Mechanical behaviour of polymer modified mortars

We investigate the mechanical behaviour of a mortar and five polymer (latex styrene–butadiene) modified mortars (PMMs) with different polymer contents. The mechanical characterisation of the materials is based on compression and three-point bending tests. As expected, compression tests reveal a decrease of the PMMs rigidity and compressive strength for increasing polymer content. On the contrary, three-point bending tests show an increase of the flexural strength from a polymer-to-cement weight ratio higher or equal to 10 wt.%. After some considerations on material porosity and cement hydration, we establish that the main cause of the PMMs flexural strength increase is the latex percolation into a continuous network. Finally, analysis on damage initiation and material rigidity seems to indicate that latex addition limits skin and inner material micro-cracking due to sample drying out.

Mechanical behaviour of polymer modified mortars

S Pascala, A Allicheb,d,∗, Ph Pilvinc

aDEN/DM2S/SEMT/LM2S, CEN Saclay, 91191 Gif-Sur-Yvette Cedex, France

bLaboratoire MSS-Mat, École Centrale, Paris, France

cLG2M, rue de Saint Maudé, 56321 Lorient Cedex, France

dGrande Voie des Vignes, 92295 Chˆatenay-Malabry, France

Received 30 May 2003; received in revised form 22 October 2003

Abstract

We investigate the mechanical behaviour of a mortar and five polymer (latex styrene–butadiene) modified mortars (PMMs) with different polymer contents The mechanical characterisation of the materials is based on compression and three-point bending tests As expected, compression tests reveal a decrease of the PMMs rigidity and compressive strength for increasing polymer content On the contrary, three-point bending tests show an increase of the flexural strength from a polymer-to-cement weight ratio higher or equal to 10 wt.% After some considerations on material porosity and cement hydration, we establish that the main cause of the PMMs flexural strength increase is the latex percolation into a continuous network Finally, analysis on damage initiation and material rigidity seems to indicate that latex addition limits skin and inner material micro-cracking due to sample drying out

© 2004 Elsevier B.V All rights reserved

Keywords: Polymer modified mortar; Latex additive; Mechanical testing; Compressive strength; Flexural strength

1 Introduction

Polymer modified mortars (PMMs) were developed to

deal with some needs in building and public works These

materials are used as tile adhesives, frontage coatings,

cov-erings, and as repair materials in road building The gaol

was to obtain materials easy to implement and with a better

useful life than usual mortars Besides, they should exhibit

good adhesive properties as well as an improved resistance

to environmental stress, as wall insulating against humidity

in particular

Various kinds of polymer [1,2] (vinyl acetate, styrene

butadiene rubber for example) are used for these

materi-als Polymer content is usually given according to the dry

polymer-to-cement weight ratio (P/C) The presence of

poly-mer in a mortar leads to different microstructures depending

on the way the polymer extends turns out into the material,

which is also closely related to the polymer content

Studies on PMMs are generally based on comparisons

between PMMs with various polymer contents and a mortar

∗Corresponding author Tel.:+33-1-41-13-13-02;

fax: +33-1-41-13-14-30.

E-mail address: alliche@mssmat.ecp.fr (A Alliche).

as reference Among these studies, we can distinguish the way in which PMMs have been worked out: (i) by adjusting the water to cement weight ratio (W/C) in order to keep constant the viscosity or the flow of the different materials [3–6], (ii) by keeping constant the W/C ratio, aiming to provide a similar hydration of the cement matrix [2,7–9]

In the first case (i), the flowing effects of the polymer due

to the surfactants that it contains require to lower the W/C ratio for increasing polymer contents In the second case (ii), the usual air entrainment effect of the polymer increases the porosity of the composite

For PMMs made up by keeping constant the viscosity between the different materials, the results indicate that the mechanical strength determined both in tension, compres-sion and three point bending increases for higher polymer contents[1,3,4,6], while elastic modulus decreases[1,5,6] Moreover, the polymer would improve the water retention capability of these materials, and then limit their drying out for the benefit to cement hydration [1,3–5] Accord-ingly, long-term hydration would be promoted into PMMs [3] These composites also turn out to be more resistant to freeze-thaw cycling and, more generally, would ensure a bet-ter protection against environmental stress[1,3,4] However, these studies involve comparisons between cement based

