2013smart.pdf

6th ECCOMAS Conference on Smart Structures and Materials

SMART2013

Politecnico di Torino, 24-26 June 2013

E. Carrera, F. Miglioretti and M. Petrolo (Editors)

www.smart2013.com

STRENGTH OF SMA/GFRP INTERFACE IN A NEW DESIGNED

RAILWAY COLLECTOR

M. Bocciolone*, M. Carnevale*, A. Collina*, N. Lecis*, A. Lo Conte*, B. Previtali*

C.A. Biffi, P.Bassani

, A. Tuissi

* Politecnico di Milano

Department of Mechanical Engineering

Via La Masa, 1

20154, Milano, Italy

National Research Council CNR

Institute for Energetics and Interphases

Corso Promessi Sposi, 29

Lecco, Italy

Key words: SMA/GFRP Composite, Interface load transfer, Laser patterning, Laser texturing.

Summary. In this paper the strength of the interface between SMA reinforcements and host

composite of a hybrid laminated composite, proposed for the lateral horns of a railway

collector, will be studied in details by means of pull-out tests. The effect of the geometrical

pattern of the SMA sheets and of the texturing of its surface on the load transfer capacity of

the interface is investigated.

1 INTRODUCTION

The application field of SMA materials as high damping reinforcement of glass fiber

reinforced polymer (GFRP) composites is disclosing interesting perspectives [1-3]. However,

the desirable function of SMA composites could not be precisely controlled without the

complete understanding of their stress transfer properties and interfacial bonding behavior

between the embedded SMA reinforcement and the surrounding polymeric GFRP. The

strength of the reinforcement exceeds considerably the strength of the host composite and,

thus, the interface is responsible for the load transfer between the two constituents the

composite material. When shear stress between constituents increase upon the interfacial

shear strength, interfacial debond will start immediately. The greater is the critical debonding

stress, the stronger the hybrid composite material will be.

In [4] a new design of a Cu based SMA/GFRP lateral horn of a railway collector was

proposed. A hybrid composite architecture was adopted as shown in Figure 1, where a

laminated GFRP was used as the host composite, and a CuZnAl SMA alloy was the high

damping material used for the reinforcement. The type and weight and the direction of

M. Bocciolone, M. Carnevale, A. Collina, N. Lecis, A. Lo Conte, B. Previtali C.A. Biffi, P.Bassani, A. Tuissi

2

orientation of fiberglass layers was carefully selected to meet engineering horn design

characteristics in both strength and modulus (stiffness/flexibility).

Synergistic contribution of the performance parameters associated with the SMA, including

specific damping, specific stiffness, volume fraction, and SMA through-the-thickness location

are taken into account. The SMA alloy was embedded, in the shape of two microcut thin

sheets (Figure 2), between the bulk and external layers of the cross section of the horn.

Figure 1: Architecture of the SMA//GFRP hybrid

composite horn. Grey [-45/+45]n layered GFRP host

composite. Brown: Embedded CuZnAl SMA thin

microcut sheets.

Figure 2: Pattern of laser microcutting of CuZnAl

sheets.

With the renewed design the structural damping of the first flexural mode of the horn was

significantly enhanced without affecting its natural frequencies. Infact, in accordance with the

initial requirements, its geometrical configuration, its flexural stiffness, and its weight were

retained. In the optimization process of the flexural stiffness, weight and damping capacity of

the horn, the microcuttig of the thin SMA sheets proved to be a key feature. At the same time,

this feature has been proposed to improve the adhesion, and the load transfer, between the

GFRP laminated composite and the SMA reinforcements.

In this paper the strength of the interface between SMA reinforcements and host composite

will be investigated. In particular, we are interested to relate the load transfer ability of the

CuZnAl SMA-GFRP interface to the damping performance at the structural level of the new

hybrid composite material [5, 6]. Adhesional parameters for many fiber-polymer systems and

also for NiTi SMA-polymer system are available in literature [7-8], but at-date no

experimental results have been published for Cu based SMA alloy-polymer system.

To characterize the interfacial interaction between fibers and matrices different parameters

have been proposed: the interfacial shear strength (i), the energy release rate (Gc), the

frictional shear stress (f), and the friction coefficient (). To obtain experimental data a large number of various micromechanical tests have been developed [7, 9, 10]. The single fibre

pull-out test (SFPO-test) is widely accepted as one of the most important test methods

considered as a mean of investigating the interfacial adhesion quality and interfacial


3

properties between fibres and matrix and the elastic stress transfer in the fibre pull-out

problem. From specific points taken from the respective load-displacement curves of the

SFPO-test it is possible to calculate the set of parameters to relate the load transfer ability of

the interface to adhesional parameters [9, 11].

