-N curve

necking prior to catastrophic failure [25]. In

that represents the actual (i.e. local) stress

the present work, this manifests itself as

present at the fracture initiation site. The

values for the true fracture stress that exceed

values for

the UTS, as summarized in Table 5.

using D and the geometry information from

6.2 Adjustment of SS-N curves due due to fracture

each of the 26 samples tested:

89

position Analysis

89

conducted on the H-225XE-T4-L materials.

= ⁄ * :89

8$

(9)

+ ;: − <: − = >

8

=:

where :89

8

is the radius of the local plane,

:

The values for stress shown in Figure 3

8

were calculated as follows

:89

Figure 3 showed the S-N curve obtained from the 26 hourglass fatigue experiments

8

()

()

(10)

is the radius of the middle plane, and :

is the radius of the hourglass geometry, which

utilized the maximum stress (i.e. peak

is 50.8 mm. Figure 10 shows the adjusted S-N

load/minimum cross sectional area) applied

curve using local stresses and includes the

to the hourglass samples tested presently.

data from Figure 3 which used the maximum

While fatigue fracture is typically expected to

stress applied at the minimum diameter. This

initiate and grow from sites at the maximum

approach shifts the S-N curve to lower

stress location (i.e. minimum diameter) in

stresses for the majority of data points while

hourglass samples, the present results clearly

also reducing the degree of scatter in fatigue

showed that most fracture initiation sites

lifetime is also reduced at certain stresses (e.g.

deviated from this location.

500 MPa).

For

the

remainder

of

the

The deviation of fatigue fracture initiation

discussion, only the adjusted S-N curve will

sites, D, from the sample minimum diameter

be used as this more accurately represents

were calculated by measuring the distance, L,

the S-N behavior since it uses the actual

from the fatigue initiation site to one end of

stresses at the fatigue initiation sites.

the gage with a straightedge, and then

6.3 The Universal Slopes Analysis

subtracting that from the half gage length, as shown in Figure 10. In addition to quantifying

One of the most important applications of

the distance from the minimum diameter of

S-N curves is to predict a material’s lifetime at

the sample, this enables calculation of the

a given stress level. While a number of

actual (i.e. local) stress present at that

models have been developed in this regard

location. Thus, the same far-field load

[35], [36], the test conditions in the present

produces a local stress

work (i.e. R=0.1) do not lend themselves to

89

8

that is smaller

than the maximum applied stress

. This

using

enables adjustment of the original S-N curve

the stress-based approaches or the

Coffin-Manson

9

and

Basquin

equations.

However, the Universal Slopes Equation [28]

performance, and was obtained by strain-

in Equation 12 provides an approach that can

controlled fatigue tests with strain ratio R=0.

be used because it has no requirements on

Fortunately, this is close to our

test conditions:

conditions where the strain ratio ranged from

Δ = ?@ .#A + ? .#B

(11)

where Δ is the total strain range, ?@ is the

test

only 0.02 to 0.08. Despite these subtle differences, Figure 11 shows that the data for AMC225 extends nicely the high cycle fatigue

strain at the elastic limit, ? is the plastic

data presently obtained for 225XE, and

strain up to the UTS, and C and D are material

expands the data population into the LCF

constants. The intercept of this equation on

regime. ?@ and ? are derived as average

the y-axis corresponds to the strain range of a

values

tensile test for the same material (i.e.

0.00322 and ? = 0.0854, respectively. A

from

both

studies,

where ?@ =

.# = 1⁄2). In practice, the value of ? and ?@

power-law fit following Equation 12 was

can be approximated from the engineering

applied to the data points with fatigue life up

stress-strain curves obtained for tension tests.

to 107 cycles. Figure 11 shows the adjusted S-

In this case, the total strain range is equal to

N curve with the Universal Slopes Equation fit

strain. ?@ and ?

be

curve. The fit in general follows the trend of

approximately measured based on their

the data with R2 = 0.81 for the following

definition. It is assumed that the stress

equation:

the

fracture

can

gradient caused by the hourglass geometry is

Δ = 0.0854 .#3

neglected, so that the total strain range

. J

+ 0.00322 .#3

.

K L

remains the same across the gage. The runout

The calculated strain range at Nf =1/2 cycle

specimen was excluded.