0921-5093/$ – see front matter © 2004 Elsevier B.V All rights reserved.

doi:10.1016/j.msea.2004.03.049

Table 1

Mix proportions of materials: polymer (P), super-plasticizer (SP),

an-tifoamer (AF), water (W), sand (S) and silica fume (SF) to cement (C)

weight ratios

Constituents P/C SP/C AF/C W/C S/C SF/C

0.05

0.075

0.125

0.15

materials in which not only the polymer content is varying

but also the W/C ratio In this case, it is not obvious to settle

what the real influence of the polymer on the overall

prop-erties of PMMs is

The compressive strength of PMMs made up with an

iden-tical W/C ratio decreases for higher polymer contents while

their flexural strength increases [1,7,9] The improvement

of the flexural strength may be due to the reinforcement of

the Interfacial Transition Zone between the cement matrix

and the aggregates[1,10], and to the bridging of the cement

matrix micro-cracks by the polymer [1,2,10] In the case

of mortars modified by latex rubbers, Justnes and Øye[11]

showed that the latex goes into a continuous network

be-yond a P/C ratio of 10 wt.% This polymer network would

be the origin of the improvement of the tensile and/or

flex-ural strength of PMMs

In a previous paper[2], we reported about the

mechan-ical behaviour of styrene butadiene modified mortars of

various polymer contents These composites were made

with an identical W/C ratio Compression and three-point

bending tests were carried out Three-point bending test

re-sults showed an increase of the flexural strength for higher

polymer contents up to 7.5 wt.% P/C ratio, then a decrease

beyond 10 wt.% Below 7.5 wt.% P/C ratio, the PMMs

re-mained the relative brittle behaviour of the original mortar

Beyond 10 wt.% P/C ratio, the PMMs got more and more

ductile, exhibiting a strong non-linearity before peak load

on the load-strain curve The compressive strength of the

materials remained identical up to 10 wt.% P/C ratio, then

decreased for higher polymer contents For both

mechan-ical tests, the strain at peak load increased for higher P/C

ratio while the Young’s modulus decreased Nevertheless,

latex addition in cement based materials induces some

Table 2

Paste density, dynamic viscosity and slump test results on the fresh materials vs polymer to cement weight ratio (P/C)

paste density (g cm −3) 2.192± 0.012 2.091 ± 0.003 2.068 ± 0.003 2.052 ± 0.002 2.029 ± 0.005 2.010 ± 0.007

Air content (vol.%) 3.7 ± 0.5 5.7 ± 0.1 5.9 ± 0.1 5.7 ± 0.1 6.0 ± 0.2 6.0 ± 0.3

Dynamic viscosity (MPa s) 73 ± 11.3 70 ± 2.3 71 ± 1.4 65 ± 1.4 68 ± 2.8 67 ± 1.4

Entrapped air content is calculated from the ratio between the paste density and the “theoretical” density.

adverse effects like variations in porosity or cement hydra-tion As a consequence, straight comparisons between the mechanical properties of PMMs of various polymer contents may not only reveal the influence of the polymer phase, but also include the influence of these adverse effects In our previous paper[2], some questions remained according

to that

In the present study, we still made up mortar and PMM samples of various polymer contents, but special care was taken to these adverse effects, especially by trying to keep constant the porosity over the materials Results about their porosity and other material characteristics [12] are to be subjected to further publications

The mechanical properties of the materials are again esti-mated using three point bend and compression tests After a description of the sample preparation, a summarize of their characteristics in terms of porosity, cement hydration, and other “material properties” is given Mechanical tests are then described and illustrated A discussion based on com-parisons between the results got on the different materials lead us to the main conclusions of this study