This paper reports the experimental results of pull-out tests of SMA/GFRP hybrid composite

in which a thin sheet of CuZnAl SMA is embedded in a layered GFRP host composite. Force-

displacement curves and the interfacial bonding properties will be presented for three

different configurations of the embedded CuZnAl SMA sheet.

The application of a stress-controlled debonding model to relate interfacial parameters to load

transfer capacity will be also proposed.

2 MATERIAL DESCRIPTION

The hybrid composite material is a GFRP host composite ([+45/-45]n 3M-SP250 S29A)

with embedded CuZnAl (Cu66Zn24Al10 at.%) Shape Memory Alloy sheets. The elastic

properties of the GFRP unidirectional layer are shown in Table 1.

Table 1. Elastic properties of the fiber glass/epoxy resin 3M-SP250 S29A.

Traction Compression

Exx [GPa] Eyy [GPa] Exx [GPa] Eyy [GPa] Gxy [GPa] nxy

45.7 13.5 47.8 12.8 5.4 0.27

The SMA alloy is induction melted under an inert atmosphere and the produced ingot,

100x70x12 mm3 in size, is hot and cold-rolled down to 300 m thick sheets. (30-mm in width

and 400-mm in length).

The transformation temperatures were measured by means of a differential scanning

calorimeter (DSC,-TA Instrument mod. Q100), calibrated with a standard indium reference.

Specimens, weighing about 10 mg, were scanned at a heating/cooling rate of 10 C/min

within the temperature range [10110 C]. The thermographs shown in Figure 4, outline the

characteristic transformation temperatures: Mf=50 C; Ms=63 C; As=60 C and, Af=68 C.

The DSC scan confirms the martensitic structure in the temperature range of interest.

Figure 3: DSC scan, heating/cooling rate: 10 C/min.


4

3 EXPERIMENTAL

The hybrid composite specimen has been particularly designed for the pull-out tests, shape

and dimensions are reported in Figure 4. A sheet of CuZnAl SMA has been embedded in the

middle of the thickness of a layered GFRP composite [+45/-45]22. The SMA sheet is 0.3 mm

thick and it is not prestrained, its embedded length is 30 mm and its free length is 70 mm.

For comparison three different configuration of the SMA sheets have been considered. For

each configuration of the embedded area three specimens has been manufactured (Figure 5).

In the first case the SMA sheets are as cold rolled. In the second case the interface surfaces of

the embedded length has been laser textured. An IPG Photonics Q-switched fiber laser was

used. The texturing pattern (Figure 6) is a square array of circular dimples. The distance

between dimples is 100 m, the diameter of the dimple is about 50 m, the depth less than 10 m. The resulting area ratio is between 20% and 25%. In the third case the embedded length has been laser microcut with a pattern of elliptical holes, according to the solution proposed

for the manufactured horn prototype (Figure 2). After the preparation of the different surface

to be embedded in the hybrid composite samples the nine sheets have been heat treated all

together at 750 C for 30 min and then water quenched (WQ).

Laser texturing allows us to obtain controlled surface roughness for improved bonding

strength between SMA reinforcement and host composite by increasing both the adhesion

surface and the friction at the interface. On the other hand laser patterning is proposed to

improve the load transfer capability between SMA reinforcement and host composite by

means of the change of the interfacial bonding profile. The goal is an increase of the bonding

strength despite of the fact that, to meet the design requirement of keeping the original

stiffness and weight of the component, the SMA surface has been laser microcut and the

adhesion surface is reduced.

The nine pull-out tests have been conducted in air at room temperature by means of an

Instron electro-mechanic testing machine, equipped with a load cell of 10 kN. The complete

test setup is shown in Figure 7. The composite specimen has been put under a metallic holder,

which has a thin slot in the center to align the axles of the sample with the axis of the holder.

The free end of the sample has been fixed at the bottom grips of the testing machine and the

tensile load has been applied to the metallic holder by means of the upper grips, at a constant

displacement rate of 1 mm/mm. The bottom surface of the layered GFRP composite provides

the reaction force to the applied tensile force by means of the contact with the holder. The

CuZnAl thin sheet has been allowed to go through the thin slot of the holder with a negligible

friction due to the clearance between the CuZnAl sheet and the slot. The sliding behavior

during the pull-out test has been monitored using a MTS exstensometer. The arm of the

extensometer records the displacement of the bottom surface of the holder with respect to the

fixed end of the sample. Data of load (F) applied to the holder keeping the hybrid composite

sample has been collected as a function of the displacement (u) of its bottom surface.