(i.e. tension test) produces 0.102, which is close to the true fracture strain calculated

The adjusted S-N curve was first recreated

from ROA (i.e. 0.072) shown in Table 2.

to plot the total strain range (i.e. at the fatigue

6.4 Unique ConeCone-shaped Fracture Features Features

initiation site) against cycles to failure, shown in Figure 11. In addition to the present test

The unique cone-shaped features appear

results, Figure 11 also includes data for an

to

earlier generation of Al-SiCp MMC (i.e.

originate

due

to

fracture

initiation

occurring at Fe-Cr rich inclusions that are not

AMC225) very similar to SupremEX® 225XE,

located at the minimum diameter of the

from the literature [37]. While, the AMC225

hourglass

MMC possessed the same metallic matrix,

sample,

thereby

requiring

subsequent fracture to propagate at roughly

reinforcement volume fraction, and average

45

reinforcement size as tested presently, the

degrees

to

the

tensile

axis

until

catastrophe intervenes. While the fracture

fatigue data for AMC225 focused on the LCF

initiation sites in these samples were 10

(12)

typically at Fe-Cr inclusions, subsequent

= ⁄4ST

fatigue crack propagation along the cone

where P = 0.637, which is a sample geometry

surfaces was predominantly by shear with

related coefficient; R = :, which is the flaw

little evidence of SiCp fracture along the cone surface.

However,

eventually

each

terminated

of

with

the

size; is the nominal stress;

cones

of a rectangular plane perpendicular to loading direction, respectively. In practice,

tensile axis, suggesting that either the UTS or

the surface area of the rectangular plane was

the fracture toughness of the MMC was

approximated by a circular plane with radius

exceeded. In order to identify which of these

of R0. Results of these calculations shown in

fracture criteria controls failure, quantitative

Table 6 indicate that values of

measurements were conducted on the cone

UTS from tension tests (i.e. 617 MPa). This

. A projected

# and

far-field stress

strongly suggests that reaching the UTS in the

#

remaining

at fracture initiation site was calculated using: #

=* #

= ⁄

of

did

not

produce

−:

(13)

#

(14) that causes catastrophic fracture can also be

fatigue samples,

toughness of the material. As shown in Table

a stress

6, the calculated NO values are rather consistent, despite the very different applied

according to [38], and used to calculate the local maximum stress

far-field stresses used and final flaw sizes

at catastrophic

observed. The critical fracture toughness

fracture using: = CM ⁄

Rather, the presence of an embedded flaw examined via calculation of the fracture

concentration factor CM = 1.03 was calculated

value NUV for #

shown

2124A/20

wt.%/SiCp

in

ASM

Engineered

Materials

Handbook (Vol. 1, 1987), and much previous

and the known defect

work reviewed by Hassan et al [33]. Although

dimensions (i.e. maximum cone diameter) by

225XE has a different composition than 2124

approximating the base of the cone as an

Al, the calculated NO approaches the known

embedded penny-shaped flaw in the material,

critical fracture toughness values reported by

following the instructions shown in [39], NO = P √*R

a

(15) composite ranges from 17 to 19 MPa√m,

Estimates of the fracture toughness NO were calculated using

ligament

catastrophic fracture.

Also taking into consideration the hourglass geometry

for all

specimens were smaller than the averaged

structure, including cone radius r and fracture surface area

is the nominal

load; S and T are half of the width and length

catastrophic

fracture that was roughly at 90 degrees to the

fracture surface radius

(17)

many investigators on MMCs [33], [40]. And (16) while these calculations do not provide valid NUV , these values approach the proper 11

toughness NUV measurements of

fracture

particles increases the stiffness of the

these materials reported by Materion [41].

material compared to monolithic Al alloy while the small reinforcement size delays

In light of the above calculations, it is very

crack initiation under cyclic loading in

likely that the cone-shaped features ceased

comparison

propagating in shear when the material

to

MMCs

with

coarser

particulates. As a result of that, the composite

reached its fracture toughness NUV , producing

will be able to support a larger stress than the

catastrophic failure in a nominally Mode I

monolithic Al alloy, if the same amount of

manner. Standard ASTM experiments for

strain is allowed. In other words, if the same

NUV should be conducted to confirm this in the

stress is applied, the composite tends to

future.