2 Materials

Polymer content is given according to the dry polymer

to cement weight ratio (P/C) The considered polymer con-tents are P/C= 0, 5, 7.5, 10, 12.5, 15 wt.% The water to

cement weight ratio (W/C) is identical for all the materials

in order to ensure a similar hydration of the cement matrix

On top of cement hydration, our goal was also to manage the porosity of the materials to be similar whatever the poly-mer content To do so, we used silica fume to optimise the compactness of the granular mix: silica fume, cement, sand, assuming that a lower porosity would be easier to control

A superplasticizer had to be used to ensure a good worka-bility of the pastes Moreover, the air entrainment effect of the latex was offset by the use of an antifoamer Finally, all the samples were prepared with the following compo-nents: Portland cement 52.5 (C), fine sand of medium grain size d50 = 330 ␮m (S), silica fume (SF), water (W), latex

(P), i.e an aqueous emulsion of styrene butadiene copoly-mer particles containing 48 wt.% of dry matter, sulfonated polymelamine based synthetic superplasticizer (SP) and an antifoaming compound (AF) The mix proportions of these components are given inTable 1

The materials were prepared according to the European

standard NF EN 1323 Four batches a material were needed

to make all the samples Paste density and dynamic viscosity

measurements as well as slump tests were carried out on

each batch Results are given inTable 2 The materials have

been moulded and cured for 24 h at 90% RH Then, the

40 mm× 40 mm × 160 mm prisms have been removed from

mould and stored an another day at 90% RH Samples have

been finally stored in a relatively dry atmosphere (50% RH

and 23◦C) for all the time.

Entrapped air content was estimated via the density of

fresh pastes compared with the “theoretical” density

ex-pected for pastes without any entrapped air (Table 2) The

density was determined by weighting a given volume of

fresh paste The “theoretical” density was calculated from

the mix proportions of the paste constituents and their own

density Entrapped air content was nearly the same between

the PMM pastes (volume fraction around 6%) but higher

than that of the mortar (around 4%) This initial gap clearly

highlights our difficulties to offset the air entrainment

ef-fect of the latex Fortunately, it has been made up for the

sample drying out with time Actually, the porosity of the

samples was estimated when the materials were 490 days

old The porosity of the mortar samples was about 18% and

that of the PMM samples about 20%[12] If the initial gap

remained, it had become no more significant with time As

a consequence, we assume that this difference will not play

a major role in our comparisons between the mechanical

properties of the materials and, thus, can be neglected

Some other “material properties” were also investigated

Scanning electron microscopy (SEM) of the latex

distribu-tion within the microstructure of the PMMs showed that

the latex phase goes into a continuous network

through-out the sample for a P/C ratio higher or equal to 10 wt.%

[12] These investigations also revealed that silica fume does

not react with other mineral species.Fig 1presents a back

scattered electron SEM image taken on a 1 year old PMM

Fig 1 Back scattered electron SEM micrograph of a 1 year old PMM

with 12.5 wt.% polymer to cement ratio White arrows denote the more

visible spherical silica fume particles.

(P/C= 12.5 wt.%) In this figure, spherical silica fume

par-ticles marked with white arrows can be easily distinguished:

no signs of degradation due to a possible pozzolanic effect

of added silica fume is visible Moreover, all the investiga-tions achieved on the other materials, and especially on the mortar, never revealed any signs of chemical activation of the silica fume particles Therefore, silica fume can be seen only as a filler Setting time measurements showed that the latex delays the setting of the PMMs Electric conductiv-ity measurements highlighted that the main reason of this

is a cement hydration kinetic decrease[12] This could be due to latex particle absorption onto the cement grain sur-face which would reduce the sursur-face of cement in contact with water[1,7,10] Infra-red Fourier transform (IRFT) and X-ray diffraction (XRD) analysis on the hardened materials showed however that their are all made up with the same mineral species Thus the latex would not modify the nature

of the cement hydration products even if it slows down the kinetic of the reaction

3 Experimental tests

3.1 Three-point bending tests

A sketch of the experimental setup is given in Fig 1 The three-point bending tests were performed on the