5

a) b)

c)

Figure 4: Hybrid composite

specimen for pull-out tests.

Dimensions are given in millimeters.

Figure 5: Different configuration of the embedded area of CuZnAl SMA

sheets. a) As cold rolled and heat treated. b) Laser texturized and then

heat treated. c) Laser patterned and then heat treated.

Figure 6: Patterns of the laser texturing.

30

70

5 22

20


6

(a)

(b)

Figure 7: (a) Experimental set-up for pull-out tests. (b) Schematic illustration of the thin sheet pull-out test.

Holder

Exstensometer

Fixed end

of sample

Slot for sliding of

the SMA sheet

F

u

Sheet

Matrix

t

S

(x)

u(x)

F/2 F/2

F

x


7

3 FORCE D-DISPLACEMENT CURVES AND THEIR INTERPRETATION

The force displacement curve for the pull-out tests performed on the hybrid composite with

the embedded surface of the SMA sheets As rolled, and Texturized, are reported in Figure 7, and 8, respectively. For these samples the force displacement curves show the

typical trend of the single long fiber pull-out test. In the first stage (0 F Fd) the SMA sheet- host composite interface remain intact and the curve is nearly linear. When the external

load reaches some critical value (debond force, Fd), the SMA sheet begin to debond off the

host GFRP composite through interfacial crack propagation. In this second stage (Fd F Fmax), the registered force continues increasing with the displacement (or with the crack

length), because frictional load in debonding region is added to the adhesional load from the

intact part of the interface. The maximum load is achieved under partial debonding

conditions. After a peak load, (Fmax), is reached, the crack propagation became unstable, the

whole embedded length fully debonds, the SMA sheet slides through the host composite with

friction and through the slot of the holder with clearance and the measured force drops

immediately. Beyond this point and until complete pull-out of the 30 mm embedded length,

the tail force is due to frictional interaction between the SMA sheet and GFRP host

composite. Table 1 shows a series of data generated for the single pull-out tests of As rolled , and Texturized samples.

Table 1 Value of the load at the initiation of the debond and at the complete pull-out of the As rolled and Textured embedded sheet.

The experimental results are reported also as mean value and standard deviation related to

three experimental values of each series of data. Clearly the texturing of the surface makes

differences in the load or stresses associated with the event of debond initiation (Fd), and the

difference became more significant when we compare the load for the complete pull-out of

the sheet (Fmax). Due to the gradual trend of the tail force in the test with Texturized sheet embedded, a comparison of the tail force for the two series of tests is not presented.

The force displacement curve for the pull-out tests performed on the hybrid composite with

the embedded surface of the Microcut SMA sheets are reported in Figure 10. Also for these samples it has been observed a nearly linear first stage, where the force displacement

curves (0 F Fd) describes the elastic response of the composite with intact interface and a second stage with F >Fd. But, in this case, after the initiation of the debonding a catastrophic

failure of the embedded sheet occurs. Figure 11 shows the section of the sample where the

failure of the sheet has occurred for all the three performed tests, and a SEM magnification of

Experiment Fd [N] Fmax [N]

Single value Mean Single value Mean

Test 1 250

238+34

500

588112 Test 2 200 555

Test 3 265 715

Test 4 310

3036

1200

1290108 Test 5 300 1410

Test 6 300 1260


8

Figure 8: Applied force (F) versus displacement (u) of the single pull-out test for the hybrid composite with

As rolled sheets embedded.

Figure 9: Applied force (F) versus displacement (u) of the single pull-out test for the hybrid composite

with Textured sheets embed

0 5 10 15 20 25 300

0.2

0.4

0.6

0.8

1

Displacement [mm]

Fo

rce

[kN

]

Test 1

Test 2

Test 3

0 5 10 15 20 25 300

0.5

1

1.5

Displacement [mm]

Fo

rce

[kN

]

Test 4

Test 5

Test 6

0 1 2 30

0.1

0.2

0.3

0.4

0.5

0.6

Displacement [mm]

Forc

e [

kN

]

Test 2

Fmax

0 2 4 60

0.2

0.4

0.6

0.8

1

1.2

1.4

Displacement [mm]

Forc

e [

kN

]

Test 4

Fmax

Fd

Fd


9

Figure 10: Applied force (F) versus displacement (u) of the single pull-out test for the hybrid composite with

Microcut sheets embedded.