deform with smaller strain amplitude than

6.5 Comparison of Fatigue Performance to

the Al alloy. Thereby, the composite with

Conventional MMCs and Monolithic Alloys

smaller reinforcement particles becomes

It is useful to compare the fatigue

more resistant to crack initiation and crack

performance of SupremEX-225XE-T4-L to

growth than the monolithic Al alloy, reflected

both 2xxx and 7xxx monolithic materials as

by its improved fatigue life under the same

well as conventionally processed MMCs by

stress condition. The effect of load sharing is

powder metallurgy. Detailed information

also significant when the applied stress is low.

regarding these materials are listed in Table

It becomes less efficient when the applied

7.

Figure

12

compares

the

stress is high and significant yielding occurs.

adjusted

SupremEX-225XE-T4-L S-N curve from Figure

Figure 13 provides a similar comparison to

10 to both 2124-T851 [40] and 7055-T7751

conventionally processed Al-SiCp MMCs, such

[42] Al alloys. All materials were tested with

as a 2xxx series Al MMC [43] and a 7034 Al

R = 0.1. It can be seen from Figure 12 that the

MMC [44]. Fatigue tests for all these materials

fatigue behavior of 225XE outperforms that

were conducted with stress ratio R = 0.1, and

for the monolithic Al alloys in both LCF and

a range of test frequencies, although changing

HCF. The fatigue improvement in HCF is much

the test frequency is not expected to

more profound than in LCF. The fatigue

significantly affect the results. The increased

strength is increased from 243 MPa for

fracture strain of 225XE that arises due to the

monolithic 2124 Al alloy to 448 MPa for the

finer particulate and more limited damage

composite.

accumulation improves the tensile ductility that

and LCF behavior in comparison to other

contribute to these improvements include

MMCs. Also, 225XE experiences the smallest

both direct and indirect strengthening, as

strain amplitude under the same stress level,

have been discussed. The addition of SiC

exhibiting great improvements in HCF as well.

The

strengthening

mechanisms

12

In addition to the comparisons indicated

strain values are lower than those for

above, it is also useful to provide comparisons

2xxx/7xxx

to the conventionally processed MMCs in a

reinforcement

normalized manner. Figure 14 provides the

contributes to the outstanding performance

adjusted 225XE S-N curve normalized against

of 225XE in HCF, outperforming the other

both 0.2% σy and UTS. In the LCF region, it is

two Al MMCs.

known that fatigue performance is dominated

MMCs

due

to

volume

the

fraction.

higher This

In addition to the above reasoning, the

by the material ductility. The single data point

effects of SiC size on damage initiation and

for the 2xxx Al MMC appears to outperform

growth are also important to consider. It is

225XE in the LCF regime when it is

well documented that reinforcement damage

normalized against yield strength, likely as a

increases with an increase in SiC particle size

result of its higher ductility (due to its lower

at a given stress [33], [45]. A reduction in

SiCp volume fraction) and relatively low 0.2%

reinforcement size reduces the amount of SiC

σy. However, when it is normalized against

particles fractured at a given stress while also

UTS, 225XE shows slightly better fatigue

increasing the stress for SiC particle fracture.

performance than the 2xxx Al MMC, even

Since HCF fatigue is dominated by crack

with its higher UTS and slightly lower

initiation, any delay in crack initiation should

elongation. Although 7034 Al MMC has

improve the HCF performance. Related work

outstanding UTS and 0.2% σy, its poor

[27] has shown a lack of change in modulus in

ductility makes it the least competitive

cyclic stress-strain experiments strongly

material in LCF when the curve is normalized

suggesting that minimal damage to the SiC

against the yield strength.

particles occurs during such cycling, in contrast to similar experiments on MMCs

In the HCF region, the beneficial effects of enhanced

with larger SiCp sizes [34]. This will also

modulus) are due to reductions in cyclic

extend the number of cycles to crack

strain amplitude in the higher modulus

initiation in the present MMCs and positively

material [33]. Increased levels of cyclic strain,

affect the HCF lives at a similar stress level.

reinforcement

(and

resulting

due to higher imposed strains, lower modulus,

7. Conclusions

or both, promotes damage along with crack opening and propagation, thereby reducing

Uniaxial

fatigue lifetime. As can be seen from the plot,

tension

tests

and

stress-

controlled uniaxial fatigue tests have been

the maximum stresses in the HCF regime

conducted

from 225XE are close to the 0.2% σy. This

for

SupremEX®

225XE

[2124A/SiC/25p (3 µm)]. Extruded round

means that the initial cyclic loading must

samples (R-225XE-T4-L) in the longitudinal,

induce some plasticity, but the imposed cyclic 13

T4 heat treatment were tested for their

suggested that catastrophic failure of the

tensile

sample occurred when the fracture toughness

samples

properties.