4505 type machine equipped with a 5 kN load cell and a PID servo control A strain measure was achieved on the face under tensile stress using an INSTRON extensometer put in the middle of the sample (seeFig 2) Details on the way the extensometer is tied to the sample are reported in Fig 2 The extensometer principle is to measure the relative

Fig 2 Experimental set-up for three-point bending tests Three-point bending tests were performed under strain control (˙ε = 10−6s−1).

spread of two articulated arms during the test Divided by

the initial distance between the two arms (l0), the

exten-someter gives a mean value of the strain between its arms

This device is fully fitted into the INSTRON electronic

interface and systematically calibrated before any new tests

series The PID servo control enabled us to rule the tests

under strain control The train rate was set to˙ε = 10−6s−1.

We systematically proceeded to successive partial

unload-ing at different strain levels to monitor the loss of sample

stiffness Loading and unloading were achieved at the same

strain rate At least three samples were tested for each

polymer content (P/C = 0, 5, 7.5, 10, 12.5, 15 wt.%)

The tests were carried out when the materials were 21

months old

These tests avoided an unstable fracture of the samples,

and allowed to reach the “softening” part of the material

response.Fig 3 shows a typical load versus tensile strain

diagram The ascending part of the curve is quasi-linear

However, near the top of the curve, the slope exhibits a

slight decrease witch indicates the damage initiation At peak

load, the micro-cracks coalesce into an unstable crack The

decreasing part of the curve corresponds to the propagation

of this crack which induces a loss of sample stiffness

Various parameters are used for our comparisons (see

Fig 3): the maximum load (Fmax) applied to the sample over

the test, the strain at peak load (εpeak ) associated with Fmax,

the initial sample stiffness (k0), and the one recorded

dur-ing the ith unloaddur-ing–loaddur-ing cycle (k i) which allowed us

to characterize the damage evolution during the test Note

that we do not convert the applied load into stresses and the

sample stiffness into elastic moduli to avoid the

introduc-tion of elastic beam calculaintroduc-tion hypothesis which would be

false in our sample geometry

The non-linear evolution of the load-strain curve indicates

the damage initiation and its growth In order to characterize

the material strength to damage, we defined a three-point

bending damage threshold This threshold (εd) is a criterion

based on a measure (δ) of the gap between the load-strain

curve and its initial linear part of slope k0 IfΓ is the set of

Fig 3 Typical load-strain curve of three-point bending tests (Fmax:

max-imum load reached over the test,εpeak: strain at Fmax, k0: initial slope of

the curve, k i : slope of the ith reload).

Fig 4 Determination of the three-point bending damage threshold.δ is

the gap between the experimental load-strain curve and the tangent to its initial part The damage threshold (εd, Fd ) is reached forδ = 1%.

points of the load-strain curve, then:

∀(ε, F) ∈ Γ, (ε, F) = (εd, Fd) ⇔ δ(ε, F)

= ε − F/k0

In order to determine with a good accuracy the couple of values (εd , Fd), we used a power function H (δ) = (100.δ)3

which lowers the values ofδ before the threshold and

in-creases them beyond.Fig 4gives an example of the (εd , Fd) determination

3.2 Uniaxial compression tests

Compression tests were carried out on 40 mm×40 mm ×

80 mm samples cut from the original 160 mm prisms (two pieces for each bending sample) The loaded faces of com-pression samples were ground to be parallel and flat The experimental assembly is described in Fig 5 These tests were achieved on a MTS hydraulic/500-kN press During these experiments, we measured and recorded the applied load, the longitudinal strain (εL) using an MTS extensome-ter similar to the one used for three-point bending tests, and the transversal strain (εT) using a set-up especially designed for these tests (seeFig 4) The transversal strain

determina-tion is based on the measure of the displacements (U1and

U2) of two opposite sample faces This measure is achieved

by two proximity sensors (sensors 1 and 2) running on eddy currents Two aluminum targets are then needed to make the

sensors working The sum of U1and U2divided by the

ini-tial width of the sample (a0) give a measure of the mean value ofεT The sensors are mounted to a support smoothly fixed on the sample Compression tests were carried out under transversal strain control The strain rate was set to:

εT= 2 × 10−6s−1 Two series of tests were achieved when

the materials were 13 and 24 months old

A typical stress–strain curve for compression test results is shown inFig 6 The stress (σ) is calculated from the applied

Fig 5 Experimental set-up for compression tests The longitudinal strain

(εL) is measured by an extensometer, the transversal strain (εT) is

mea-sured via the relative displacements (U1, U2) of two opposite faces of the

sample divided by its initial width (a0) U1 and U2 are given by

sen-sors 1 and 2 The applied load is also recorded Compression tests are

performed under transversal strain control (εT = 2 × 10 −6s−1).

load (F) divided by the sample initial section (S0) (σ =

F/S0) An automatic routine of data capture and processing

allows to determine the initial Young’s modulus (E0), its

value at the ith unloading (E i), the maximum stress reached

during the test (σmax), the longitudinal and transversal strains

(εpeak

L , εpeak

T ) associated withσmax

As for the three-point bending tests, a compression

dam-age threshold is defined in the following way:

∀(εL, εT) ∈ Γc, (εL, εT) = (εd

L, εd

T) ⇔ δ(εL, εT)

= εT − υ i (εL − ε i,irL )

where Γc is the set of points of the strain–strain curve,

(εd

L, εd

T) are the longitudinal and transversal strain at the

damage threshold, υ i is the Poisson’s ratio at the current

loading,ε i,ir is the irreversible strain at the current

load-ing As for the three point bending damage threshold, we

use the power function H(δ) = (100δ)3to determine the

Fig 6 Typical stress–strain diagram of compression tests (σmax : maximum

stress recorded over the test, εpeak

L , εpeak

T : longitudinal and transversal strains atσmax, E0 , T0: initial Young’s modulus and transversal stiffness,

E i , T i : Young’s modulus and transversal stiffness at the ith reload).

Fig 7 Determination of the compression damage threshold in the transversal–longitudinal strains diagram. νapp is the effective Poisson’s ratio for the reload where a non-linearity occurs in this diagram.δ is the

gap between the experimental curve and its initial tangent at this reload The damage threshold (εd

L, εd

T ) is reached forδ = 1%.

couple(εd

L, εd

T) An illustration of the (εd

L, εd

T) determination

is given inFig 7

4 Discussion

Result analysis is based on comparisons between the me-chanical properties of the materials We successively analyse the variation in Young’s modulus, damage threshold, maxi-mal stresses reached during the tests, their associated strains and, finally, the shape of the stress–strain curves according

to the polymer content

4.1 Variation in Young’s modulus and damage threshold

Fig 8 presents the variation in Young’s modulus de-termined by compression tests according to the polymer content of the materials It is obvious that the Young’s modulus decreases almost linearly with the P/C ratio In

Fig 8 Initial Young’s modulus vs polymer to cement weight ratio (P/C) for compression and three-point bending tests.

addition, we note that there is no significant difference

between materials about 1 year old and those 2 years old

This indicates that the materials were in quite a stable

state with regard to cement hydration or samples drying

out

In three-point bending, the Young’s modulus is

identi-fied by finite element computations from the sample initial

stiffness k0 However, the value of this module is not exact

since it does not take into account the possible difference of

stiffness between the part of the beam sample under tensile

stress and that under compressive stress in a cross-section

The three-point-bending Young’s moduli thus identified

are generally lower than those obtained by uniaxial

com-pression tests This difference points out that micro-cracks

pre-exist in these materials before any loading However, the

difference between the Young’s moduli determined by

com-pression and three-point bending tests, and thus the density

of “pre-existing” micro-cracks, decreases for higher P/C

ra-tio and cancels for P/C= 12.5 wt.%.