Figure 11: Details of the failure of the SMA sheet during the pull-out tests of Hybrid composite with Microcut sheets embedded.

0 1 2 3 4 50

0.2

0.4

0.6

0.8

1

Displacement [mm]

Fo

rce

[kN

]

1 stage Test 9

Test 7

Test 8

Test 9

Cross section of sheet failure SEM image of the failed sheet

Fmax Test 9

Fd Test 9


10

a failed sheet. In this case the elliptical microholes represent a discontinuity of the interface

that modify both the stress profile in the matrix and in the fiber and the interfacial shear stress

distribution along the embedded length. As the load increases beyond the load of the debond

initiation, a concentration of normal stress in the section of the sheet occurs and the sheet fail

Table 2 shows data generated for the single pull-out tests of Microcut samples.

Table 2 Value of the load at the initiation of the debond and at the complete pull-out of the Microcut embedded sheet

As for textured sheets, also the presence of the pattern of elliptical microcut of the surface makes differences in the load or stresses associated with the event of debond initiation (Fd).

Despite of the reduction of interfacial area introduced to have a light and not stiffen sheet, the

load for debond initiation is comparable with the value obtained by means of the texturing of

the surface. If we consider the maximum load (Fmax) we can observe that the value is higher

than the value observed for sample with as rolled sheets, but the main results is that the catastrophic failure of the sheet account for a better performance of the interface able to avoid

the sheet delamination.

To characterize the quality of interfacial bonding, the apparent interfacial shear strength is

calculated according to:

A

Fappi

*2

max, (1)

where A is the interfacial area.

According to equation (1) and the average value of the maximum load (Fmax) (Table 1 and

Table 2), the apparent interfacial shear strength cab be obtained as i,app less then 1MPa, about 1 MPa, and about 2 MPa for As rolled, Textured, and Microcut sheet respectively.

This i,app value usually allow us to estimate the relative efficiency of matrix and fiber surface adhesion and suffice to distinguish between better/worse interfacial bonding.

However, a quantitative characterization of the fiber-matrix interface properties require a

more adeguate appoach which should tahe into account the actual mechanism of interfacial

failure and include local interface parameters instead of averaged or apparent ones.

4 THEORETICAL MODEL OF INTERFACIAL FAILURE

Quantitative interpretation of single pull-out test for single continuous sheet as for As rolled and Texturing experimental test can be accomplished from both the data associated with the start of the fiber debond and with the fiber pull-out. In the first stage (bonded

Experiment Fd [N] Fmax [N]

Single value Mean Single value Mean

Test 7 310

305

625

725114 Test 8 -- 700

Test 9 300 850


11

condition), it is assumed that the interface between the sheet and the matrix is perfectly

bonded. In the second stage (bonded and debonded transition), debonding starts when the

interfacial shear stress reaches the ultimate shear stress (d). The shear stress over the debonded region is replaced by the frictional stress (f). As the load is increased, the debonding zone extends over the entire interface. Once the sheet is completely debonded,

frictional forces are the only means of resisting the slip. In the final stage (slipping condition),

the static frictional stress was replaced by a dynamic frictional stress, (s) which was lower in magnitude. The load-slip relationship in this range is a linear descending model.

To relate the interfacial shear stress of the sheet to the debond force and the embedded

length the most widely used model is the shear lag model, originally proposed by Cox and

subsequently developed by others [13, 14] which focuses on the transfer of tensile stress from

elastic matrix to elastic fiber by means of interfacial shear stress. Let the x direction be

parallel to the sheet axis Figure 7, the axial stress distribution along the axis of the sheet and

the interfacial shear stress has been derived as:

2/1

/2sinh

/2sinh)(

tnL

txLn

tb

Fxf (2)

t

Lnech

t

xLn

tb

nFx

2cos

2cosh

2)( (3)

where

2/1

)/ln(1*2

tSE

En

mf

m

(4)

In the previous equations (2), (3), and (4), according to the schematic of Figure 7: F is the

applied load, t the thickness of the sheet, and S the thickness of the host composite. Moreover,

b is the width of the sheet, L the embedded sheet length, Em is the Young modulus of the host

composite in the x direction, Ef is the Young modulus of the CuZnAl SMA sheet, and m the Poisson coefficient of the host composite.