Extruded

(H-225XE-T4-L))

of

hourglass the

same

of fatigue sample was reached due to the

conditions were tested at 20 Hz and R = 0.1

embedded flaw (i.e. cone) and applied far-

for their fatigue behaviors.

field stress.

Comparing 225XE to monolithic 2124 Al

The adjusted S-N curve together with data

averaged

elastic

modulus

from the literature based on ∆ε was modelled

ultimate

tensile

strength

with the Universal Slopes equation up to 107

(UTS=617 MPa), and 0.2% yield strength

cycles. The fit result was expressed as ∆ε =

(0.2% σy=467 MPa) was increased by 34.7%,

0.0854 Nf-0.271 +0.00322 Nf-0.00519, with R2 =

31.6%, and 11.0%, respectively, The elastic

0.81.

alloy,

the

(E=102 GPa),

modulus is close to the lower bound

Compared to another 2xxx Al MMC [43]

predicted by the Hashin-Shtrikman model.

and

The yield strength is lower than the

a

7034

Al

MMC

[44],

225XE

outperformed these conventionally processed

predication from the modified shear lag

composites in both LCF and HCF. Its HCF

model.

strength

exceeded

75%

UTS.

The

Fatigue S-N curves were initially generated

improvements to LCF behaviour were related

based on the maximum stress (σmax) at the

to the higher ductility provided by the smaller

minimum sample diameter. The fatigue

SiCp size reducing damage accumulation and

strength at 107 cycles was 448 MPa, an

increasing the ductility. The improvements to

increase of 205 MPa (83.2%) compared to

HCF were related both the higher elastic

monolithic 2124 with stress ratio R = 0.1 [40].

modulus (i.e. reducing the cyclic strain

The fatigue strength increase is also over

amplitude) as well as the finer SiCp delaying

200 MPa compared to 7055 Al alloy at

fatigue damage initiation.

106

cycles [42].

8. Acknowledgement

It was observed that edge fracture initiation dominated in the low cycle fatigue

The

authors

gratefully

acknowledge

(LCF) regime and internal fracture initiation

funding provided by Lightweight Innovations

dominated in high cycle fatigue (HCF). A

for Tomorrow (LIFT) and the Office for Naval

unique cone-shaped morphology dominated

Research.

the internal fracture initiation in HCF, with

conclusions or recommendations expressed

Fe-Cr inclusions at the fatigue initiation site.

in this material are those of the author(s) and

Calculations of fracture stress (Smax) and

do not necessarily reflect the views of the

fracture toughness (KQ) for these samples

Office of Naval Research. This material is 14

Any

opinions,

findings,

and

based on research sponsored by Office of Naval Research under agreement number [6]

N00014-14-2-2002. The U.S. Government is authorized to reproduce and distribute reprints

for

notwithstanding

Governmental any

copyright

purposes notation

thereon.

[7]

Materion Brush Inc. of Elmore, OH. has graciously supplied SupremEX® materials for [8]

this work. Faculty and colleagues in Case Western Reserve University have also offered great help in experiments. Their support of this work is acknowledged and appreciated.

[9]

9. Data Availability Statement [10]

The raw/processed data required to reproduce these findings cannot be shared at this time due to technical or time limitations.