Fig 9 presents the variation in damage threshold

ac-cording to the polymer content For both mechanical

tests, the damage threshold increases with the polymer

content The compression-damage threshold measured

when the materials were 2 years old is systematically

lower than that measured 1 year before This result

prob-ably points out the ageing of the materials during this

period

Moreover, the bending damage threshold is lower than

that measured in compression It would be possible to do

straight comparisons between these two damage thresholds

by using a criterion on the strain tensor, but it is not obvious

whether this comparison has any sense Actually, the

com-pression damage threshold is based on a global measure of

the transversal strain over the full width of the sample On

the contrary, the tensile strain measure accounting for the

three-point bending tests is rather local The sensibility of

Fig 9 Transversal trains at damage threshold (εd

T ) and at peak load (εpeak

T ) for compression tests, and strains at damage threshold (εd) and at peak

load (εpeak) for three-point bending tests, vs polymer to cement weight

ratio (P/C).

Fig 10 Mechanical strength (Fmax) and load at damage threshold (Fd)

vs polymer to cement weight ratio (P/C) for three-point bending tests.

the strain measurements to damage is then certainly more important in the latter case than in the former Therefore, these data have to be handled with care if ones aims at es-tablishing a damage-threshold criterion

4.2 Load and maximal stress

Peak load and maximal stress reached during three-point bending and compression tests are plotted versus the P/C ratio in Figs 10 and 11 Load and stress at the damage threshold are also mentioned in these figures

In three-point-bending tests, the maximum load remains relatively constant for all the composites with a P/C ratio lower or equal to 7.5 wt.% For higher polymer contents, i.e P/C≥ 10 wt.%, the flexural strength increases with the

poly-mer content However this increase of the flexural strength slows down for the highest polymer content For PMMs with

a P/C ratio higher than 15 wt.%, a decrease of the mechan-ical strength can be expected Finally, we notice that the overall damage threshold tends to increase for higher poly-mer contents

Fig 11 Mechanical strength (σmax )and stress at damage threshold (σd )

vs polymer to cement weight ratio (P/C) for compression tests.

The reasons accounting for the evolution of the bending

strength are not obvious First of all, the porosity cannot play

a major role, since the volume fraction of pores is similar

throughout the materials, and especially in the PMMs The

influence of a possible difference in cement hydration is not

easy to handle However, it is clear that the cement hydration

and the sample drying out hardly evolve between the age of

1 and 2 since the compressive strength of the materials is

the same between these two dates (seeFig 11) So, we can

consider the materials in a stable state when the three-point

bending tests are performed Moreover, we know that cement

hydration strengthens the cement-based materials by filling

the capillary pores with hydrates, especially during the first

month after their making[13] If some differences between

the materials in cement hydration were responsible for the

bending strength evolution, this would also have a strong

influence on the evolution of the compressive strength But

this is not the case.Fig 11presents the compressive strength

of the materials versus the P/C ratio: it decreases almost

linearly with P/C, latex addition acting like an increase of

porosity due to its low rigidity Therefore, we assume that

cement hydration cannot account for the bending strength

evolution presented inFig 10

Finally, the last reason we consider is the percolation of

the polymer phase over the sample This phenomenon has

already been evoked to explain the tensile-strength increase

of polymer-modified cement-based materials[7,10] In our

case, 2x10x20 mm3 pieces of material were put into

hy-drochloric acid in order to dissolve the mineral phase For 5

and 7.5 wt.% P/C ratio, the pieces were dissolved in a

pow-der probably mainly composed of the fine sand Sometimes,

the dissolution of 7.5 wt.% P/C ratio pieces was not

com-plete and sheets of residual material of a few millimetre

re-mained, showing that the polymer starts to get organized in

a continuous network at this ratio For a P/C ratio higher or

equal to 10 wt.%, the pieces kept the same geometry, only

their colour changed from grey to white[12] Some of these

pieces were broken and put into a scanning electron

mi-croscope for energy-dispersive analysis of X-rays No more

calcium was found on the fracture surfaces, hydrates were

clearly dissolved [12] In conclusion, it seems clear now

that the main reason for an increase of the PMMs flexural

strength is polymer continuous network formation

4.3 Envelope of the stress–strain curves

We cleared the stress–strain curves from their loading–

unloading parts (seeFigs 10 and 11) to get their envelope

curve These curves sum up rather well the whole variation

in the mechanical behaviour of the materials according to

the polymer content For compression tests,Fig 12clearly

shows the decrease in the initial Young’s modulus as in the

mechanical strength, and the increase in the longitudinal

strain at peak load for higher polymer contents

For three-point bending tests, we notice in Fig 13

the early initiation of damage in the mortar sample (the

Fig 12 Envelope of stress–strain curves in compression for four polymer

to cement weight ratios (P/C).