Applying eq. (2) at x=0, the debonding shear stress (d) is deduced from the debond force, Fd, corresponding to the kink in the force displacement curve (Figure 8 and 9) [12]:

tb

t

nLnFd

d2

2coth

(5)

In the model, the debonding shear stress (d) (local shear stress near the propagating crack


12

tip) is supposed to be constant during the test and not depend on the crack length as shown in

Figure 12 for the Test 2 (As rolled). The following values have been used hereafter: t= 0.3

mm, S=5 mm, L=30 mm, Ef=17 GPa, Em=60 GPa, m=0.27. In correspondence of the crack

initiation (Figure 12a) the maximum of the interfacial shear stress is d=11 MPa at x=0 and decreases along the embedded length according to the distribution given by equation (3). In

the same figure the distribution of the axial stress along the sheet axis given by equation (2)

also is reported. For an advanced stage of the crack propagation, in the debonded region the

shear stress is determined by interfacial friction and is assumed to be constant. In the intact

zone, the shear stress start for the same value d and decreases along the remaining embedded length of the sheet. This condition is shown in Figure 12b for the same Test 2 (As rolled)

when the displacement is 1 mm. In this case the distribution of the axial stress along the sheet

axis that is constant in the debonded region and decrease in the remaining embedded length.

(a)

(b) Figure 12: Schematic of the distribution of the axial stress and the interfacial shear stress along the axis of

the CuZnAl sheet for Test 2 (As rolled): a) At the moment of the crack initiation. b) When the displacement is

equal to 1 mm.

0 1 2 3 4 50

5

10

15

20

25

30

35

x [mm]

(x

)/ (

x)

[MP

a]

(x)

(x)

0 1 2 3 4 50

10

20

30

40

50

60

70

x [mm]

(x

)/ (

x)

[MP

a]

(x)

(x)

F=Fd

u(x)=0 x

F>Fd

u(x)=1 x

debonded

F/2 F/2

x=0 x=debonded x=L length

x=0 x=L


13

Table 3 reports the value of the maximum interface shear stress d for the two series of tests As rolled, and textured, and the related value d of the axial stress in the sheet, at x=0, when the debond starts.

Table 3 Value of of the maximum interface shear stress d for the two series of tests As rolled, and Textured, and the related value d of the axial stress in the sheet, at x=0, when the debond starts.

An attempt to quantitative interpretation of single pull-out test for Microcut sheet can be done on the basis of the following observations:

1) the thin ligaments of the elliptical microholes pattern guarantee the important feature of continuous load transfer at the interface between the CuZnAl SMA insert and the host

composite, as typically occurs for long fiber embedded.

2) the microholes of the embedded sheets became a bridge of the host composite across the thickness of the sheet. When, during the pull-out test, debond of the sheet starts,

this bridge virtually works as a strong reinforcement of the interface.

According to the observation 1), attempt can be done to use the equation (2) and (3) to

interpret the results of the pull-out tests for Microcut sheets. We refer to the cross section of the microcut sheet failure and define an equivalent area of

the sheet as the total area of the ligaments cross section A=tb. In this case the width b is an equivalent width of the sheet to be used in equation (2) and (3).

Table 4 reports the value of the maximum interface shear stress d and the related value d for the tests Microcut. An improvement of the performance of the interface is clearly observed.

Table 4 Value of the maximum interface shear stress d for the two series of tests Microcut, and the related value d of the axial stress in the sheet, at x=0, when the debond starts.

When the applied load increases the bridge of the host composite across the thickness of

the sheet make the interface to stick and the ultimate strength of the material in the ligament is

reached before the debonding of the interface occurs.

Experiment d [MPa] d [MPa] Test 1 14 41

Test 2 11 33

Test 3 15 44

Test 4 18 51

Test 5 17 50

Test 6 17 50

Experiment d [MPa] d [MPa] Test 7 50 143

Test 9 48 138


14

5 CONCLUSIONS

In this paper the strength of the interface between SMA reinforcements and host composite

of a hybrid laminated composite, proposed for the lateral horns of a railway collector, has

been studied in details by means of pull-out tests. In particular the attention has been focused

on the effect of the geometrical pattern of the SMA sheets and of the texturing of its surface

on the load transfer capacity of the interface is investigated. Three configurations have been

considered. a) As cold rolled and heat treated. b) Laser texturized and then heat treated. c)

Laser patterned and then heat treated.

A model derived from the single fiber configuration has been adapted for the interpretation

of the results.

Acknowledgements

The authors gratefully acknowledge the financial support from the SMILE project Young Researcher Award of the Department of Mechanical Engineering, Politecnico di Milano 2008.

The authors wish to thank the Composite Materials Laboratory of the Department of

Aerospace Engineering of the Politecnico di Milano for the manufacturing of the hybrid

composite samples.

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