[11]

10. References [1]

[2]

[3]

[4]

[5]

G. Davies, “Materials Overview,” in Materials for automobile bodies, Elsevier Ltd., 2012, pp. 4–10. D. B. Miracle, “Aeronautical Applications of Metal-Matrix Composites,” in ASM Handbook - Composites, vol. 21, 2001, pp. 1043–1049. W. H. Hunt and D. B. Miracle, “Automotive Applications of MetalMatrix Composites,” in ASM Handbook Composites, vol. 21, 2001, pp. 1029– 1032. T. Zeuner, P. Stojanov, P. R. Sahm, H. Ruppert, and A. Engels, “Developing trends in disc brake technology for rail application,” Materials Science and Technology, vol. 14, pp. 857–863, 1998. N. Chawla and Y. L. Shen, “Mechanical Behavior of Particle Reinforced Metal

[12]

[13]

[14]

15

Matrix Composites,” Advanced Engineering Materials, vol. 3, no. 6, pp. 357–370, 2001. V. C. Nardone and K. M. Prewo, “ON THE STRENGTH OF DISCONTINUOUS SILICON CARBIDE REINFORECED ALUMINUM COMPOSITES,” Scripta METALLURGICA, vol. 20, pp. 43–48, 1986. S. Suresh, A. Mortensen, and A. Needleman, Fundamentals of Metal Matrix Composites. MA: ButterworthHeinemann, pp. 3-41, 1993. R. J. Arsenault and R. M. Fisher, “MICROSTRUCTURE OF FIBER AND PARTICULATE SiC in 6061 Al COMPOSITES,” Scripta METALLURGICA, vol. 17, pp. 67–71, 1983. T. G. Nieh, “Creep Rupture of a Silicon Carbide Reinforced Aluminum Composite,” Metallurgical Transactions A., vol. 15A, pp. 139–146, 1984. T. Christman and S. Suresh, “MICROSTRUCTURAL DEVELOPMENT IN AN ALUMINUM ALLOY-SiC WHISKER COMPOSITE,” Acta Metallurgica, vol. 36, no. 7, pp. 1691–1704, 1988. S. Suresh, A. Mortensen, and A. Needleman, Fundamentals of Metal Matrix Composites. MA: ButterworthHeinemann, pp. 273-292, 1993. I. Dutta and D. L. Bourell, “A Theoretical and Experimental Study of Aluminum Alloy 6061-SIC Metal Matrix Composite to Identify the Operative Mechanism for Accelerated Aging,” Materials Science and Engineering, vol. A112, pp. 67–77, 1989. T. Ozben, E. Kilickap, and O. Cakir, “Investigation of mechanical and machinability properties of SiC particle reinforced Al-MMC,” Journal of Materials Processing Technology, vol. 198, pp. 220–225, 2008. N. Chawla, C. Andres, J. W. Jones, and J. E. Allison, “Effect of SiC Volume Fraction and Particle Size on the Fatigue Resistance of a 2080 Al/SiCp Composites,” Metallurgical and

[15]

[16]

[17]

[18]

[19]

[20]

[21]

[22]

Materials Transactions A., vol. 29A, pp. 2843–2854, 1998. D. J. Lloyd, “ASPECTS OF FRACTURE IN PARTICULATE REINFORCED METAL MATRIX COMPOSITES,” Acta Metallurgica et Materialia, vol. 39, no. 1, pp. 59–71, 1991. A. R. Vaidya and J. J. Lewandowski, “Effects of SiCp size and volume fracture on high cycle fatigue behavior of AZ91D magnesium alloy composites,” Materials Science and Engineering, vol. A220, pp. 85–92, 1996. N. Chawla, C. Andres, and J. W. Jones, “CYCLIC STRESS-STRAIN BEHAVIOR OF PARTICLE REINFORCED METAL MATRIX COMPOSITES,” Scripta Materialia, vol. 38, no. 10, pp. 1595– 1600, 1998. J. N. Hall, J. W. Jones, and A. K. Sachdev, “Particle size, volume fraction and matrix strength effects on fatigue behavior and particle fracture in 2124 aluminum-SiCp composites,” Materials Science and Engineering, vol. A183, pp. 69–80, 1994. Z. Z. Chen and K. Tokaji, “Effects of particle size on fatigue crack initiation and small crack growth in SiC particulate-reinforced aluminum alloy composites,” Materials Letters, vol. 58, pp. 2314–2321, 2004. C. Kaynak and S. Boylu, “Effects of SiC particulates on the fatigue behaviour of an Al-alloy matrix composites,” Materials and Design, vol. 27, pp. 776– 782, 2006. K. H. Chung et al., “Comparison of consolidation processes of mechanically alloyed Al-SiC metal matrix composite powders,” presented at the Proceedings of the 2018 International Conference of Powder Metallurgy & Particulate Materials, Princeton, NJ, 2018, pp. 643– 658. “Aluminum 2124-T851,” in Properties and Selection: Nonferrous Alloys and Special-Purpose Materials, ASM International 10th Ed., vol. 2, ASM International, 1990. and International