ascending part of the curve breaks approximately at 1300 N) and, on the contrary, the linear shape of the ascending part

of the other curves almost to the peak: the polymer seems to delay the damage initiation into the PMMs This difference between the mortar and the PMMs could be due to the ex-istence of a film of latex on the PMM sample surfaces This film was revealed by optical microscopy analysis achieved

on the PMM sample faces Our hypothesis is that this film limits the PMM sample from drying out and, therefore, from surface micro-cracking On the contrary, micro-cracks

on the skin of the mortar sample, due to sample drying out, would be responsible for the early damage initiation This explanation should not be mistaken with the one accounting for the increase of the PMM flexural strength; a thin film

of latex on the sample faces cannot improve the overall mechanical strength of the beam It holds for the extremely non-linear shape of the load-strain curve of the mortar which reveals that micro-cracking happens early upon loading Pre-existing skin micro-cracks due to sample drying out may be the cause of such a behaviour On the opposite, the absence of skin micro-cracks on the PMM surfaces would enable their samples to exhibit a very linear behaviour

Fig 13 Envelope of load-strain curves in three-point bending for four polymer to cement weight ratios (P/C).

After peak load, the curve of the 5 wt.% P/C ratio sample

falls down to the one of the mortar and exhibits the same

shape for higher strain levels The curves of the 7.5 wt.%

P/C ratio samples, which are not given inFig 13, showed

the same behaviour after peak load The shape of the curve

of the 10 wt.% P/C ratio is quite different after peak load,

and the one of the 15 wt.% P/C ratio is definitely different

Therefore, as for the variation in the flexural strength (see

Fig 10), the post-peak behaviour of the materials exhibits

a strong change between 7.5 and 10 wt.% P/C ratio This

result has to be put in connection with the percolation of

latex into a continuous network for P/C ratio higher or equal

to 10 wt%

5 Conclusions

The mechanical behaviour of a mortar and five PMMs

have been investigated using compression and three-point

bending tests In compression, the results are the ones we can

expect from the addition of a phase with a low rigidity, the

latex, to a mortar: the Young’s modulus and the compressive

strength decrease for higher polymer contents, while the

strain at the maximum stress increases, which can be seen

as a gain of ductility

Three-point bending test analysis gives less obvious

re-sults The flexural strength increases for P/C ratios higher

or equal to 10 wt.%, and reaches a maximum for a P/C

ratio equal to 15 wt.% The increase on the PMM

flexu-ral strength and the latex continuous network formation

match Since a difference in cement hydration or in sample

porosity cannot explain this evolution, it is suggested that

the percolation of the polymer phase into a continuous

net-work sample increases the flexural strength of the polymer

modified mortars

The analysis of the shape of the load-strain curves showed

the early damage initiation in the mortar, unlike the PMMs

whose curves exhibit a very linear ascending part We

ex-press an hypothesis to explain this change in mechanical

be-haviour: the presence of a film of latex on the PMM sample

faces may prevent these materials from skin micro-cracking which could be the cause of the early damage initiation into the mortar

Comparisons between the Young’s modulus obtained from compression and three-point bending tests revealed the existence of micro-cracks in the materials before any loading However, this difference of rigidity decreases for higher polymer content and vanishes for P/C ratios higher than 12.5 wt.% Latex addition seems then to reduce this

“intrinsic” damage

Acknowledgements

This research was carried out in collaboration with the Centre de Recherche d’Aubervilliers de RHODIA The au-thors are grateful to G Orange and M Sari

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