[23]

[24]

[25]

[26]

[27]

[28]

[29] [30]

[31]

[32]

16

Registration of 2124A (PN15-59 Project) from The Aluminum Association, June 2915. ASTM International, “ASTM E8-09 Standard Test Methods for Tension Testing of Metallic Materials.” ASTM International, 2009. ASTM International, “ASTM E466-15 Standard Practice for Conducting Force Controlled Constant Amplitude Axial Fatigue Tests of Metallic Materials.” ASTM International, 2015. C. Park, “Mechanical Performance and Structure-Property Relations in 6061B Aluminum Metal Matrix Composites.,” M.S. Thesis, Case Western Reserve University, 2018. ASTM International, “ASTM E111-10 Standard test method for Young’s modulus, tangent modulus, and chord modulus.” ASTM International, 2010. J. Xia, “Tension and Fatigue Behavior of Al-2124/SiC-particulate Metal-Matrix Composites,” M.S. Thesis, Case Western Reserve University, 2019. S. S. Manson, “Inversion of The StrainLife and Strain-Stress Relationships for Use in Metal Fatigue Analysis,” Fatigue of Engineering Materials and Structures, no. 1, pp. 37–57, 1979. Level 3 Database, “Silicon Carbide (sintered).” CES EduPack 2018. Z. Zhang and D. L Chen, “Contribution of Orowan strengthening effect in particulate-reinforced metal matrix nanocomposites,” Materials Science and Engineering: A, vol. 483–484, pp. 148– 152, 2008. J. J. Lewandowski and P. Lowhaphandu, “Effects of hydrostatic pressure on mechanical behaviour and deformation processing of materials,” International Materials Reviews, vol. 43, no. 4, pp. 145–187, 1998. J. J. Lewandowski and P. Lowhaphandu, “Effects of hydrostatic pressure on the flow and fracture of a bulk amorphous metal,” Philosophical Magazine A, vol. 82, no. 17–18, pp. 3427–3441, 2002.

[33] H. A. Hassan and J. J. Lewandowski, “Fracture Toughness and Fatigue of Particulate Metal Matrix Composites,” Comprehensive Composite Materials II, vol. 4, pp. 86–136, 2018. [34] J. J. Lewandowski and P. M. Singh, “Fracture and Fatigue of DRA Composites,” in ASM Metals Handbook No. 19, Section 7, ASM International, 1996. [35] N. E. Dowling, “Fatigue of Materials: Introduction and Stress-Based Approach,” in Mechanical Behavior of Materials, Fourth Edition vols., England: Pearson, 2013, pp. 416–478. [36] N. E. Dowling, “Strain-Based Approach to Fatigue,” in Mechanical Behavior of Materials, Fourth Edition vols., England: Pearson, 2013, pp. 745–790. [37] I. Uygur and M. K. Kulekci, “Low Cycle Fatigue Properties of 2124/SiCp AlAlloy Composites,” Turkish J. Eng. Env. Sci., vol. 26, pp. 265–274, 2002. [38] W. D. Pilkey, Peterson’s Stress Concentration Factors, Second Edition. NY: John Wiley & Sons, Inc., 1997. [39] N. E. Dowling, “Fracture of Cracked Members,” in Mechanical Behavior of Materials, Fourth Edition vols., England: Pearson, 2013, pp. 334–361.

[40] X. Li, Z. M. Yin, L. Zhong, Q. L. Pan, and F. Jiang, “High cycle fatigue characteristics of 2124-T851 aluminum alloy,” Front. Mater. Sci., vol. 1, no. 2, pp. 168–172, 2007. [41] Materion Corporation, “SupremEX Forged Plate Materials Data Sheet.” 2016. [42] T. S. Srivatsan, S. Anand, S. Sriram, and V. K. Vasudevan, “The high-cycle fatigue and fracture behavior of aluminum alloy 7055,” Materials Science and Engineering, vol. A281, pp. 292–304, 2000. [43] J. J. Bonnen, J. E. Allison, and J. W. Jones, “Fatigue Behavior of a 2xxx Series Aluminum Alloy Reinforced with 15 Vol Pct SiCp,” Metallurgical Transactions A., vol. 22A, pp. 1007–1019, 1991. [44] T. S. Srivatsan and M. Al-Hajri, “The fatigue and final fracture behavior of SiC particle reinforced 7034 aluminum matrix composites,” Composites: Part B, no. 33, pp. 391–404, 2002. [45] T. Mochida, M. Taya, and D. J. Lloyd, “Fracture of Particles in a Particle/Metal Matrix Composite under Plastic Straining and Its Effect on the Young’s Modulus of the Composite,” Materials Transactions, vol. 32, no. 10, pp. 931– 942, 1991.

17

Figure Captions: Figure 1. Microstructure by SEM taken from the grip portion of a broken tension sample (R-225XE-T4-L3), shown with different magnifications. The dark region is Al matrix, and the bright particles are SiC. Figure 2. Engineering and true stress-strain curves for R-225XE-T4-L2. The strain is measured by UVID non-contact extensometer. The discontinuity of the curves (at 0.04 strain) in the plot is caused by removing the contact extensometer that interfered with the UVID measurement. Figure 3. The S-N curve plotted with maximum stress vs. cycles to failure. Fracture Initiation positions are indicated for all specimens with SEM or Keyence optical microscope. Triangles in the plot refer to specimens fractured from the inside with a unique cone-shaped feature. Figure 4. SEM fractography for tension sample with T4 heat treatment tested in the longitudinal direction. Ductile fracture features were present. Shear lips are indicated by red arrows. Figure 5. Fracture surface of a fatigue sample with T4 heat treatment tested in the longitudinal direction, with σmax = 468.9 MPa. The defect (arrow) is beneath the polished surface. Figure 6. Fracture surface of a fatigue sample with T4 heat treatment tested in the longitudinal direction, with σmax = 482.0 MPa. Fracture initiated from a large (i.e. 30 µm) SiC particulate well in excess of the average reinforcement size (3 µm). Figure 7. Fracture surface of a fatigue sample with T4 heat treatment tested in the longitudinal direction, with σmax = 481 MPa. Fracture initiated from a Fe-Cr rich inclusion (solid lined boxes) and grew along the cone surface (dashed lined box). Figure 8. Height map of a cone apex measured using optical microscopy. The cone apex angle is 86 degrees. Figure 9. Typical SEM images from fatigue specimens, showing with cone features (a) side surface of the cone exhibiting shear features, and (b) typical catastrophic overload failure illustrating dimpled features. Figure 10. Illustration of fracture position deviation D with respect to the minimum diameter. The original S-N curve was adjusted based on local stresses calculated from D. Figure 11. S−N curve based on total strain range ∆ε. The fit curve following Universal Slopes Equation is shown in the plot, with R2 = 0.81. The additional red data points are obtained from literature for AMC225 [40], which is very similar to SupremEX® 225XE, but of earlier generation. Figure 12. Comparison of fatigue performance between SupremEX-225XE-T4-L and monolithic 2xxx/7xxx Al alloys. Figure 13. Comparison of fatigue performance between SupremEX-225XE-T4-L and other conventional MMCs. Figure 14. Normalized S-N curves based on ratio of σlocal against 0.2% σy (top) and UTS (bottom). The dashed line separates the LCF and HCF regions at 105 cycles. The 2xxx series Al-SiCp MMC [43] is shown in green and the 7034 Al-SiCp MMCs [44] are shown in red

18

Table 1. Composition of 2124A aluminum alloy, in weight percent. Differentiation between 2124 and 2124A resides in the allowable oxygen content. (limits % max unless range provided) [22]

Al Balance

Cu 3.8 − 4.9

Mg 1.2 − 1.8

Mn 0.3 − 0.9

Fe 0.3

Zn 0.25

19

Si 0.2

Ti 0.15

O 0.6

Others Each, 0.05 Total, 0.15

Table 2. Mechanical properties calculated for R-225XE-T4-L tension samples.

Sample R-225XE-T4-L1 R-225XE-T4-L2 R-225XE-T4-L3 Average

E (GPa) 108 102 96 102 ± 5

0.2% σy (MPa)

UTS (MPa)

Elongation (%)

ROA (%)

468 464 470 467 ± 2

565 638 648 617 ± 37

5.8 7.2 5.2 6.1 ± 0.8

6.8 7.0 7.7 7.2 ± 0.4

20

Table 3. Summary of fracture features for all fatigue specimens. Samples labeled with * refer to those with the cone feature.

Sample ID L1 *L2 L3 L4 *L5 L6 *L7 L8 L9 *L10 L11 L12 *L13 *L14

Local σmax (MPa) 494 481 469 483 461 448 474 483 474 469 521 538 452 497

Local Initiation Edge Internal Edge Edge Internal Internal Edge Edge Internal Edge Edge Internal Internal

Defect type Surface defect Fe-Cr Inclusion Surface defect Surface defect Fe-Cr Inclusion RUNOUT Fe-Cr Inclusion Surface defect Surface defect Oxide inclusion Surface defect Oxide inclusion Fe-Cr Inclusion Fe-Cr Inclusion

21

Sample ID

σmax

Initiation

Defect type

*L15 L16 L17 *L18 L19 L20 L21 *L22 L23 L24 L25 L26 *L27 L28

(MPa) 440 528 / 442 / 382 483 454 471 488 476 481 474 476

Internal Edge / Internal / Edge Internal Internal Edge Edge Edge Edge Internal Edge

Fe-Cr Inclusion Fe-Cr Inclusion / Fe-Cr Inclusion / Fe-Cr Inclusion Large SiC Fe-Cr Inclusion Surface defect Fe-Cr Inclusion Fe-Cr Inclusion Surface defect Fe-Cr Inclusion Fe-Cr Inclusion

Table 4. Summary of measurements for most samples with unique cone fracture features from Keyence optical microscopy.

Sample ID L2 L7 L10 L13 L14 L15 L18 L22 L27

Apex angle (Degree) 100.1 106.9 98.3 113.5 106.5 96.1 85.5 106.7 92.9

Max cone diameter (mm) 2.63 2.82 1.57 2.86 1.26 2.78 3.37 2.16 2.25

22

Table 5. Average values and percentage change in mechanical properties of R-225XE-T4-L samples, com-pared to monolithic 2124A aluminum alloy.

Average Values R-225XE-T4-L Monolithic 2124 Al [25] Percentage Change

E (GPa) 102 73.1

0.2% σy (MPa) 463.7 441

UTS (MPa) 617 483

Elongation (%) 6.1 9.0

True fracture stress (MPa) 721.0 /

39.7%

6.2%

40.9

-32.6%

/

23

Table 6. Calculations of fracture stress and fracture toughness for specimens with cones.

Load Range (kN) 1.61 − 16.1 1.60 − 16.0 1.55 − 15.5 1.55− 15.5 1.55 − 15.5 1.55 − 15.5 1.53 − 15.3 1.50 − 15.0 1.44 − 14.4

Af (mm2) 33.9 36.5 30.8 30.8 36.3 35.9 28.1 33.6 31.0

Smax (MPa) 489 452 519 518 440 445 560 459 479

σmax (MPa) 503.3 500.6 482.7 482.7 482.7 482.7 475.8 468.9 448.2

24

KQ (MPa√m) 13.4 16.1 17.7 17.7 16.8 15.3 13.5 15.8 16.3

®

Table 7. Detailed information for SupremEX 225XE, Al alloys, and other MMCs, including composition and basic mechanical properties. The amount of SiC reinforcements are presented in volume fracture. All mechanical properties are measured at room temperature.

Material 2124A/SiC/25p_3 µm Monolithic 2124-T851 [25] Monolithic 7055-T7751 [45] 2xxx/SiC/15p_5 µm [46] 7034/SiC/15p (PA)_~20 µm [47] 7034/SiC/15p (UA)_~20 µm [47]

E (GPa) 103 73.1 70 99 91 90

UTS (MPa) 617 441 630 552 683 709

25

0.2% σy (MPa) 467 483 610 397 612 622

Elongation 6.1% 9.0 % 12% 8.9% 1.9% 1.8%

Declaration of interests ☒ The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper. ☐The authors declare the following financial interests/personal relationships which may be considered as potential competing interests: